A copper template was temporarily clamped down on a flat steel plate, and a rim was welded around the template. After removal ... with CNC or EDM equipment.
Dissertation
Shape Deposition Manufacturing ausgeführt zum Zwecke der Erlangung des akademischen Grades eines Doktors der technischen Wissenschaften eingereicht an der Technischen Universität Wien Fakultät für Elektrotechnik von
Dipl.-Ing. Robert Merz Sparkassenstr. 7 A-5020 Salzburg Matr.Nr. 8426325 geboren am 15.03.1966 in Salzburg
Begutachter: O. Prof. Dipl.-Ing. Dr. techn. DDr. hc. Fritz Paschke O. Prof. Dipl.-Ing. DDr. Helmut Detter
Wien , am 16. Mai 1994
Kurzfassung Rapid Prototyping Verfahren stellen dreidimensionale Strukturen mit beliebig komplexer Geometrie durch inkrementale Materialabscheidung in Form von 21/2 dimensionalen Schichten her. Automatische Planung und Exekution der Fabrikation wird dadurch ermöglicht, und die Verwendung von werkstückspezifischen Werkzeugen und Haltevorrichtungen entfällt. Allerdings sind Einschränkungen in der erzielten Oberflächengüte, die durch die 21/2 dimensionale Natur des Fertigungsprozesses hervorgerufen werden, in Kauf zu nehmen, und die bisher entwickelten Verfahren verfügen nicht über die Möglichkeit der direkten Produktion von völlig dichten, gut zusammenhaltenden Strukturen. Globale Märkte und zunehmender Konkurrenzdruck erfordern aber die rasche und kostengünstige Entwicklung von hochwertigen Produkten und rasche Anpassung an Änderungen in Konstruktion und Funktion. Dies wiederum erfordert die rasche Herstellung von funktionsfähigen Teilen, wie Prototypen und Werkzeuge für Massenproduktion (z.B. Spritzgußformen). Diese Dissertation befaßt sich mit der raschen und automatischen Herstellung von funktionsfähigen Metallteilen direkt von CAD Modellen. Ein neu entwickeltes Verfahren, “Shape Deposition Manufacturing” (SDM), wird beschrieben. Dieses Verfahren basiert auf dem Konzept der Herstellung von Teilen in Schichten, das von den Rapid Prototyping Verfahren bekannt ist, verwendet aber getrennte Prozeßschritte zur Ablagerung des Materials und zur Formgebung der Schicht. Dreidimensional geformte Lagen werden durch Schneiden mit fünfachsigen CNC Maschinen hergestellt, um die für funktionsfähige Teile notwendige Genauigkeit zu erreichen. Thermische Auftragsverfahren (thermisches Spritzen, Schweißen) werden zur Erlangung der notwendigen Materialeigenschaften verwendet. Ein neuartiges Auftragsfahren, Mikrogießen “Microcasting”, das auf der Ablagerung von flüssigen Tropfen basiert, wurde entwickelt, um gut haftende Schichten mit hoher Materialstärke, bei gleichzeitiger Minimierung der in die darunterliegenden Schichten geleiteten Wärme, herzustellen. Zur Herstellung von voll dreidimensionalen Schichten sind detailiertere Strategien notwendig, als die, die von konventionellen Rapid Prototyping Prozessen verwendet werden. Ein auf CAD basierendes Planungssystem addressiert dieses Problem durch das Zerteilen eines CAD Modells in Schichten und herstellbare Segmente mit voller dreidimensionaler Gemometrie. Eine automatische Teststation, die im Shape Deposition Labor der Carnegie Mellon Universität eingerichtet wurde, ist beschrieben, und zeigt die Möglichkeit der Automatisierung des Verfahrens. Der Microgußprozeß und seine Leistungsfähigkeit in der SDM Umgebung werden detailiert erklärt. Verschiedene Strategien und Materialkombinationen zur Herstellung der Supportstruktur wurden entwickelt, und werden mit detaillierten Strategien zur Herstellung von Teilen mit unterschiedlicher geometrischer Komplexität und unterschiedlichen Anforderungen an die Materialqualität beschrieben. Die Materialeigenschaften von mit dem SDM Verfahren hergestellten Strukturen werden ermittelt. Probleme, die die Genauigkeit und die Materialintegrität beeinflussen, und die hauptsächlich auf thermisch aufgebaute Spannungen zurückzuführen sind, werden für spätere Arbeiten aufgezeigt. Am Ende werden verschiedene Teile mit unterschiedlicher Komplexität präsentiert, die mit dem SDM Verfahren hergestellt wurden. Die Herstellungsdauer und der notwendige Materialaufwand sind aufgezeigt, und werden mit herkömmlichen Rapid Prototyping Verfahren verglichen. Neben der Fabrikation von voll funktionstüchtigen Teilen, ist es mit dem SDM Verfahren auch möglich, Strukturen aus mehreren Materialien, funktionstüchtige, bereits montierte Strukturen mit mehreren Teilen und konforme, eingebettete elektro-mechanische Systeme herzustellen.
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Abstract Solid Freeform Fabrication (SFF) processes “rapidly” create three dimensional shapes of arbitrary complex geometries by incremental material deposition of 21/2 dimensional layers. Automated planning and execution of the fabrication process is possible, and the need for part-specific tooling or fixturing is eliminated. However, there are inherent limitations in the resulting surface quality, due to the 21/2D nature of the building process, and the current approaches are not capable of directly producing fully dense, well-bonded metal structures. Global economies and increasing competition require the fast and cost effective development of high quality products and rapid changes in design and functionality, demanding the rapid creation of functional parts, such as prototypes and tools for mass productions (e.g. injection molds). This thesis addresses the issue of rapidly and automatically fabricating functional metal parts directly from CAD models. A newly developed process called Shape Deposition Manufacturing (SDM) is introduced. The process is based on the concept of layered manufacturing in SFF, but uses separate deposition and shaping steps to create a layer. Three dimensionally shaped layers are created using 5-axis CNC machining, to achieve the required geometric accuracy for fully functional shapes. Thermal deposition technologies (thermal spraying, welding) are used to achieve the required material properties. A novel, droplet based deposition process, microcasting, has been developed, to create well-bonded, high-strength material, while minimizing the heat input into previously shaped layers. To create layers with a true three dimensional geometry, more detailed building strategies, than used by conventional SFF processes, are required by the SDM process. A CAD based planning system, which addresses these issues by decomposing a solid model of a part into layers and manufacturable, fully three dimensional segments is described. An automated testbed facility installed at Carnegie Mellon’s Shape Deposition Laboratory is discussed, and shows the feasibility of automating the process. The microcasting process is explained in detail and it’s performance in the SDM environment is evaluated. Different strategies and material combinations for the support structure have been developed and are presented with detailed descriptions of several building strategies for parts with various complexity and material quality. Material properties of structures created by the SDM process are evaluated. Problems affecting the accuracy and material integrity of SDM created structures, which mainly involve the buildup of thermal stresses during material deposition, are identified and opened for future research. Finally, various parts, with different complexity, have been built with the SDM process, to show the feasibility and performance of the process. Building time and material usage are evaluated and compared to conventional SFF processes. Next to the creation of fully functional shapes, the SDM process has also shown the capability to manufacture multi-material structures, functional assemblies and conformable, embedded electromechanical systems.
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Acknowledgments First of all, I wish to thank my advisors at the Technical University of Vienna, Prof. Dr. Fritz Paschke and Prof. Dr. Helmut Detter for their support, help and inspiration over the past three years. Prof. Dr. Fritz Prinz and Dr. Lee Weiss deserve credit for their guidance at Carnegie Mellon University. Frequent discussions with them, and many of their suggestions were very helpful for the successful completion of this work. Discussions with Prof. Dr. Helmuth Kirchner have also been very helpful and inspiring. Larry Schultz has contributed significantly to the project. His exceptional knowledge, enthusiasm and open mind were a driving force in the laboratory and provided solutions for many problems. This work could not have been completed without him. Many other former or current members of the Engineering Design Research Center at Carnegie Mellon and the Shape Deposition Laboratory deserve credit for their contributions: Dr. Paul Fussell, Dr. Levent Gürsöz, Kevin Hartmann, Dr. Jim Hemmerle, Dr. Gennady Neplotnik, Krishnan Ramaswami, Ursula Sadiq, Kevin Schmaltz, Dr. Mike Terk, Dave Thuel, and John Zinn. On the industrial side, Andy Spizak and Robert Stosky of Packard Electric, Mark Novotarsky of Praxair, Inc., and Dan Gibson of Babcock and Wilcox deserve credit for providing equipment and their time to conduct various tests. Members of the staff of the Engineering Design Research Center, Tim Sager, Tara Taylor, and Silvia Walters, and Theresia Bruckner from the “Institut für Allgemeine Elektrotechnik und Elektronik”, have been very supportive with administrative details. Clauss Strauch provided excellent support for computer resources. I also with to thank the various sponsors for providing financial support for the Shape Deposition Laboratory: The Advanced Research Project Agency (ARPA), the National Science Foundation (NSF) and a series of industrial sponsors: Alcoa, Babcock and Wilcox, Ford Motor Co., Packard Electric and Delco Chassis (both divisions of General Motors) and United Technologies. Special credit has to be given to Prof. Dr. Fritz Paschke for initiating my stay at Carnegie Mellon University and the Austrian “Bundesministerium für Wissenschaft und Forschung” for providing the initial support through a “Kurt Gödel Stipendium”. Finally, and most of all I wish to thank my parents, for their support during my studies and their understanding, encouragement and care throughout the years.
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Table of Content 1. Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 1.1. Motivation
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1.2. Background . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 1.2.1. Solid Freeform Fabrication
. . . . . . . . . . . . . . . . . 2
1.2.2. A Survey of Current SFF-Technologies . . . . . . . . . . . . 3 1.2.3. Spray Based Approaches Towards Functional Parts . . . . . . 5 1.2.4. A Survey of Weld Based Deposition Processes . . . . . . . . 6 1.3. The Concept of Shape Deposition Manufacturing . . . . . . . . . . 7 1.4. Thesis Outline . . . . . . . . . . . . . . . . . . . . . . . . . . 8 2. Planning Shape Deposition . . . . . . . . . . . . . . . . . . . . . . 10 2.1. The Concept of SFF Processes . . . . . . . . . . . . . . . . . . 10 2.1.1. The Geometric Principle of SFF . . . . . . . . . . . . . . . 10 2.1.2. Theoretical Accuracy of SFF . . . . . . . . . . . . . . . . . 11 2.2. The Concept of SDM . . . . . . . . . . . . . . . . . . . . . . . 12 2.2.1. The Geometric Principle of SDM . . . . . . . . . . . . . . . 12 2.2.2. Surface Classification . . . . . . . . . . . . . . . . . . . . 13 2.3. Basic SDM Manufacturing Strategies . . . . . . . . . . . . . . . 15 2.3.1. Manufacture of Non-Undercut Features . . . . . . . . . . . . 16 2.3.2. Manufacture of Undercut Features . . . . . . . . . . . . . . 17 2.3.3. Manufacture of Arbitrary Layers . . . . . . . . . . . . . . . 17 2.4. The SDM Planning System . . . . . . . . . . . . . . . . . . . . 18 2.4.1. The CAD Model . . . . . . . . . . . . . . . . . . . . . . 21 2.4.2. The Surface Classification Module . . . . . . . . . . . . . . 22 2.4.3. The Adaptive Slicing Module . . . . . . . . . . . . . . . . 22 2.4.4. The Precedence Ordering Module . . . . . . . . . . . . . . 25 2.4.5. The Compact Splitting Module . . . . . . . . . . . . . . . . 26 2.4.6. Geometric Decomposition for General (Nonlinear) CAD Models . . . . . . . . . . . . . . . . . . . . . . . . 27 2.4.7. The Process Definition and Scheduling Module . . . . . . . . 28 2.4.8. The Path Planning Module
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. . . . . . . . . . . . . . . . . 29
2.4.9. Path Generation for Cutting Processes . . . . 2.4.9.1. 2D Contouring . . . . . . . . . . . . 2.4.9.2. 3D Shaping . . . . . . . . . . . . . . 2.4.9.3. Geometric Accuracy of Cutting Processes
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. 30 . 30 . 31 . 34
2.4.10. Path Generation for Deposition Processes . . . . . . . . . . 35 2.4.11. Machine Code Generation Module
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2.4.12. Conclusion and Future Developments . . . . . . . . . . . . 37 3. Automated SDM Testbed Facility . . . . . . . . . . . . . . . . . . . 39 3.1. Configuration of the Experimental SDM Cell . . . . . . . . . . . . 40 3.1.1. Robotic Transfer and Palletizing System
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3.1.2. Deposition Chamber . . . . . . . . . . . . . . . . . . . . 41 3.1.3. Shaping Station . . . . . . . . . . . . . . . . . . . . . . . 42 3.1.4. Grit-Blasting and Shot-Peening Stations
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3.2. SDM Process Control System . . . . . . . . . . . . . . . . . . . 43 4. The Microcasting Process . . . . . . . . . . . . . . . . . . . . . . 46 4.1. Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . 46 4.1.1. Thermal Spraying
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4.1.2. Conventional Welding
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4.2. The Principle of Microcasting . . . . . . . . . . . . . . . . . . . 49 4.2.1. Comparison of Substrate Heat Transfer . . . . . . . . . . . . 49 4.3. The Microcasting Apparatus
. . . . . . . . . . . . . . . . . . . 51
4.3.1. Early Versions of the Microcasting Device . . . . . . . . . . 51 4.3.2. The Plasma-Arc Microcaster . . . . . . . . . 4.3.2.1. Configuration . . . . . . . . . . . . . 4.3.2.2. Typical Microcaster Adjustments . . . . 4.3.2.3. Droplet Formation . . . . . . . . . . . 4.3.2.4. Droplet Trajectories . . . . . . . . . . 4.3.2.5. Deposition Voids . . . . . . . . . . . 4.3.2.6. Shrouding . . . . . . . . . . . . . . . 4.3.2.7. Typical Microcasting Operation Sequence
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. 52 . 52 . 54 . 55 . 56 . 58 . 60 . 61
4.4. Microcasting Parameters . . . . . . . . . . . . . . . . . . . . . 61 4.4.1. ER70S-6 Mild Steel . . . . . . . . . . . . . . . . . . . . . 62 4.4.2. 308 Stainless Steel . . . . . . . . . . . . . . . . . . . . . 62
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4.4.3. Deoxidized Copper . . . . . . . . . . . . . . . . . . . . . 64 4.5. Droplet Temperatures . . . . . . . . . . . . . . . . . . . . . . 65 4.5.1. Calorimetric Evaluation of Droplet Temperatures 4.5.1.1. ER70S-6 Mild Steel . . . . . . . . . . . 4.5.1.2. 308 Stainless Steel . . . . . . . . . . . . 4.5.1.3. Deoxidized Copper . . . . . . . . . . .
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4.5.2. Thermocouple Measurements of Droplet Temperatures 4.5.2.1. ER70S-6 Mild Steel . . . . . . . . . . . . . 4.5.2.2. 308 Stainless Steel . . . . . . . . . . . . . . 4.5.2.3. Deoxidized Copper . . . . . . . . . . . . .
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. 66 . 67 . 70 . 71
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. 73 . 76 . 77 . 78
4.5.3. Conclusions . . . . . . . . . . . . . . . . . . . . . . . . 80 4.6. Microstructure of Microcast Droplets
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4.6.1. Single Microcast Droplet . . . . . . . . . . . . . . . . . . . 81 4.6.2. Multiple Microcast Droplet
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4.7. Deposition of Dissimilar Materials . . . . . . . . . . . . . . . . . 85 4.8. Conclusions and Future Improvements . . . . . . . . . . . . . . . 86 4.8.1. Laser Microcaster
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4.8.2. Closed Loop Microcasting Control . . . . . . . . . . . . . . 88 5. Support Structures
. . . . . . . . . . . . . . . . . . . . . . . . . 89
5.1. Support Structure Geometry
. . . . . . . . . . . . . . . . . . . 89
5.2. Material Requirements . . . . . . . . . . . . . . . . . . . . . . 90 5.2.1. Support for Arbitrary Geometries . . . . . . . . . . . . . . . 90 5.2.2. Support for Parts without Undercuts . . . . . . . . . . . . . 91 5.2.3. Removal of the Support Structure . . . . . . . . . . . . . . 91 5.3. Materials Penetration . . . . . . . . . . . . . . . . . . . . . . . 92 5.4. Support Materials for Different Applications . . . . . . . . . . . . 93 5.4.1. Support for Thermally Sprayed Structures . . . . . . . . . . . 93 5.4.2. Low Melting Support with Thermal Barrier for High Temperature Metals . . . . . . . . . . . . . . . 5.4.2.1. Interface Temperature . . . . . . . . . . . 5.4.2.2. Deposition of Dissimilar Materials . . . . . 5.4.2.3. Substrate Temperature during Deposition . . 5.4.2.4. Thermal Barrier Protection of Low Melting Support Structure . . . . . . . . . . . . .
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. 94 . 94 . 95 . 96
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5.4.3. Powder Support for Microcast Structures without Undercuts . . 99 5.4.4. Copper Support for Microcast Steel Structures . . . 5.4.4.1. Remelting Conditions for Dissimilar Materials 5.4.4.2. Estimated Temperatures for Stainless Steel and Copper . . . . . . . . . . . . . . . . 5.4.4.3. Support Material Removal . . . . . . . . .
. . . . . 101 . . . . . 101 . . . . . 104 . . . . . 107
5.4.5. Additional Support Material Strategies . . . . . . . . . . . 108 6. Manufacturing Strategies . . . . . . . . . . . . . . . . . . . . . . 110 6.1. Interlayer Penetration . . . . . . . . . . . . . . . . . . . . . . 110 6.1.1. High Interlayer Boundary . . . . . . . . . . . . . . . . . 110 6.1.2. Low Interlayer Boundary . . . . . . . . . . . . . . . . . 111 6.2. Compensation for Penetration and Surface Abrasion . . . . . . . . 112 6.2.1. Selective Protection . . . . . . . . . . . . . . . . . . . . 112 6.2.2. Stepped Compensation . . . . . . . . . . . . . . . . . . 114 6.3. Sequence of Shaping Operations
. . . . . . . . . . . . . . . . 116
6.4. Adaptive Compact Geometries . . . . . . . . . . . . . . . . . 117 6.5. Process Definitions
. . . . . . . . . . . . . . . . . . . . . . 118
6.5.1. Arbitrary 3D Structures . . . . . . . . . . . . . . . . . . 119 6.5.2. 21/2D One-Material Structures . . . . . . . . . . . . . . . 122 6.5.3. 3D Structures without Undercuts and Powder Support . . . . 123 6.5.4. Microcast, Non-Overhang Structures with Shrinkage Compensation . . . . . . . . . . . . . . . . . . . . . . 124 6.5.5. Simplified Strategy for 3D Microcast Structures with Solid Support . . . . . . . . . . . . . . . . . . . . . . . 125 7. Mechanical Properties of Shape Deposited Materials and Stress-Related Problems in Layered Forming . . . . . . . . . . . . 127 7.1. Material Properties of Sprayed Structures
. . . . . . . . . . . . 127
7.1.1. Properties of a Sprayed Zinc - Nickel/Aluminum Laminate . . 127 7.1.1.1. Tensile Testing . . . . . . . . . . . . . . . . . . . 128 7.1.1.2. Adhesion Testing . . . . . . . . . . . . . . . . . . 129 7.1.2. Properties of Microcast Steels . . . . . . . . . . . . . . . 129 7.1.2.1. Microcast ER70S-6 Mild Steel . . . . . . . . . . . . 130 7.1.2.2. Microcast 308 Stainless Steel . . . . . . . . . . . . . 132 7.2. Residual Stresses in Layered Deposits . . . . . . . . . . . . . . 132
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7.2.1. Temperature Gradients After Solidification . . . . . . . . . 132 7.2.2. Effects of Residual Stresses on the Artifact 7.2.2.1. Artifact Warping . . . . . . . . . 7.2.2.2. “Christmas Tree” Effect . . . . . . 7.2.2.3. Debonding . . . . . . . . . . . .
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133 133 134 136
7.2.3. Counteracting Internal Stress Buildup . . . . . . . . . . . . 136 8. Example SDM Parts . . . . . . . . . . . . . . . . . . . . . . . . 139 8.1. Sprayed Parts . . . . . . . . . . . . . . . . . . . . . . . . . 139 8.1.1. Zinc Laminate Test Parts . . . . . . . . . . . . . . . . . 140 8.1.1.1. Thin, L-Shaped Wall . . . . . . . . . . . . . . . . . 140 8.1.1.2. Interlocking Frames with 3D-Shaped Surfaces . . . . . 141 8.1.2. IMS-T1 Test Part . . . . . . . . . . . . . . . . . . . . . 142 8.2. Microcast Parts
. . . . . . . . . . . . . . . . . . . . . . . . 145
8.2.1. Steel Parts with Sprayed Support . . . . . . . . . . . . . . 145 8.2.1.1. Injection Molding Die with Internal, Serpentine Cooling Channel . . . . . . . . . . . . . . . . . . . 145 8.2.1.2. Steel Cube with Internal Cavity and Embedded Sphere . 146 8.2.2. Copper-Steel Multimaterial Parts . . . . . . . . . . . . . . 147 8.2.2.1. Laminated Tube with Copper and Steel Layers . . . . . 147 8.2.2.2. Copper Wheel on a Steel Axle . . . . . . . . . . . . 147 8.2.3. Steel Mold Half . . . . . . . . . . . . . . . . . . . . . . 149 8.2.4. IMS-T2 . . . . . . . . . . . . . . . . . . . . . . . . . 150 8.2.4.1. Materials Usage for IMS-T2 . . . . . . . . . . . . . 156 8.3. Building Times . . . . . . . . . . . . . . . . . . . . . . . . 158 8.3.1. IMS-T1 . . . . . . . . . . . . . . . . . . . . . . . . . 159 8.3.2. IMS-T2 . . . . . . . . . . . . . . . . . . . . . . . . . 161 8.4. Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . 163 8.4.1. Summary of Current Work
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8.4.2. Future Research and Goals
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References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 165 Appendix A: Pictures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A-1
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1. Introduction 1.1. Motivation Solid Freeform Fabrication (SFF) processes “rapidly” create three dimensional shapes of arbitrarily complex geometries by incremental material deposition of 21/2 dimensional, cross-sectional layers1 embedded in complementary shaped, sacrificial support material. This approach facilitates the automatic planning and execution of fabrication by eliminating the need for partspecific tooling or fixturing. However, there is an inherent limitation in the resulting surface quality due to the 21/2D nature of the building process. In addition, current SFF approaches are limited in their capability to directly deposit fully dense, well-bonded metal structures. Functional metal shapes, such as production-quality custom tooling (e.g. injection molds) typically require superior surface and material properties. Therefore they cannot be directly2 created with current SFF approaches. This thesis addresses the issue of rapidly and automatically fabricating functional metal parts directly from CAD models using a modified SFF process, called Shape Deposition Manufacturing (SDM). In this process layers are deposited in form of truly three dimensional slices. Material deposition is based on welding-type processes to achieve the material properties required for fully functional parts.
1.2. Background Today’s global economies and increasing competition require fast and cost effective development of high quality products and rapid changes in design and functionality to respond to market demands. A variety of measures can be taken to meet the challenges imposed by the desire for decreased time to market for mass products and decreased delivery times for low production “one-of-a-kind” parts. Having prototype models available for various kinds of testing and visualization can dramatically decrease the time needed to design a product. This need has been successfully addressed by several SFF processes. Manufacturing functional products, such as prototypes or tools for mass production (e.g. injection molds) is still a very time consuming and expensive process, and mostly done “manually” by skilled personnel. Typical lead times for manufacturing tooling and prototypes range from several month (e.g. typical injection molds for automotive parts) up to one and a half years for prototypes with complicated geometry and tight specifications on tolerances (e.g. ship propellers).
1. for detailed explanation see chapter 1.2.1. 2. indirect creation possible through post-processing or conversion; see Chapter 1.2.1.
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1.2.1. Solid Freeform Fabrication Over the past ten years a new manufacturing approach known as Solid Freeform Fabrication (SFF) [1], [2], [3] has been developed through the merging of several previously distinct technologies, as one method to help shorten the time between initial conception of a product and its final production. With the original idea of quickly making models of parts a variety of SFF processes have been developed with a common general concept. A three dimensional solid or surface model CAD representation of the part is sliced into thin, planar layers with a constant thickness of typical 0.125 to 0.25 mm. Cross-sectional data is then sent layer by layer to a SFF-machine. Each layer is created by incremental material buildup of 21/2-dimensional sections, i.e. the geometry of the layer consists of the two-dimensional cross-section built up to a uniform thickness. Depending on the thickness of the layers the surface of the completed part will show a more or less distinct stair-step texture. To accommodate geometries with overhanging features the parts are embedded in a sacrificial supporting material during buildup. The principal advantage of SFF processes is the ease and speed with which one can go from part design to part fabrication in a CAD/CAM environment. Conventional manufacturing through CNC machining involves extensive reasoning about three-dimensional geometries and the design or selection of part-specific fixturing and tooling. For complex geometries multiple re-fixturing might be necessary, resulting in time-consuming and expensive procedures for fixture preparation and indexing. Automatic planning of CNC trajectories remains a challenging task and is restricted to a limited set of geometric features. An object oriented, feature based design environment and feature refinement are necessary to automatically manufacture simple parts [4]. Laminated approaches have been taken for geometries without undercuts and limited aspect ratios, which shape a part in layers from the top down and result in limited surface accuracy and quality [5]. Computational geometric reasoning for path generation is not yet possible for all geometries without the intervention of an experienced human operator. Manual path generation for arbitrarily complex shapes, such as turbine blades or air foils, remains a time-consuming task, especially when re-fixturing is needed to avoid collisions for features which are obstructed by other parts of the geometry. Other features, such as integrated internal cooling lines or narrow, deep channels cannot be built directly. Parts requiring those features have to assembled from different, separately machined pieces. In contrast, the SFF processes operate on simple planar geometries that do not require part-specific fixturing or tooling information. The planning and execution effort for SFF is essentially independent of part complexity. Because of SFF being an additive material method, it is possible to create single-component structures that would be impractical or impossible to build with traditional approaches. Examples include monolithic parts with internal channels or parts contained inside of boxes. Operating SFF equipment also requires a minimum of human intervention. The part designer can even personally fabricate a prototype. The usefulness of SFF technologies is dramatically demonstrated by an example from Ford Motor Company’s Alpha Engineering Group (Dearborn, Michigan) [6]. A rocker arm for the 1992 5L engine was designed using a stereolithography model for visualization. In addition to a reduction of the design time from 90 to one day, the final design resulted in a higher quality part because of a problem which was only discovered after inspection of the prototype. Also, when models of the
2
part where sent to the suppliers in addition to the blueprints, the price was reduced from a previous quote, made from the blueprints alone, which resulted in an annual savings of $3 million1.
1.2.2. A Survey of Current SFF-Technologies One of the first known approaches to layered manufacturing was used by the Egyptians to build the pyramids. According to Edwards [7] an entire level was completed, including core, packing blocks and casing blocks, before another level was added. To move the massive stones to the growing pyramid a sacrificial supply ramp was used, in concept quite similar to the support structure of SFF processes. The first commercially available SFF process was stereolithography (SLA) [8] from 3D Systems, Inc. in 1988. It is based on liquid photopolymer resins that solidify when exposed to ultraviolet radiation. The part is supported by an elevator platform which is submerged in a vat of liquid resin. For each layer the top of the previous layer is kept just underneath the surface to allow resin to coat the top surface. An ultraviolet laser scans the surface of the liquid to partially cure the layer. The elevator is then lowered further into the vat to build the next layer. For structures like overhanging features a support structure is needed. The support structure consists of thin, vertical walls of cured resin. In most practical applications a somewhat skilled operator manually creates the geometry of the support structure. After the scanning process is finished, the part is removed from the elevator. The part is cleaned from liquid resin, and the support structure is removed. After postcuring the part under ultraviolet light, to solidify any uncured resin, the part is finished by sanding or glass-beading. A process similar to stereolithography is Cubital’s Solid Ground Curing Technology (Solider) [9]. It uses erasable masks to expose only those sections of the liquid resin that should be cured, to light from an ultraviolet lamp. Uncured resin is removed from the layer and the cavities are filled with a water-soluble wax. After the layer is milled flat, the process starts again by spreading a thin layer of photopolymer for the next layer. When the SFF part is completed, the wax support structure is removed by melting or dissolution. The major advantages of this process are the elimination of creating a support structure, no final curing, a small increase in accuracy because of the solid environment in which the part is built, and faster build time due to simultaneous curing of the whole layer. Selective Laser Sintering (SLS) [10] is a process being developed by The University of Texas at Austin and DTM Cooperation. It uses heat-fusable powders, that are spread over a surface, and a laser scans and sinters the cross-section of the layer. A new layer of powder is then spread and the process continues until the part is finished. During buildup the part stays embedded in the unsintered powder which serves as support material for subsequent layers. Currently used materials are investment casting wax, various plastics and ceramic or metal powders coated with a polymeric binder. Postcuring of the SLS-parts is required to remove the binder similar to the thermal degradation or solvent extraction techniques used for injection mold production.
1. Figures courtesy of Peter R. Sferro, Ford Motor Company.
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Another SFF process using a powder technology was developed by the Massachusetts Institute of Technology (MIT). Three-Dimensional Printing (3DP) [11] uses ink jets to spray a cross-sectional pattern of binder into a layer of powder. Another layer of powder is then applied onto of the preceding one and the process is repeated. The part remains supported in the unbound portion of the powder. Upon completion the binder is removed from the “green” parts and the part is being sintered. In best cases the porosity of the “green” and debindered parts lays between 20 and 30% [3]. Typical powders used are ceramic powders such as aluminum oxide, silica, zirconia and silicon carbide. Experiments with stainless steel resulted in a low density of the part (around 78%) and significant shrinkage during the sintering process [12]. Ceramic cores and shells generated by 3DP were successfully used for investment casting. Soligen, Inc. has commercialized the process for ceramic core and mold applications under the trade name Direct Shell Production Casting. Fused Deposition Modelling (FDM) [13] is a one step process, that uses continuous wires of thermoplastic polymers or wax. The wire is heated and extruded through a nozzle onto a stationary base. Computer controlled x-y motion of the nozzle traces the cross-section of each layer. Due to solidification within 0.1s horizontal overhangs can be built without any support structure. Laminated Object Manufacturing (LOM) [14] uses paper or ceramic, plastic or metal sheets coated with a thermally activated adhesive. A new layer is glued to the previous layer using a heated roller. The contours of the cross-section is cut into the sheet with a laser. Portions of the layer not belonging to the part are crosshatched, and can be broken away upon completion of the part. Postprocessing of the parts is not required and they have a woodlike appearance. Table 1.1 summarizes the most important and commercially available SFF processes. Even SFF-Process
Material
Support Structure
Application
Applications with Postprocessing
Stereolithography
photopolymer resin, postcuring required
thin photopolymer walls
visualization, form fitting
patterns for mold making
Solid Ground Curing
photopolymer resin
wax
visualization, form fitting
patterns for mold making
Selective Laser Sintering
investment casting wax, polycarbonate, PVC, ABS plastic, nylon, binder coated metals
unsintered powder
injection molded parts
mold patterns, investment casting patterns, metal parts through infiltration
3D Printing
ceramics
unbound powder
ceramic shells and cores
metal parts through casting or infiltration
Fused Deposition Modelling
investment casting wax, nylon
none
plastic models
investment casting and mold patterns
Laminated Object Manufacturing
paper or polyester sheets
scored laminate
modelling
Table 1.1: Most Important SFF Processes though the listed processes have produced a significant impact in design and manufacturing, none
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have the capability of directly producing fully functional parts, i.e. parts of high structural integrity (i.e. mainly metal parts) with excellent dimensional tolerances, that could be used in operational systems. Currently, to achieve parts with properties equivalent to parts made with conventional manufacturing technologies, SFF parts are used as a starting point for additional processing steps, the SFF parts are further processed or converted. Postprocessing methods modify the microstructure, properties or characteristics of the material. In Addition to grinding, they include conventional sintering and hot isostatic pressing, which are normally combined with significant shrinkage of the part. This must be taken into account in the original part design. For porous parts, such as the ones produced by SLS or 3DP infiltration can provide densification and mechanical integrity. Conversion methods try to convert the SFF-part into metal, ceramic or shapes from a different polymer without loss of form or dimension. Ceramic SFF-parts, for example, are used as investment casting molds for metal parts, wax SFF-parts can be used in the lost wax investment casting process, other SFF-parts can be used to shape the mold cavities by molding with RTV (room-temperature vulcanizing) silicone, epoxy molding or sand casting. Other methods for creating molds from SFF-parts include metal plating and thermal spraying of metals.
1.2.3. Spray Based Approaches Towards Functional Parts The inception of thermal spray technologies for metals goes back to M. U. Schoop, who invented the first spray gun in 1909 [15]. For decades sprayed metals have been used for corrosion, heat and wear resistant coatings and for parts restoration. About 30 years ago, the spraying process has been adapted to produce metal tooling. With the development of SFF technologies, this application became widely used. In the early stages spraying metal tooling has been limited to materials with low melting points such as zinc or zinc-alloys. The process steps for manufacturing a zinc injection molding tool is described in detail in [16]. Only a short outline should be given here to understand the basic principle. To create the first half of the injection mold, a plastic pattern shaped like the complement of the interior of the first mold half (i.e. the partial part shape embedded in a block resembling the parting plane and the runner system) is created using a SFF process (e.g. SLA). After spraying a release agent and mounting a metal frame, a metal shell is sprayed onto the pattern. The shell is then filled with a backing material (typical epoxy filled with metal shot) and the substrate pattern is removed from the mold half. To build the second half of the mold, a SFF model of the part being molded, including the runner system and gate is inserted in the first mold half and the process is repeated. Based upon this concept, the technique has been expanded to the manufacture of thick shelled, sprayed-steel tooling [17], [18] to increase wear resistance. The process for spraying zinc molds had to be modified to avoid problems with sprayed steel not adhering to smooth surfaces due to higher internal stresses and excessive heat input into the plastic pattern during the spray operation. First a tin-bismuth alloy was sprayed onto a SFF pattern of the desired shape of the mold (not the inverse as in the zinc process). After reinforcing with a backing material and releasing the plastic pattern, this “mold” could be used to create the steel mold half exactly like the sprayed zinc molds. To release the steel mold half, the tin-bismuth pattern is melted. While the sprayed tooling approach is attractive for several applications, such as prototype tooling, it has several limitations. Due to the nature of spray processes geometric shapes with high aspect ratios are difficult to manufacture. It is also difficult to maintain accuracy since several process steps are involved. The
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composite structure of the final tool (metal shell with typically non metallic backing) causes delamination of the metal shell from the backing due to thermal stresses and fatigue. Lifetime of typical zinc injection molds can be expected at around 500, typical steel injection molds at around 1500 shots. To address these problems thermal spray technologies were combined with the concept of layered manufacturing. This led to a new SFF process, which was developed at Carnegie Mellon in 1991 [19]. The MD* (for recursive mask and deposit) process creates parts by successively spraying cross-sectional layers. Each layer is shaped by a disposable mask, which has the shape of the cross-section of the current layer. Each mask is cut with a CO2 laser [20] and then placed upon the top layer of the growing part. A robotically manipulated thermal spray gun traverses the areas exposed by the mask. As an additional advantage over other SFF process, the successive use of multiple masks per layer enables the manufacture of multi-material structures. Typical masking materials are paper supplied on rolls, metal foil or thin metal sheet for materials with higher melting points. As a support material for parts made of materials with low melting points (e.g. zinc) parts of the paper mask can be left in place. For materials with higher melting temperature (e.g. steel) or for multi-material parts a sacrificial low melting support material can be sprayed. When the part is complete, it can be removed by melting. While a few artifacts have been successfully built with a semi automated version of the MD* process, it has several limitations. Full automation of the masking process is extremely difficult, especially for features with small dimensions, and can not be accomplished in a reliable fashion. Other limitations are also common to other SFF processes. Any layered approach, in which material deposition involves thermal bonding, residual stress builds up due to unavoidable temperature gradients. This will lead to warpage and delaminations. The process still exhibits the for SFF processes typical stair-step texture, and it is difficult to achieve the accuracy, precision and resolution which can be accomplished with conventional CNC machining.
1.2.4. A Survey of Weld Based Deposition Processes In order to achieve the strength required for structural components, deposition welding methods have been successfully used in the past to restore worn machine parts and to manufacture near net shapes. One very early invention in this field was a method for making molds for vulcanized rubber soles [21]. A copper template was temporarily clamped down on a flat steel plate, and a rim was welded around the template. After removal of the template and repeating the process to create more molds on one plate, the rims were ground to the desired height. Another application of deposition welding is used for the repair or surfacing of heavy rollers [22]. While the roller is slowly rotated, one or more layers of weld material are applied by a torch traversing along the axis of the roller. Finally, after enough material has been applied, the roller is turned to the desired diameter. More delicate parts, such as blades for turbine engines, are normally repaired or hardfaced using special welding processes such as plasma arc welding or microplasma powder welding [23]. Low heat input into the substrate is essential to keep the heat affected zone to a minimum and to prevent distortion, which would occur from traditional welding processes. A machining operation is performed after the application of the material and requires special fixturing and indexing. Other weld based processes have tried to incorporate deposition and shaping into one step without
6
using a preform. The Shape Melting technology [24] developed by the Babcok & Wilcox Company tries to build heavy, axis-symmetrical workpieces entirely from weld material. This method uses a rotating, water-cooled, reusable shoe, which is positioned underneath a torch head [25]. The shoe provides support for the weld puddle and cooling for the weld material as it is deposited. The shoe stays in constant friction contact with the workpiece and travels with the weld head, thus providing a fresh shoe surface for the weld puddle and therefore shaping the workpiece. A similar method, which has been claimed in [26], builds up axis-symmetrical workpieces vertically (in contrast to the horizontal building direction of Shape Melting) using two independently moveable shoes. While the workpiece is being rotated, a weld head mounted and travelling between the shoes deposits material. The molten material is confined by the two shoes, thus forming the walls of the workpiece. Since no direct surface machining is involved in both methods, they are ideal for economically building big, high quality parts, where surface tolerances are not extremely critical, such as in pressure vessels or similar applications. Rapid Prototyping research at the University of Nottingham (UK) was based on a robotic 3D welding system [27]. Initial work included the welding of simple, unsupported straight and sloping walls. Even though some more complicated structures were build later, a number of basic limitations, such as excessive heat input, distortion, penetration, the necessity of reorienting the workpiece for construction of certain geometries, and the lack of a smooth surface, prevent the manufacture of perfect parts. In another approach [28], similar to the LOM process, mild steel or stainless steel sheet metal laminates were cut by laser machining, stacked and welded together with a laser. However, tolerance issues (stair-step effect) with thicker sheet metal (1 or 2 mm), and problems with work holding, heat dissipation and distortion with thinner laminates make it unlikely for this approach to produce high quality parts.
1.3. The Concept of Shape Deposition Manufacturing The usefulness of a part for certain applications is mainly determined by the properties of two areas, the properties of the material and the geometric tolerances. High quality parts for use in fully functional machinery and structures mainly demand excellent values for both. Materials properties (strength, wear resistance) of e.g. metal parts and structures have to be close to the ones of cast or welded materials, the geometric tolerances have to be comparable to the ones achieved with CNC or EDM equipment. Current rapid prototyping technologies are not able to meet those demands. Parts produced by SFF processes do not have the required material properties and due to the stair-step surface texture and process limitations geometric accuracies are at least one order of magnitude smaller than the ones from conventional CNC equipment. Weld deposition processes can achieve the required material properties, but are far worse in terms of accuracy. CNC machining on the other hand, can meet material quality as well as accuracy constraints, but are limited in geometry and expensive (special fixturing, time-consuming path planning and programming) for prototype applications. Considering the benefits and deficiencies of the processes laid out above, it seems evident, that combining the accuracies achieved with CNC equipment, the material qualities of weld based processes and the layered principle of SFF technology can lead to a superior manufacturing pro-
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cess. Such a process, Shape Deposition Manufacturing (SDM), will be described in this thesis. SDM will use various methods of thermal deposition for creating a near-net shape for each layer, which will be shaped with a 5-axis CNC milling machine. The thermal deposition technologies include thermal spraying of metals, plastics and ceramics with wire arc and plasma spray torches, a newly developed microcasting process for deposition of molten metal droplets, casting of materials with low melting points, such as waxes, tin, or zinc alloys, and conventional welding techniques. In addition to fully functional parts, the SDM process will also allow the manufacture of multi-material parts and conformable embedded electro-mechanical structures. Several patents have been issued related to the Shape Deposition process ([42], [61] - [68]).
1.4. Thesis Outline In contrast to current SFF processes, which create 21/2 dimensional layers with constant crosssections, built up to a fixed thickness, SDM will slice the part geometry according to local curvatures with an adaptive layer thickness and also shape the “side-band” according to the real, three dimensional geometry of the part. Chapter 2 explains the general geometric principles used for manufacturing arbitrary geometries with three dimensionally shaped layers. The requirements for a CAD based planning tool are identified. The steps to decompose a model of a part with arbitrary geometry into manufacturable segments is shown. A rule based syntax is developed to define and store a set of different process schedules. Finally, a sequence of manufacturing steps is translated into actual path trajectories and instructions readable by the different stations in the manufacturing cell. In Chapter 3 an experimental, automated testbed facility is described, which has been used to develop the SDM process, and which has shown the feasibility of industrial automation of the process. A robot based system, incorporating individual processing stations has been built. Each of the stations houses one or more processes known from conventional manufacturing. The flexible, robotic setup allows reconfiguration of the testbed facility to adapt to changing process strategies and permits the incorporation of additional subprocesses if necessary, thus providing an ideal environment for the development of the SDM process. To overcome the materials and adhesion problems of spray technologies, the penetration problems and the problems created by excessive heat input into substrate materials of traditional deposit welding techniques, a microcasting process has been developed. This process is based on the deposition of liquid metal droplets, which are generated away from the substrate, thus reducing the thermal impact on the substrate to the effects caused by the impinging droplets. Droplets can be created from different materials with a wide range of temperatures. The microcasting process is described in Chapter 4. The layered creation of three-dimensional structures requires the use of support materials to protect underlying, previously shaped layers, and to provide support for overhanging features. Various support material strategies for different deposition processes and different geometric features are discussed in Chapter 5. Remelting conditions for a droplet based deposition of two different materials are discussed, and the required properties for a combination of part material - support
8
material are presented. In Chapter 6 several process strategies for the creation of parts are presented. Each strategy describes the sequence of manufacturing operations necessary to build a certain structure. The individual sequences are influenced by the geometry of the desired part, the material and the deposition process, and the support material strategy. Problems with materials penetration and a geometry based compensation are also shown. The mechanical properties of sprayed and microcast materials, which were used for the creation of test parts, are presentedin Chapter 7. The microstructure of the materials and the ability of the microcasting process to produce high strength, multi-material structures with clearly defined boundaries are shown. Further, problems related to the buildup of thermal stresses during the deposition process, affecting the shape and structure of the individual layers, are identified. Finally, in Chapter 8 a variety of test geometries and example parts, which have been built with the SDM process are shown. An analysis on material usage and building time is given, to compare the SDM process with conventional rapid prototyping systems. A summary of the accomplishment of the current work, and an outlook towards future research and the future direction of the SDM process conclude this chapter.
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2. Planning Shape Deposition 2.1. The Concept of SFF Processes Conventional CNC machining techniques use the principle of material removal to fabricate parts with certain geometries. PlanningCNC operations involves several complex problems including reasoning about three dimensional geometries. In contrast, SFF processes work on an additive basis and deal mainly with two dimensional geometric information. This simplifies the planning effort to the point, where it is essentially independent of part complexity.
2.1.1. The Geometric Principle of SFF
a) =
U
CAD Model
2D Cross-Sections
Set of Planes
b)
c)
SFF Part
Embedded 2-1/2D Layers
Figure 2.1: The Principle of SFF Processes
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For most SFF processes the geometric information of the part is contained in a surface or solid CAD model. A set of equidistant, parallel planes, perpendicular to the building direction is intersected with the model to obtain cross-sectional layers (Figure 2.1 a)). Each cross-section is then built up in form of a 21/2 dimensional disk. Figure 2.1 b) shows a cross-sectional view of the layers contained inside the support material. In Figure 2.1 c) the support material has been removed. The part shows the stair-step texture typical for SFF processes.
2.1.2. Theoretical Accuracy of SFF Under the assumption of perfect process parameters, the theoretically achievable accuracy of SFF can be estimated. Real world process parameters, such as finite laser beam thickness, nonuniform material deposition, or shrinkage and warpage, play an additional role in the accuracy of the part, but shall not be taken into account here. Two effects play an important role in the theoretical accuracy of the part; the finite and fixed layer thickness and the stair-step texture. First, the height of the part is not always a multiple of the layer thickness, and therefore the height of the SFF-part (hSFF) can deviate from the height of the model (hModel) by up to half a layer thickness (hLayer). h Layer h SFF – h Model ≤ ------------2
(2.1)
The second effect on the accuracy, caused by the stair-step texture, depends on the local curvature of the part surface. Since layer thicknesses are generally kept thin, it can be estimated by linearizing the surface. surface of model part deviation d from model surface
local surface normal n
h Layer --------------2
α
local tangential plane cross-section of model part direction of growth
previous layer
Figure 2.2: The Influence of the Stair-Step Effect on SFF Accuracy
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Figure 2.2 shows part of a cross-section taken parallel to the building direction through one layer. The tangential plane is drawn at the intersection of the cross-section, that is used to build the layer, and the surface of the part. For a layer thickness hLayer the approximated maximum deviation d of the surface of the created disk from the desired surface of the model is calculated in (2.2), where α is the angle between the local surface normal n and the building direction. h Layer d = -------------- cos ( α ) 2
and therefore
h Layer d ≤ -------------2
(2.2)
In the worst case of a surface normal to the building direction the maximum deviation could be half a layer thickness. Since the effect on the part-height described in (2.1) can also be viewed as a special case of (2.2) with α=0, the two effects do not add in terms of inaccuracy of the part. Depending on constraints for the geometry of the part, the theoretical deviations could be as low as zero (e.g. cube with height being a multiple of the layer thickness). For a general shape with no geometric constraints, the theoretical accuracy of a SFF-part is half of the layer thickness.
2.2. The Concept of SDM The basic concept of the SDM process improves upon the accuracy of the part by avoiding the stair-step effect, which is typical for SFF processes. Instead of using two dimensional cross-sections, built up to a fixed layer thickness, SDM uses true, three dimensional slices of the part. In addition to shaping the layers according to their three dimensional geometry SDM also provides for adaptive layer thicknesses. For local features with big changes in curvature the accuracy is improved by using thinner layers, for features with small changes in curvature layers will be kept thicker to optimize material usage and shorten processing times. Due to the consideration of three-dimensional layer geometries planning strategies are more complex than the ones used by SFF processes, but are still far less complicated than for conventional CNC operations. The additive nature of “growing” the object enables the part to be split into separate layers in critical spots. Therefore, planning remains straight forward and completely predictable.
2.2.1. The Geometric Principle of SDM To plan for the SDM process more detailed geometric information about the part is required than by SFF processes. Instead of a simple surface CAD model of the part a solid model is used, that also allows access to information such as surface boundaries of solids, boundaries of surface patches and surface normals, and that can determine whether vertices or edges are inside of surfaces or solids. One such CAD system called “NOODLES” has been described in [29]. Depending on the curvature of the surfaces of the model a set of blocks with varying thickness in the building direction is created. The criteria essential for the adaptive thickness will be described later in this chapter. The set of blocks is intersected with the model to create a set of slices (Figure
12
2.3.a)) which still contain the full three dimensional information for the part.
a) =
U
CAD Model
Model of 3D Slices
Set of Blocks
c)
b)
SDM-Part
Embedded 3D Slices
Figure 2.3: The Principle of the SDM Process For each of the slices the three dimensional geometry is then manufactured according to the principles outlined later in this chapter (Figure 2.3.b). Figure 2.3.c shows the part after removal of the support structure. No surface texture due to the stair-step effect is present, and the geometry of the original model is closely matched.
2.2.2. Surface Classification While conventional SFF processes build their layers in form of disks from a cross-sectional layer, surface information is not available and is not needed for the planning process. The only distinct features are flat overhangs, which can be manufactured without change in strategy because the previous layer, embedded in the support material, provides the necessary flat surface. The SDM process creates parts from three-dimensionally shaped slices. The three dimensional geometry of the “side-band” of each slice has an important influence on the process sequence for manufacturing the layer. The term side-band is explained in Figure 2.4. It refers to the surface of the slice excluding the topmost and bottommost faces which are perpendicular to the building direction. Those two faces, in case they exist (i.e. a pyramid would not have a top face), will be
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referred to as top face and bottom face, respectively. All surfaces belonging to the side-band are drawn shaded in Figure 2.4. side-band faces (shaded)
top face
direction of growth side-band faces (shaded)
bottom face (not visible)
Figure 2.4: “Side-Band” Definition
Five different types of surfaces can be defined with respect to the direction of growth: straight, undercut and non-undercut surfaces (Figure 2.5). The classification of the surfaces can be easily done by looking at the surface normals which point to the outside of the object. top undercut outward surface normals
straight
direction of growth
nonundercut bottom cross-section
Figure 2.5: Surface Classification Straight surface patches are parallel to the direction of growth. The surface normal of straight patches is perpendicular to the direction of growth. Undercut surfaces are surfaces which “point” downwards (in the sense of the direction of growth). Undercut surfaces are not visible from the top of the object. The angle between their outwards surface normals and the direction of growth is greater than 90 degrees. Non-undercut surfaces are surfaces that “point” towards the top of the object (again in the sense of the direction of growth). They are accessible from the top. The angle between the outward surface normal of non-undercut surface patches and the direction of growth is smaller than 90 degrees. Table 2.1 summarizes the surface type definitions. In normal applications, the direction of growth coincides with the positive z-axis. Therefore, the z-coordinate of the surface normal can be used
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for the surface classification. The types top and bottom are used after the part has been decomposed into layers, to identify the top and bottom face (if they exist) from the remaining faces, which belong to the side-band. surface type definition
angle α between surface normal and direction of growth
z-coordinate of outwards pointing surface normal
Straight
α = 90°
nz = 0
Undercut
90° < α ≤ 180°
nz < 0
Non-Undercut
0° ≤ α < 90°
nz > 0
Top (must be topmost face)
α = 0°
nz = 1
Bottom (must be bottommost face)
α = 180°
nz = -1
Table 2.1: Surface Type Definitions The simple example shown in Figure 2.5 to clarify the surface type definitions contains only linear faces. For non-linear CAD models the concept of surface definitions remains valid without change. Instead of classifying individual faces of the part surface, the non-linear surface of the part would be divided into patches by finding the parting curves. This problem is equivalent to the problem of finding the parting lines for automatic mold creation [30].
2.3. Basic SDM Manufacturing Strategies In this section an overview of the basic strategies for manufacturing layers with different surface classifications is discussed. The basic strategies cover the principle of SDM and do not contain any details about the deposition or shaping processes. They are meant to be a frame and will be expanded and adapted to specific deposition and shaping processes in Chapter 6. To manufacture an arbitrary part, the geometry will be adaptively split into layers (see Chapter 2.4.3), such that only three different scenarios have to be taken into account. First, a method is needed to manufacture a slice which contains only non-undercut surfaces along its side-band. Second, a method for creating a layer only containing undercut surfaces in the side-band is presented. Finally the two methods are combined to manufacture slices containing both undercut as well as non-undercut surfaces. Straight surface patches can be manufactured with both strategies. Depending on optimization criteria, they will be built according to either the strategy for undercut surfaces or the one for non-undercut surfaces. The principle of the three manufacturing strategies will be shown on a simple part which is made up from three layers (Figure 2.6). The first layer only contains non-undercut surfaces, the second only undercut features and in the third layer a combination of both occurs. Any part is typically started on a substrate or a substrate with a thin layer of support material to allow for easy removal. Since every layer is finished as a flat surface, those different starting conditions for the first layer will not affect the manufacturing sequences of the basic strategies.
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3
both
2
undercuts non-undercuts
1
Figure 2.6: Three Different Combinations of Surface Types
2.3.1. Manufacture of Non-Undercut Features Manufacturing layers containing only non-undercut features is a fairly uncomplicated task. On top of the substrate, or the previous layer, a sufficiently thick layer of material is deposited (Figure 2.7a). The material is then planed at the height of the current layer to remove excess material from the top. Excess material is also removed from the sides, and the side-band of the layer is shaped according to the 3D geometry (Figure 2.7b). After the slice has been shaped, support material is deposited in order to protect the current layer from subsequent operations (Figure 2.7c). In the final step the layer is planed to remove excess support material from the top and to provide a flat surface for the next layer (Figure 2.7d).
a)
deposit main material
b)
plane and shape main material
c)
deposit support material
d)
plane layer
Figure 2.7: Manufacture of Non-Undercut Surfaces
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2.3.2. Manufacture of Undercut Features Shaping undercut surfaces is a more challenging task. They are not directly accessible from the top and therefore shaping would require specialized tooling and non-generic procedures. However, for certain combinations of materials it is possible to create undercut shapes indirectly by first depositing and shaping the support material (the undercut surfaces would be non-undercut surfaces on the support material) and then using this structure as a mold to create the correct shape for the main material.
a)
deposit support material
b)
plane and shape support material
c)
deposit main material
d)
plane layer
Figure 2.8: Manufacture of Undercut Surfaces
The following steps would be required to create a slice containing only undercut surfaces: First the support material is deposited on top of the previous layer (Figure 2.8a). Excess support material is removed by planing the top and shaping the side-band of the slice (Figure 2.8b). the main material is then deposited into the support material mold (Figure 2.8c). Finally, after planing the top, the layer is completed (Figure 2.8d)
2.3.3. Manufacture of Arbitrary Layers Arbitrary layers containing both undercut and non-undercut features will be manufactured by combining the previous two methods. Those parts of the side-band with undercut surfaces are created by first depositing and shaping the support material, the part containing the non-undercut sur-
17
faces is manufactured by depositing the main material first. To accomplish this, the geometry of the slice and the surrounding support material have to be broken into smaller pieces. This process, called compact splitting, is described in detail in chapter 2.4.5.
a)
deposit and shape support material
b)
deposit and shape main material
c)
deposit support material and plane layer
Figure 2.9: Manufacture of Arbitrary Layers With the layer split into several segments its creation follows the logic of the previous two sections. Layer 3 of the example shown Figure 2.6 would consist of the support material, broken into two segments, and the slice of the main material, which does not have to be split. In the first step the part of the support material which is necessary to create the undercut surfaces of the side-band will be deposited and shaped (Figure 2.9a). Then the main material is deposited, planed, and the non-undercut surfaces of the side-band are shaped (Figure 2.9b). Finally the layer is completed by depositing the remaining part of the support material and planing the top (Figure 2.9c).
2.4. The SDM Planning System The manufacturing strategies in the previous chapter show the principle of building layers with simple, basic features. In most cases the geometry of an arbitrarily shaped part is more complex and the part has to be decomposed into segments, each of which can be manufactured with one of the three basic strategies. Then, according to predefined process plans a sequence of operations has to be established for each segment. Finally machine code to manufacture each segment will be derived. According to this functionality the planning system for the SDM process consists of three groups
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of modules (Geometric Decomposition, Process Planning, Machine Code Generation). The individual modules in each group are shown in Figure 2.10.
CAD Modelling System
Surface Classification Module
Adaptive Slicing Module
Precedence Ordering Module
Compact Splitting Module
Geometric Decomposition Process Definition and Scheduling Module
Path Planning Module
Process Planning CNC Code Translator
GMF Code Translator
Other Code Translators
Machine Code Generation
SDM Process Control Interface
Figure 2.10: Architecture of the SDM Planning System
During a planning session a CAD model is analyzed and processed by the different modules. The flow of information and the sequence of modules during a typical planning session is shown in Figure 2.11. In the Geometric Decomposition segment of the planner the CAD model is analyzed and decomposed into manufacturable segments. First, the surface of the CAD model is classified into patches of different surface types by the Surface Classification Module. The classified surface is then processed by the Adaptive Slicing Module to split the part into a set of layers. The Precedence Ordering Module analyses the segments of each layer and tries to establish an order in
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which the different segments can be manufactured. In case of conflicting orders one or more segments have to be divided. This is done by the Compact Splitting Module, which tries to resolve the ordering conflicts by splitting one of the segments involved in the conflict into sub-segments that do not pose any problems. The loop between the Precedence Ordering Module and the Compact Splitting Module continues until a manufacturing order can be established. In the Process Planning portion the Process Definition and Scheduling Module uses process definitions, applicability criteria and process data to create the sequence of operations necessary to build each segment of the part. Combining the sequence of operations from all segments results in the Master Plan, which is used to manufacture the whole part. Operations requiring a geometry specific path are further processed by the Path Planning Module. Depending on process and tool data and the geometry of the respective segment a tool path is generated for each operation. The final group of modules (Machine Code Generation) contains code translators for each individual machine involved in the manufacturing process. The “generic” path generated by the Path Planning Module is translated into command files which can be used to drive the individual machines in the manufacturing cell.
2.4.1. The CAD Model The type of CAD model used as the input for the SDM planning system has an important effect on how the different tasks are handled inside the modules. To date, there is a variety of CAD packages available, but not all of them make enough geometric information transparent to the user and a language library to manipulate created models is not always provided. The particular package which has been chosen for an experimental implementation of the planning system is called NOODLES [31]. It uses a linear, non-manifold boundary representation (b-rep) to construct all entities. Objects of all dimensions (vertices, edges, faces and solids) are represented the in the same data structures, thus enabling unrestricted operations between objects of different dimensions (e.g. edges can be intersected with faces or solids or the intersection between two faces which share one edge would have that edge as the result). All objects are represented by their bounding entities (e.g. edge by two vertices) augmented with a flag indicating the directional sense or the notion of inside or outside. Simple operations available to manipulate objects are translation, rotation, scaling and the binary operators union, subtract and intersect. A set of system primitives is available to create models of simple geometric objects, such as vertices, edges, faces, cubes, linearized cylinders or spheres and others. An arbitrary geometric model would be constructed from those primitives manipulated by the different operations. Other operations allow the interrogation of the b-rep structure of objects to find bounding or adjacent entities, directions or angles between different entities. The NOODELS CAD system is available as a C program library, and was therefore ideal for developing the SDM planning system. The functional descriptions of the different planner modules in the following paragraphs are based upon linear models from the NOODLES CAD system. Whereas the basic functions of the modules will remain unchanged for different CAD systems, the linearity of the NOODLES models imposed some restrictions and limits to the functionality of some modules. Especially the performance of adaptive slicing and compact splitting will change when a non-linear CAD system is
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CAD Model
Surface Classification CAD Model w/ Surface Classification Min/Max Layer Thickness
Adaptive Layer Splitting Layer Model w/ Surface Classification
Compact Splitting
Sets of Compacts w/Surface Classification and Precedence Graph
Precedence Ordering Manufacturable Compacts w/Surface Classification
Process Definition and Scheduling Compacts w/Surface and Tool Info
Process Description Process Data Applicability Criteria
Process Database
Path Type Path Data
Path Planning Tool Data
Master Plan Tool Path
Machine Code Translation
Machine Code
.......
Figure 2.11: Flow in the SDM Planning System
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Tool Database
used. A non-linear version of the planning system has not fully been developed yet, but, an overview of the functional change of the modules will be given in Chapter 2.4.6. For the ease of explanation it will also be assumed, that the direction of growth coincides with the positive z-axis of the model coordinate system. References to the height of the buildup would therefore be the z-coordinate at the particular level and the decomposition of the layers would be done in parallel to the x-y plane. Having the growing direction along the z-axis of the model coordinate system also simplifies the generation of trajectories for the individual machines in the manufacturing cell, since machine coordinate systems normally have their z-axis pointing upwards.
2.4.2. The Surface Classification Module The surface classification module takes the linear CAD model and classifies all the faces in the model. Each face of the surface will be assigned an attribute according to its orientation with respect to the direction of growth. The classification scheme has been described in Table 2.1. For this classification only the types straight, undercut and non-undercut will be assigned. To classify the surfaces of non-linear models, first the boundaries of surfaces with different orientation have to be found. Then the attributes can be assigned to the surface patches.
2.4.3. The Adaptive Slicing Module This module takes the model of the part with the surface type attributes, which have been assigned by the surface classification module and produces a set of new models representing the decomposition of the part into layers. For the decomposition purpose faces with type top and bottom are treated like non-undercut and undercut faces respectively. To determine the heights at which the model has to be subdivided into layers, the edges on the surface of the model are examined. Every edge which serves as boundary for faces of different classification are presumed to cause a problem for the SDM process and are marked as problem edges. The problems are caused by transitions from undercut to non-undercut faces or vise versa, and the associated faces have to be put into separate segments for the manufacturing process. If the problem edge is perpendicular to the direction of growth (i.e. parallel to the x-y plane), the zcoordinate of its vertices are put in the list of heights at which the model has to be split into layers. Since the layer partition cuts through the problem edge (which lays inside the partition) the faces with different surface classification will belong to different layers and no longer pose a problem for the manufacturing process (Figure 2.12 a and b). Faces associated with problem edges that are not parallel to the x-y plane cannot be cleanly separated by a partition parallel to the x-y plane. For those edges the z-coordinates of both vertices are put into the list of heights. This does not eliminate the manufacturing problem, but limits the conflict to one single layer (Figure 2.12 c). To resolve this conflict a partition along the problem edge is later inserted in the compact splitting module. Faces classified as straight require additional processing after the initial decomposition into layers. From a manufacturing point of view straight surfaces can be built using either the method for
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a)
b)
planar transition non-undercut - undercut
planar transition undercut - non-undercut
c) transition edge non-planar conflict local to single layer
Figure 2.12: Adaptive Layer Slicing undercut or the one for non-undercut surfaces. To optimize the process several factors can be taken into account to decide upon which method is used for building the straight surfaces. In general it seems preferable to keep the number of layers to a minimum to shorten the build time for the part. Therefore straight faces would be built according to the strategy used for its neighboring faces. Wherever straight faces are surrounded by faces of different classifications, the method which directly shapes them would be chosen (i.e. non-undercut). The top and bottom face of the layer do not have to be considered for this evaluation. After the reclassification of the straight reclassified as non-undercut
non-undercut straight undercut straight
reclassified as undercut
undercut
original layer
reclassified and subdivided layer
Figure 2.13: Reclassification of Straight Surfaces edges, the layer is checked for edges with adjacent faces of conflicting type to determine if the layer has to be subdivided. As shown in Figure 2.13 straight faces can “hide” a conflict from being detected at first. After reclassification of the straight faces the conflict can easily be
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detected and the layer is subdivided. Two parameters depending on data from the process database associated with the particular process definition, which will be used to manufacture the part, further influence the adaptive generation of layers. Depending on the deposition process, the shaping process and the materials of a part, it is only possible to manufacture layers within a range of thicknesses. Limiting factors for the maximum thickness of a layer are mainly the maximum depth that a material can be deposited without requiring additional processing steps (such as relieving of internal stresses or surfaceplaning to reduce or avoid voids in the deposit) and the maximum depth at which the side-band can be shaped (i.e. cutting length of the cutter with the shortest cutting length). The minimum layer thickness is influenced by factors like the tendency that thin layers could be peeled away (e.g. by grit-blasting operations) and by considerations, that a certain minimum layer thickness is desired to keep the total number of layers and the build time within reasonable limits. Layers that fall outside the process imposed range of thicknesses have to be geometrically adjusted. Layers exceeding the maximum layer thickness are simply subdivided into equally thick layers, so that each of the layers is thinner or equal the maximum layer thickness. For layers
maximum layer thickness layer split into sublayers of equal thickness
minimum layer thickness a)
b)
layers combined for minimum thickness
c)
geometry adjustment
Figure 2.14: Maximum and Minimum Layer Thickness Adjustments below the minimum layer thickness the adjustment cannot be done without modification to the geometry of the part (which ultimately caused the thin layers). From the thin layer on upwards, as many layers are combined into one layer as are necessary to reach or exceed the minimum layer thickness (Figure 2.14). To adjust the geometry of the side-band a variety of approaches are possible. The ideal solution would be an approximation (a) of the side-band geometry. Especially in a linear CAD environment, because of a multitude of linear facets, this can be a very difficult and computationally expensive task. Another approach would keep the geometry of the thickest of the original layers (b) and build the preceding or following layers as extrusions of the bottom and top face of the thicker layer. Next to the approximation method this would most likely result in the second best accuracy, but problems could result depending on which 3D shaping strategy will be chosen (see Section 2.4.9.2). The third approach for modifying the side-band geometry of the
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combined layer is the least accurate one, but also the one resulting in the least amount of geometric complications. It is similar to the second approach with the exception that the exact geometry of the topmost (c) of the original layers is kept and the rest of the combined layer is the extrusion of the bottom face of the topmost original layer.
2.4.4. The Precedence Ordering Module In its nature SDM is a multi-material process. Because of each part being encased in a sacrificial support material the process deals with more than one material even for homogeneous parts. Whether being from a multi-material part or not, each layer consists of several segments with different materials. In this module the order in which the different segments have to be manufactured according to the basic strategies is analyzed. In case of conflicts the segments have to be subdivided into entities called compacts. When a layer is dissolved into its compacts, each compact meets the following definition: A compact is a segment1 of a layer composed of only one material. The surfaces adjacent to each of the other compacts must only be of one type (undercut or non-undercut), i.e. the surfaces adjacent to some of the other compacts in the layer can only be undercuts, the surfaces adjacent to the remaining compacts can only be non-undercuts. Compact A
Compact B
Compact A precedes Compact B
Figure 2.15: Compact Precedence Rule The compact precedence rule is a conclusion from the obvious fact, that those compacts, that are adjacent to undercut surfaces, have to be deposited before the current compact can be built. Compact Precendence Rule: A compact A precedes another compact B if any surface between compact A and compact B is classified as undercut with respect to compact B. Using this relationship, a graph can be constructed which determines the order in which the compacts have to be manufactured. The resulting graph is then examined for precedence loops. A precendence loop indicates a conflict in the manufacturing order of two or more compacts, i.e. one or more compacts appear multiple times in the list, indicating that they should be manufactured 1. Incidentally a compact could consist of more than one “piece” (solid),e.g. multiple island features, if the surface geometry permits manufacturing those features in the same step.
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before, as well as after some other compacts. If no precedence loops exist, the planning process can proceed with defining and scheduling the manufacturing process. In case of precedence loops the compacts have to be further subdivided, until the precedence graph is without loops. This manipulation is the responsibility of the compact splitting module.
2.4.5. The Compact Splitting Module The compact splitting module takes the model of a layer and the precedence graph produced by the precedence ordering module, and resolves the scheduling conflicts by subdividing the compacts in the layer. This operation will also include resolving the problems caused by edges non parallel to the x-y plane with adjacent faces of different type, which were left after the adaptive slicing procedure. The approach used by this module to resolve the conflicts requires identifying a compact involved in a precedence loop and subdividing it in order to break the loop. Figure 2.16 shows an example for a layer with six compacts. For each compact an analysis of its predecessor and successor compreceding compacts 3, 5 1, 3, 6 1, 2, 3, 5, 6 3 1, 3, 5
succeeding compacts
compact
3
2, 4, 6 4, 5 1, 2, 4, 5, 6
1 2 3 4 5 6
eC 5
1
2
5
4
precedence loop
1, 2, 4, 6 2, 4
precedence graph with conflict
precedence list 3
6
5-1
1
6
2
5-2
4
loop resolved by splitting compact 5 Figure 2.16: Compact Precendence Graph pacts is listed. The precedence graph resulting from this list shows a loop for compact 5. To resolve the problem, the compacts enclosed inside the loop (1-6-2) can be combined into one compact (eC). Now the problem compact (compact 5) can be subdivided into two subcompacts sC1 and sC2 which meet the following three conditions: 1. The surfaces of sC1 which are adjacent to eC are only undercut surfaces with respect to eC. 2. The surfaces of sC2 which are adjacent to eC are only non-undercut surfaces with respect to eC.
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3. The new surface generated between sC1 and sC2 imposes no additional precedence conflicts, i.e. is only an undercut (or straight) surface with respect to sC2. To decompose the problem compact, the compact splitting module computes a parting plane that splits the compact into its subcompacts sC1 and sC2. The current approach involves identifying the common boundary between the problem compact and the combined compacts enclosed in the loop eC. The nature of this boundary determines into how many pieces the problem compact has to be split. The surface adjacent to eC is divided into patches of different type (or groups of connected faces of the same type). The problem compact is then broken into pieces, so that each piece contains exactly one of the patches. The pieces with patches of type non-undercut (with respect to the problem compact) make up compact sC1, the pieces with patches of type undercut (with respect to the problem compact) make up compact sC2. To determine the internal boundary for splitting the problem compact, the current implementation uses the Medial Axis Transform1 (MAT) [32] of the intersection of the top and the bottom face of the layer. The segment of the MAT connecting the end-points of the common boundary is extruded as a straight wall to serve as the parting surface inside the problem compact. The current implementation of the compact splitting module is not fully developed. Although the principle of identifying and splitting the surface adjacent to the compacts inside the loop will resolve the precedence problem, finding a parting surface inside the problem compact using the MAT is not fully stable. Cases where no intersection exists between the top and the bottom surface have to be specially handled. For certain cases better geometries, such as simple rectangular faces, instead of more complicated faces extruded from the MAT, produce simpler compacts which are easier to manufacture. Other cases again, such as an inclined cylinder or an inclined hole, could be split by a plane which is defined by the edges which are the partition between the undercut and non-undercut surface patches of the cylinder or hole. To improve the current version of the compact splitting module a variety of different approaches, each handling a special set of cases, needs to be implemented. Some of them could be used based on a simple set of application criteria, the others could be invoked on a trial and error basis (some methods might not produce a result) and their results could be evaluated to find the best approach.
2.4.6. Geometric Decomposition for General (Nonlinear) CAD Models Linear modelling systems approximate nonlinear geometries by discretizing them with a multitude of linear entities. Increasing the accuracy of the CAD model by finer discretisation leads to models containing many tiny surfaces. As a result of intersections between discretized nonlinear geometries and of triangulated linear surfaces, which are in addition subject to mathematical inaccuracies due to the limited precision of computer mathematics, many transitions between faces of different surface type can be present in the model. This can cause severe problems, if the transitions occur that frequently, that the required layer thickness is much lower than the minimum layer thickness, and/or the compact splitting operation results in too many compacts to keep the 1. The medial axis (MA) for a 2D object is defined as the locus of the centers of all maximal discs inside the object. A maximal disc is one which can not wholly be contained by another disk. The MA is essentially a skeleton or the symmetric axis of an object.
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layer build time within reasonable limits. If this is the case, producing the layer essentially results in building the part according to the SFF principle, with the minimum layer thickness as the slicing thickness and each layer being built up as a 2 1/2 dimensional slice. Using a nonlinear modelling system can greatly improve upon this problem. Nonlinear solid models are in general constructed from nonlinear surface patches and avoid the discretisation into many tiny surfaces, therefore keeping the number of entities to construct the model to a minium. Unnecessary and unwanted transitions between surfaces with slightly deviating angles will be avoided. Instead, in the non-linear system the surface can be classified into different segments according to the true, desired geometry of the model. In addition, the geometric decomposition can be improved by combining the layer and compact splitting tasks. Using the classified surface patches as one part of the boundaries, new non-linear surfaces can be constructed inside the model to facilitate the decomposition into manufacturable compacts. The new surfaces would use the transition curves between surface segments of different classification as part of their boundaries. The hulls defined by the new and original surfaces serve as the boundaries for the compacts of the part. The newly created surfaces must not impose any additional manufacturing problems, i.e. they must be of type non-undercut with respect to the compact which comes earlier in the list. Whenever a compact exceeds the maximum layer height, the whole part can be split by planes perpendicular to the building direction.
2.4.7. The Process Definition and Scheduling Module The process definition and scheduling module takes each compact of a part and establishes a set of operations necessary for manufacturing according to a specific process definition. Following the order of layers and compacts from the geometric decomposition, the operations for all compacts are combined into the master plan, which provides the correct sequence of manufacturing instructions for the whole part. A variety of pre-established process definitions for different applications, materials and manufacturing processes, which are involved in the production of the part, are stored in the process database. Each process definition is a rule based sequence of manufacturing steps for a single compact. The left-hand side of each rule specifies an applicability criterium, which determines if a particular operation is needed for the current compact. The right-hand side of the rules defines one or more operations and their parameters which are used in the manufacturing sequence of the compact, if the applicability conditions are met. Variables used in the applicability criteria and for the parameters of the operations are layer thickness, layer number, number of compacts in the current, previous or next layer, compact number, and the material of the current, previous or next compact. The applicability criteria use a syntax similar to those of many programming languages. Possible constructions are relations between variables and process parameters from the process database, list-lookup functions and multiple executions for items in a list. The operations on the right-hand side can be one-step operations or substrategies, which are (similar to subroutines) a longer sequence of processing steps, which could be repeated throughout the process definition.
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Rule syntax: (if/for) applicability criterium do operation or substrategy (parameters, variables) Applicability criteria: • if variable relational operator process parameter or variable relational operators: , equal, is element of (list) • for items in process parameter or variable (must be list) EXAMPLES: if thickness of current layer > maximum thickness do special operation (model of compact, thickness of current layer) if material is element of stress_reliefing_materials do stress relieving procedure (material) for items in cutter_list do cut the contours (model of compact, cutter, material)
Detailed rules for different process descriptions which have been developed and used for parts manufacturing are discussed in Chapter 6.5. The master plan is the combination of the process steps, which have been developed for each compact according to the process description. It is used by the SDM process control interface to guide the manufacturing of the part. In addition to the individual operations it also contains the commands for transporting the part and for downloading trajectories to the individual cells (see Chapter 3). Operations requiring the creation of trajectories for cell according to a geometry specific path are further processed by the path planning module.
2.4.8. The Path Planning Module The path planning module takes a geometric model of a compact with additional parameters from the process planning and scheduling module and creates a geometric model of the trajectory, which a certain tool has to follow in the manufacturing operation. The information provided in addition to the geometric model consists of the type of operation requiring the path, material assignment, specific tool information (e.g. size of cutter for cutting operation), and when applicable the request for a specific type of path. In the process database a variety of path geometries are stored in form of rules. Augmented with process and material specific data from the process and tool databases (e.g. trajectory speed, distance between trajectory segments, standoff) the “generic” path geometry is adapted for a specific compact and process. The following two sections describe the basic characteristics of typical “generic” pathes, that are stored in the process database, with respect to shaping or deposition processes.
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2.4.9. Path Generation for Cutting Processes Manufacturing a part layer by layer, with each layer being split into compacts, allows for a generic strategy, independent of part complexity, to generate cutting trajectories. Per definition, compacts with undercut surfaces will be built after all compacts of the same layer which are adjacent to those surfaces have been built before. This ensures, that only non-undercut features have to be shaped. To prevent previous compacts of a given layer from being destroyed by cutting operations, the model supplied to the path generation for cutting processes has to consist of the union of the current compact with all previous compacts of the current layer. The task of shaping the combined compacts can then be split into two subtasks. The first task will contour the shape with constant cross-section along the z-axis (2D contouring) to remove most of the material. The second task (3D shaping) will then create the true 3D geometry of the combined compacts. 2.4.9.1. 2D Contouring Due to the nature of compacts the combined model of the current and previous compacts of a layer does not contain any undercuts. The bottom cross-section of this model can therefore be used to derive the trajectories for 2D cutting. To create the path for a certain cutter, the boundaries of the bottom cross-section have to be offset by the tool radius. A simple way of doing this is to union the cross-section with surface patches, whose width is the tool radius, and circles (radius of circles = tool radius) for convex corners (Figure 2.17). The boundary of the new surface is the path along which the center of the cutter has to be moved. this method automatically adjusts the geometry of the cross-section to prevent cutting into features, such as tiny slots, which are to narrow for a specified cutter diameter. Other methods [20] for cutter offset generation shifted the boundaries and then corrected for inconsistencies in the geometry of the path. To increase the quality of the surface finish of the cut, rough-cutting operations are performed before the final cut. These operations cut a specified distance away from the actual surface of the part. Generating the path for rough-cutting is done by adding the specified distance to the tool diameter in the offsetting procedure. original cross-section
path for 2D contouring
mow-path
tool radius
Figure 2.17: Path Offsetting for 2D Contouring
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In order to economically remove all the unwanted material in a layer and to cut the contours of the compacts as accurate as possible, different size cutter are to be used. If the planning system only adjusts the geometry for the cutter offset and does not accurately track the material that has been removed, a sequence of cutters has to be used, where each smaller cutter is at least half of the diameter of the previous one, in order to ensure, that all the unwanted material is removed. A sophisticated version of the path planning software would include a mechanism for optimizing cutter usage, therefore saving cutting time. The biggest cutter would contour the geometry of the cross-section and remove most of the material. The remaining material, which has to be removed, could be analyzed for maximum width, the next smaller cutter could be chosen accordingly (it could be smaller than half the diameter of the previous one, thus saving time by not using as many cutters for a certain accuracy) and only cut in spots were material remained (tiny slots, concave corners). Parameters (and dependencies) for cutter path generation: • tool diameter and type1 • feedrate (material, tool diameter and type, type of cut2) • speed [rpm] (material, tool diameter, cutter type, type of cut) • max. depth per cut (thicker layers have to be cut in multiple pathes) • max. tool depth (this limits the max. layer height)
Other cutting strategies related to the 2D contour cut are planing operations and “mow’-cutting. 2.4.9.2. 3D Shaping After the initial 2D contouring a variety of different strategies are available for three-dimensionally shaping the side-band of a set of compacts. If 5-axes NC-equipment is available, the surface can be milled with the side (peripheral cut) of a cylindrical cutter (Figure 2.18a). The part (or the cutting head) would be rotated, so the surface normal is perpendicular to the cutter axis. the second approach possible with 5-axes equipment is face-milling of the surface (Figure 2.18b). The surface normal and the cutter axis are in parallel. Figure 2.18 c) and d) show a “zig-zag” strategy for cylindrical and spherical cutters, which can be used without 5-axes capability. The cutter is traversed up and down along the surface. In between the up and down strokes, the cutter is advanced along the top or bottom contour in small increments. For calculating the toolpath, different offsetting strategies have to be used. For 5-axis peripheral milling, the surface needs to be offset by the tool radius along the surface normal (Figure 2.20a). A strategy similar to the one for offsetting contours in the 2D case has to be used to avoid cutting into tiny features with larger cutters. Offsetting for 5-axes face milling has to be done by offsetting the boundaries towards the inside of the surface and is dependent on the surfaces adjacent to the boundaries (Figure 2.20b). If the angle γ between the two adjacent surfaces, which is measured inside the object around the boundary, is bigger than 180° the boundary needs to be offset by the tool radius to the inside of the surface which is being cut. For angles γ smaller or equal 1. tool type: tool material and geometry 2. type of cut: face, slot, peripheral; rough, finish
31
a) 5-axis peripheral
b) 5-axis face
c) cylindrical zig-zag
d) spherical zig-zag
Figure 2.18: 3D Shaping Strategies 180° no offsetting is necessary. For zig-zag cutting with a cylindrical cutter the center of the cutter is offset by the cutter radius along the horizontal (Figure 2.20c). For linear faces, the offset amount d can also be determined along the surface normal, and is given by the angle α between the tool axis and the surface normal (d = R · sin α). For a spherical cutter the center-point of the tool can be used as the reference point for the cutter trajectory. The offset along the surface normal is then simply the radius of the tool and independent of the surface inclination (Figure 2.20d). Using 5-axis peripheral milling can in some cases pose a problem because of the geometry of the side-band of the compact. For adjacent faces, where the angle γ around their common boundary, measured inside the part, is bigger than 180°, a section of the part would be cut away, when one of the faces is being shaped (Figure 2.19). Possible solutions to this problem are the approximation of the geometry by a single surface, the use of a different strategy to shape one of the surfaces (i.e., the one that is further away with respect to the approaching cutter), or imposing an additional constraints to the adaptive slicing and compact splitting algorithms, to eliminate transitions between such surfaces by creating additional layers and compacts for those features.
cutter
this portion would be milled away
γ
approximation
Figure 2.19: Interference with 5-Axis Peripheral Cutting
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R = tool radius
boundary of face
tool path
γ
γ
no offsetting necessary offsetting necessary
R
tool path R
a) 5-axis peripheral
b) 5-axis face cutter axis
surface normal α R
R
tool path
tool path
R
c) cylindrical zig-zag
d) spherical zig-zag
Figure 2.20: Path Offsetting for 3D Shaping
Choosing a 3D shaping strategy is influenced by a variety of variables. Geometric constraints, such as explained in the previous paragraph, can limit the use of 5-axis peripheral shaping, and 5axis face cutting can be limited by the size of the surfaces (the cutter does not fit into tiny faces that need to be offset). Accessibility constraints of the NC equipment can also play a role. With common cutters and chucks, tilt angles for the part or cutter head are typically only possible up to approximately 45° before parts of the chuck interfere with the flat surface of the current or previous layer. The use of zig-zag strategies is time consuming because of tedious pathes with up and down motion and only tiny increments in the x-y plane. Constraints of keeping the build time within reasonable limits can influence a preference for the 5-axis strategies. Additional parameters for 3D cutter path generation: • 3D type (force certain type of 3D cut) • allowable zig-zag surface roughness • type of zig-zag cutter (cylindrical or spherical if applicable)
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2.4.9.3. Geometric Accuracy of Cutting Processes In general the geometric accuracy of cutting is limited by the diameter of the cutter. While convex corners and surfaces can be cut to accurate shape, concave corners will appear filleted with the radius of the tool. Concave surfaces can only be traced accurately with cutters whose radius is smaller than the curvature radius of the surface.
tool axis d
R
r
surface normal
R sin α
α
Figure 2.21: Surface Roughness of Zig-Zag Cutting
For 3D zig-zag cutting, the final surface roughness is determined by the spacing between the up and down segments, and the radius of the tool. For cylindrical cutters, the angle between the surface normal and the cutter axis has also an influence. If the cutter movement along the surface is according to an “up-over-down-over” path, the machined surface will show a wave pattern with an elliptical cross-section (Figure 2.21). The equation for the elliptical shape is as follows: 2
2 y 2 x + ------------= R 2 sin α
(2.3)
with R being the radius of the tool and α the angle between the surface normal and the tool axis. The roughness r of the surface can be determined by solving (2.3) for y and taking the difference of y(0) and y(d/2) where d is the distance between the up and down segments. For the cylindrical cutter the surface roughness results to 2 d r = R ⋅ sin α 1 – 1 – --------2- 4R
for
0 ≤ d ≤ 2R
(2.4)
Given a specified surface roughness r, the required path distance d is 2
2r r d = 2R --------------- – ------------------2 R sin α R sin 2α
with
0 ≤ r ≤ R sin α
(2.5)
For sin α = 1 (2.4) and (2.5) show the relationship between path distance and surface roughness
34
for zig-zag shaping with a spherical cutter. 1 -3
.9
.4
x 10
α =90°
.2
α =90°
.8 .7
1
α =70°
.8
α =50°
.6
.6
α =30°
α =70°
.4
--rR
.5 .4
.2
α =10°
α =50°
0
.3
α =30°
.2 .1
α =10°
0
d⁄R
Figure 2.22: Path Spacing vs. Surface Roughness To decrease the surface roughness of zig-zag milling with cylindrical cutters on NC-equipment with 5-axis capability, the angle between the surface normal and the cutter axis can be decreased by tilting the part or cutting head. For α = 0° cylindrical zig-zag cutting is identical to 3D face milling. For α approaching 90° the surface roughness of (2.4) is further decreased by parts of the wave pattern being cut away at the down-over motions. At α = 90° the side of the cutter traces the surface and the cut would be identical to 3D peripheral milling.
2.4.10. Path Generation for Deposition Processes Depending on the distribution of their particle stream, the deposition processes used in SDM can be divided into two classes. The spray processes on the one side have very fine, statistically distributed particles, with a relatively large deposition area and do not require trajectories which are derived with every detail of the geometry. The deposition processes with discreetly deposited particles or streams, such as welding or microcasting, on the other hand, require controlled and accurately placed trajectories. Most spray deposition processes have a particle stream with a near Gaussian distribution, with a standard deviation ranging from approximately 5 to 20 mm. Assuming a perfect gaussian distribution, uniform deposit, with exception of the edges, can be achieved by using a meandering trajec-
35
tory with a path spacing smaller than 1.5 times the standard deviation. Corrections for asymmetries in the deposition pattern can be made by measuring the real distribution and adjusting the tilt angle of the torch [33]. Lower deposition on the edges of the meandering path (the neighboring trajectory lines are missing) requires overspray, i.e. the sprayed area is larger than the actual cross-section. Typically, meandering lines, alternating between traversing in the x and y direction (x-y cross-hatch) are used for spray deposition. To derive the trajectories, the projection of the compact onto the x-y plane is “grown” (i.e. offset similar to 2D cutter path) by the overspray distance. The cross-hatch pattern can either be bound by the resulting geometry, i.e. some lines would be longer, some shorter, or simply the bounding box is used to create the pattern (Figure 2.23a). Additional material deposited by the overspray or deviations in thickness due to unsymmetrical distributions do not impose a problem, since it is removed during the cutting operations. The generation of trajectories for deposition methods with discrete streams, such as welding or microcasting, are more critical, since the integrity of the deposit depends on a strictly controlled path. while the area of particle distribution is decreased, the deposition rate per path is much higher than for spray processes. Keeping accurate path spacing and torch angles is therefore important to avoid gaps in the deposit or overlapping of the individual lines. Torch angles could also be dependent on the geometry of the compact, i.e. to deposit along a previously deposited compact, the torch might use a different angle. Unsymmetrical deposition properties might also restrict the deposition to traverse only along one axis, or even to only one direction. Typical trajectories for discrete deposition are created by offsetting the projection of the compact. Depending on the behavior of the particular process, offsetting could be done to “grow” as well as “shrink” the area, since discrete deposits tend to have a significant width. When shrinking the deposition area, special care has to be taken, so that small features do not disappear, and no material would be deposited. Typical path geometries used for discrete deposition processes are meandering lines along one axis, with fast travel to advance between the lines (Figure 2.23b), unidirectional deposit (Figure 2.23c) or a strategy where first the perimeter is deposited, and then the inside is filled with any of the other patterns (Figure 2.23d).
a) cross-hatch
b) uni-axial
c) uni-directional
d) perimeter+ fill
Figure 2.23: Patterns for Deposition Trajectories
In certain cases the deposition strategy can also involve more processing steps, than the deposition process alone. To prevent the creation of voids with slightly instable trajectories of the deposition stream, first only every second line of a unidirectional path was deposited with slightly
36
decreased path spacing. Then, a cylindrical groove, slightly smaller then the width of one deposited line, was cut in between the lines, to remove any stray particles. To finish the deposit, The grooves were filled, now guiding any slightly stray particles into the grooves. Parameters (and dependencies) for deposition path generation: • path type (application) • path spacing (torch, material, operating parameters) • standoff (torch, material, operating parameters) • speed (material, operating parameters) • deposition per path (torch, material, spacing, standoff, operating parameters) • overspray/offset distance (torch, material, standoff, operating parameters) • operating parameters1 (torch, material, application)
2.4.11. Machine Code Generation Module The machine code generation module consists of various code translators, which take a path generated by the path planning module and transfers it into instructions which can be read by the individual machines in the manufacturing process. All information relevant to the process, including commands for switching certain processes on and off, will be contained within the path. To increase flexibility after the planning process is finished, each path is stored individually. Commands in the master plan enable the process control interface to download each path to the individual cells and arrange for their execution.
2.4.12. Conclusion and Future Developments The current linear version of the planning system shows the principle requirements for planning in the SDM process. A scheme and the basic elements for the geometric decomposition of a part with arbitrary geometry have been identified to create manufacturable segments and use the advantages of the SDM process to produce parts with geometries and properties, which are difficult to reach with conventional processes. A simple, rule based syntax to describe different process sequences, generic tool trajectories and a storage mechanism for process and tool parameters has been developed. Finally, a protocol with instructions and trajectory data was established to transfer the manufacturing information to the control system and the actual manufacturing stations. The current planning system was successfully used to prove the principle of the SDM process and to manufacture several test parts. However, there are many areas offering opportunities for improvement. Linear models of more complicated parts with tiny linear surfaces can lead to many unnecessary transitions because of geometric noise and can cause the geometric decomposition to fail. An non-linear CAD and planning environment, allowing fewer, bigger surface patches, can 1. i.e. settings on the deposition equipment, such as power, material flow- or feedrates, gas mixtures and flowrates.
37
eliminate this problem. The rule based system for the process sequence description needs to be extended into a meta language to allow more sophisticated and flexible process programming environment. The most challenging improvement lays in the combination of the SDM planning system with a complete design and advisor environment. Currently a specific process description and a small number of process parameters are manually chosen from the database. By including advisors and optimization strategies, the model of the part and criteria for the required part quality and properties could be supplied instead, and the optimum parameters and building strategies would be extracted from the database automatically. The use of feature recognition systems could further greatly enhance the efficiency and accuracy of creating the manufacturing plan.
38
3. Automated SDM Testbed Facility Shape Deposition is a layered manufacturing technology to incrementally build shapes and functional parts. Compared to existing rapid prototyping processes, SDM uses separate steps of material deposition and shaping to create layers, or portions of layers (compacts). Material deposition is not limited to one material, or one particular deposition process, thus allowing the creation of parts from a wide range of materials or multi-material structures. Different intermediate steps, for pre- or postprocessing of the material are required. To facilitate the usage of different processes, flexible and programmable manufacturing strategies are used for part manufacturing. To prove the concept of SDM and to develop the process, a modular manufacturing cell, consisting of individual subprocesses, has been built (Figure 3.1). In addition to a fully programmable manufacturing sequence, the modular setup allows reconfiguration and the incorporation of additional
Figure 3.1: Testbed SDM Facility at Carnegie Mellon University
39
processes.
3.1. Configuration of the Experimental SDM Cell The experimental configuration of the SDM cell consists of several subprocessing stations for material deposition, shaping and intermediate processing steps [34]. The growing part is built on a pallet, which is robotically transferred from station to station, to allow full flexibility of the configuration and position of the individual stations, as well as the manufacturing sequence. Currently, the manufacturing cell consists of the following subprocesses, which are positioned along a circle around the transfer robot: a deposition chamber, incorporating thermal spraying, microcasting (see Chapter 4), welding and deposition of waxes, a 5-axes CNC milling machine to shape each layer, grit-blasting for surface preparation, shot-peening for residual stress release, and a washing station to remove residual cutting fluids or residue from the deposition process. Figure 3.2 shows a schematic layout of the subprocesses and the robotic transfer system.
Figure 3.2: Subprocesses in the Shape Deposition Manufacturing Cell
40
3.1.1. Robotic Transfer and Palletizing System A custom built palletizing system is used to transport the part between the different subprocesses of the manufacturing cell. Each station is equipped with a pallet receiver mechanism, which will hold the pallet during the individual operations of the process. Where necessary (milling machine, tilt table in deposition chamber) the receiver incorporates hydraulic clamping at a force of 5.3 kN and indexing of the pallet with an accuracy of 5 µm (see Figure A.1 in the Appendix). A “Motoman K100S” robot is used for pallet transfer. The wrist of the robot is equipped with a special gripper/clamping mechanism to pick up and hold the pallet during transfer. The maximum possible part size is currently 300 x 300 mm with a height of 200 mm at a maximum weight of approximately 50 kg, limited by the size of the pallet, the height available on the milling machine and the payload of the transfer robot and the tilt/rotary table of the milling machine. Part size and weight, however, do not impose a limitation on the process. The manufacturing cell can be adapted to different part sizes by simply scaling the equipment for the individual subprocesses and the transfer system. Choosing a robotic transfer and palletizing system is ideal for the experimental setup, to allow the flexibility to rearrange the configuration of subprocesses, if required by changing process strategies. However, the time required for transferring the pallet from one station to another is comparably large. Slow speeds are required to avoid excessive forces, which could be exerted onto the receiver or pallet during pick up or drop off, and the trajectories required for the robot to reach into doors or to rotate the pallet into position are rather long. A shuttling system, with dedicated tracks and linear loading mechanisms could speed up the transfer process significantly.
3.1.2. Deposition Chamber The operations for material deposition take place inside a sound proof booth, to contain dust generated by some deposition processes and to reduce noise levels. An air removal/filtration system is connected to the chamber, to extract fumes, shielding gasses and dust particles created by spraying operations. Special consideration has to be given to avoid collection of unoxidized metal dust, to prevent possible explosions. The pallet is brought into the chamber by the transport robot through pneumatic doors. The deposition station (Figure 3.3) is divided into two areas, with two separate receiver mechanisms. The right hand side of the station is used for spraying operations, the left hand side for welding and microcasting (see Figure A.4 -A.7 in the Appendix). The receiver on the left has a hydraulic clamping and indexing mechanism and is mounted on a 2-axes rotary/tilt table, to allow deposition onto substrates at an inclined position. The deposition torches and guns are mounted on racks to the left and right of the pallet receivers. A “GMF-700S” robot, equipped with a pneumatic tool-changing wrist, is used to manipulate the individual deposition torches and guns. To deposit material, the deposition robot docks to one of the torches (see Figure A.4 in the Appendix), takes it off the rack and traverses according to trajectories created by the process planner over the growing part. The material feeders and power supplies are located upon a mezzanine above the acoustic chamber. The currently available processes on the side for spraying are a plasma-arc spray gun, for the deposition of metals, ceramics and plastics, two metal wire-arc spray torches, and a device for depositing low melting materials, such as tin or tin-zinc alloys. On the welding/microcasting side, currently two microcasters for the deposition of metal
41
droplets and a hot melt gun for waxes and glue are mounted. In a previous configuration, a MIGwelding process was also in operation, but was later replaced with the second microcaster. If necessary, additional deposition equipment can be easily added to the racks, and be incorporated into the manufacturing cell.
Figure 3.3: Deposition Chamber
3.1.3. Shaping Station In the current configuration of the manufacturing cell, a vertical CNC milling machine (“FADAL VMC 6030”) is used for shaping the individual layers of the growing part. A rotary/tilt table (“Tsudacoma TTNC-301”) was added to achieve 5-axes capability for three dimensional contouring (see Figure A.2 and A.3 in the Appendix). The machine is equipped with an automatic 21head tool changer and has a spindle power of 11.2 kW. Pneumatic doors allow the loading of the pallet with the parts handling robot. A hydraulic pallet receiver is mounted on top of the rotary/tilt table. Accurate location of the pallet is extremely important for the layered shaping process, to avoid steps in the surface of part due to inaccurate alignment. A positioning sensor (LVDT) is installed in the pallet receiver, to ensure correct seating of the pallet. Ideally, to avoid any misalignment of the pallet, the correct position and orientation of the pallet could be measured after the pallet is clamped on the receiver, and the cutting trajectories could be adjusted accordingly. Milling is currently the only shaping process which was included in the SDM system. Other shaping processes, such as electro discharge machining (EDM) or micro machining stations, could
42
potentially be incorporated in a future SDM cell.
3.1.4. Grit-Blasting and Shot-Peening Stations The grit-blasting and the shot-peening stations are identical, with the exception of the ballast used for operation. Conventional, manual blast cabinets have been modified to allow automatic operation and pallet loading and unloading. A pallet receiver has been installed on the inside of each cabinet, and a mechanism has been added to allow for the parts handling robot to manipulate the blaster nozzle over the surface of the part. To grit-blast or shot-peen a part, the transfer robot first loads the pallet into the cabinet through a pneumatic door. Then, the robot grabs the mechanism on the top of the cabinet to manipulate the blaster (peener) nozzle according to a part specific trajectory. The standoff distance between the nozzle and the part surface can also be adjusted by the robot. Finally, the robot puts the mechanism back into its resting position and retrieves the part from the cabinet (see Figure A.12 and A.13 in the Appendix).
3.2. SDM Process Control System To automate the experimental SDM system, a two-level control system has been established (Figure 3.4). A Sun 3-160 workstation is used to schedule the individual operations and serves as the main controller for the cell. The controllers for the individual stations (i.e., milling machine, parts handling robot and deposition robot) on the second level are connected to the main controller via Ethernet or RS-232 interfaces, and carry out all trajectory and low level control. The digital IO of the transport robot controller is used for switching and checking of statuses of the mechanisms for the pallet transfer (doors, clamping), as well as to control the grit-blaster, shot-peener and the cleaning station. The deposition processes are controlled by digital and analog IO of the deposition robot. Direct handshaking from the milling machine and the deposition robot to the transfer robot is used to indicate states for process initiation and pallet transfer. A strategy dependent master plan, which contains the sequence of manufacturing operations, and the layer dependent trajectory data are generated off-line by the SDM process planner (see Chapter 2.4) and downloaded to the main controller. The instructions of the master plan consist of a set of 5 simple commands for downloading, executing and deleting trajectory data on a certain station, or putting the pallet into or retrieving it from a certain station. A process control interface evaluates the individual commands of the master plan and initiates the appropriate action. The trajectories and IO commands for pallet placement operations are taught and permanently stored on the transport robot controller, and are only initiated by the control interface. The trajectories for the shaping and deposition operations are geometry dependent and need to be downloaded for each individual operation before they can be executed. In addition to the SDM process control interface a task planning platform for robotic arc welding [35] has been adapted to simulate to process. Using the master plan and trajectory data created by the SDM process planner, the entire manufacturing process can be simulated. Figure 3.5 shows a
43
captured image from a simulation of the final layer of the IMS-T1 test part (see Chapter 8.1.2). The upper left corner shows the process control interface and part of the master plan, which was used to create the part. The ongoing operation, placement of the pallet onto the milling machine, is highlighted in green. The left side of the picture shows the spray side of the deposition chamber, with the deposition robot holding the plasma-arc spray torch. The two partly obstructed stations in the back are the grit-blaster and the shot-peener, the station in the front is the parts cleaner.
Internet (Worldwide)
Process Planner DEC 5000/200 or DEC 3000 (Alpha)
RS-232
Ethernet
Cell Controller SUN 3-160
GMF - Server PC 486-33
RS-232
RS-232
CNC Controller
Motoman Cont.
GMF Controller
Milling Machine
Transfer Robot
Deposition Robot
DIO
DIO
DIO/AIO
Parts Cleaner Grit-Blaster Shot-Peener Transport System
Microcaster Plasma Spray Wire-Arc Spray Wire-Arc Spray Hot Melt Microcaster
Figure 3.4: SDM Process Control Diagram
44
Figure 3.5: SDM Process Control Interface and Process Simulation
45
4. The Microcasting Process 4.1. Introduction High quality, fully functional metal parts for engine components or tooling require, next to a high degree of accuracyand surface finish, superior mechanical properties. In order to produce those parts directly with SDM they must have strong interlayer adhesion and the deposited metal must be cohesive and dense with low oxide content and with a controlled microstructure throughout. In thermal deposition the capability to control the energy (i.e. both thermal and kinetic) of the molten material and the energy transfer to the underlying layers is critical to achieve these goals. On one hand, sufficient energy is desirebale to slightly remelt the surface of previously deposited material in order to promote adhesion between the layers through metallurgical bonding and to enable sufficient flow of the newly deposited material for full densification. On the other hand, the energy has to be kept to a minimum to prevent remelting of larger areas of previously deposited and shaped layers and to avoid penetration into the support material, thus destroying the shaping mechanism for undercut features. Controlling heat input, heat transfer and heat affected zones further minimizes the buildup of large temperature gradients to prevent distortion of the support structure and previously deposited materials. Trade-offs in conventional welding or thermal spraying processes between material quality and requirements for layered manufacturing limit their use with respect to their capability for achieving these goals. Excessive heat input on one side and the lack of metallurgical bonding and full densification on the other required the development of a novel deposition process called microcasting, which can satisfy those conditions.
4.1.1. Thermal Spraying To deposit material by spraying, a variety of techniques are available. While some have rather large heat transfer into the substrate (such as the ‘Osprey’ process, flame or detonation spraying), plasma or arc spray processes are characterized by a relatively low heat-input into the target material. The energy of the sprayed material stream is stored in form of thermal energy (i.e. latent heat of the molten droplets) and kinetic energy, which is transformed into heat upon impact of the droplets on the surface. The sprayed particle stream exits the nozzle of the spray torch in form of a mist of relatively small droplets (typical droplet diameter from submicron to approximately 50 − 100µm) with a near Gaussian distribution. The tiny droplets do not store much heat, and because of a high surface to volume ratio they cool off quickly during their flight. Typical temperatures for droplets upon leaving the spray nozzle are in the range from melting temperature to approximately half way between melting and vaporization temperature of the respective material. When impacting the substrate, most droplets have significantly cooled down and solidify within a few microseconds. The heat stored in the droplets and the energy available for bonding to the substrate can be controlled by the power level for melting (arc power or combustion parameters), by regu-
46
lating the feed material (feed rate) and the gas velocity for propelling the droplets. The kinetic energy of the droplets varies between the different spray processes and can be quite significant. Droplet velocities vary from submach (for arc and conventional plasma spraying) to supersonic (for combustion and vacuum plasma spraying). When the droplets impact the target surface they are deformed and flattened. Metallographic investigations of sprayed coatings [36] show, that the droplets solidify in isolation, i.e. there is no liquid pool on the surface. Upon impact the droplets flatten out to form lamellae with a thickness of approximately 5 to 10% of their original diameter. The material density of sprayed materials is directly related to the droplet velocity. Arc or plasma sprayed deposits typically achieve porosities varying form 1 to 10% [36], depending on the atomization gas and shrouding technique. The high velocity processes are able to produce dense materials with evidence of fusion bonding. While this would be desirable for use in the SDM process, the particles also tend to penetrate the target. This makes high velocity spray processes highly abrasive, and because they can destroy the support structure needed for manufacturing parts, they are unsuitable in an SDM approach. The low velocity processes are able to spray materials with high melting points like steel (Tmelt ~ 1600°C) onto tin/zinc support structures (Tmelt ~ 200°C) by superficially melting the surface and thus adhering to it. However, power levels have to be on the low end, and due to the relatively low velocities and heat content, the droplets do not have enough energy to remelt the underlying material. The nature of interparticle bonding in these processes seems to be predominately mechanical, i.e. the particles interlock in a jig-saw like matrix. The bonding strength of low-power sprayed material is on the order of 10% compared with conventionally manufactured material1. Furthermore, machined surfaces roughened through grit-blasting operations do not achieve the same surface roughness as an as sprayed surface, thus further decreasing the bond-strength between the individual layers of a part. Another inherent problem in spraying processes is the entrainment of gases from the surrounding atmosphere. Spraying in air, for example, entrains oxygen which causes oxidation of the hot droplets. Relatively small amounts of oxidation (in the order of less than 1%) can significantly reduce the intermaterial bonding strength of spray-formed materials. The relative amounts of oxidation can be rather large. Arc spraying in air with air atomization causes oxidation values up to 30% for steels, which can be reduced to about 5% with inert gas atomization and shrouding, and to about 2% in an inert chamber [36]. High velocity processes have less oxidation (typical on the order of 0.1%) because of limited transit times. Spraying in a controlled atmospheric chamber is a very costly alternative, shrouding solutions, where the particle stream is shielded from the surrounding air, require significant volumes of inert gas because of high gas flows from the spray torch itself. Looking at the properties of the different spray processes it is apparent, that none of them can satisfy all conditions necessary for manufacturing parts in the SDM process.
4.1.2. Conventional Welding In comparison with thermal spraying technologies, conventional welding processes produce a significant heat flux into the substrate. In typical arc welding applications [38], an electric arc is 1. Comparison between tensile test data of low-power arc sprayed zinc with rolled zinc sheets.
47
established between the welding torch and the substrate. Among the available welding processes two seem to have potential for use in the SDM process. Gas metal-arc welding, also called metal inert-gas (MIG) welding, uses a continuous, solid electrode wire, which serves as a source of weld metal and as the electrode for the electric arc. Gas tungsten-arc welding, or tungsten inert-gas (TIG) welding, uses a non-consumable tungsten electrode to establish the arc to the substrate. The weld material is fed externally into the arc. For both processes the energy of the impacting arc remelts part of the substrate to create a liquid weld puddle, into which the weld material is fed (Figure 4.1). The electrode tip (welding torch), the weld puddle and the adjacent areas of the workpiece are protected by shielding gas which is supplied through the torch nozzle. As the puddle resolidifies it establishes a metallurgical bond between the substrate and the weld material. The electric arc also breaks up any residual oxide layers on the underlying surfaces. While remelting the substrate results in excellent metallurgical bonds and shielding the weld puddle guarantees oxide free material, the excessive heat from the arc in combination with slow cooling rates creates relatively large heat affected zones with disrupted microstructure. Welding torch Weld metal
Electric arc
Feed wire Shield gas Substrate
depth of remelting
Weld puddle Heat-affected zone
Figure 4.1: MIG Welding Process
A welding technique derived from TIG welding is plasma-arc welding (PAW). It is characterized by a constricted arc with higher plasma temperatures and velocities. PAW has been successfully used to weld metals as thin as 0.05 mm, but deformation of the base material, and deeper penetration which is required for thicker materials, limits the use for SDM purposes. In a process called “Shape Melting” [24] PAW was used to incrementally build engineering parts. The shapes created with this process have been limited in size, complexity and geometry since suitable sacrificial support materials, that can tolerate the excessive heat input, have yet to be identified. This approach also required a customized, geometry dependent cooling apparatus. Typical parts built with shape melting include large, symmetrically shaped components such as pipes, elbows and flanged cylinders. Excessive energy transfer into underlying material by direct transfer of the welding arc, causes penetration of the weld puddle and disrupts the geometry and microstructure of previous deposits. Mixing (alloying) and therefore destroying the geometry when welding over support material, and the lack of a support structure which is conductive, resistant to the harsh conditions of the transferred arc, that can be shaped with the necessary accuracy, and which also can be removed after the part is completed, are additional problems, which make conventional welding processes unsuitable for deposition in SDM.
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4.2. The Principle of Microcasting It is apparent that neither conventional thermal spraying nor conventional arc welding alone exhibit all the capabilities required to achieve the deposition properties required in the SDM process. An alternative, economical deposition process is needed that allows deposits with metallurgical bond to the underlying material and that can create fully dense deposits without oxidation. To keep penetration to a minimum (µm range), which is required for the metallurgical bond, the energy stored in the deposited material and the energy transfer into the base material must be controllable. Together with fast cooling rates of the deposit this will ensure a uniform microstructure of the deposit and will effect the geometry of previous layers and the sacrificial support material in the range of a few microns, which can be tolerated for use in the SDM process. An attempt to achieve these goals by combining the benefits of thermal spraying with the benefits of conventional welding is the development of a process called microcasting. This novel approach creates molten droplets with sizes in the range of approximately 1 to 4 mm. The droplets are deposited in a discrete fashion with a droplet rate of approximately 1 to 3 droplets per second. The impact speed of the droplets on the substrate are relatively low, and in general in the order of 1 m/ s. The droplet temperatures at the impact can be varied from melting temperature to almost vaporization temperature of the material. The droplets are generated without effecting the substrate material, i.e. no arc is transferred into the substrate. The energy transfer into the underlying material is therefore mainly controlled by the droplet size, rate and temperature. If necessary, an independent power source (induction heater, electric arc) can be used to locally preheat the target material. On the other hand, if cooling rates need to be tightly controlled to achieve materials with certain microstructures, controlled heating or cooling of the deposited droplet will be possible. Typical natural cooling times for the individual droplets being deposited on metal substrates are in the range of 1 to 3 seconds, which is mainly dependent on the heat transfer of the droplet into the substrate material. The discrete nature of the droplet deposition enables selective heating, e.g with flames or electric arcs, or selective cooling, e.g. with bursts of CO2 or liquid nitrogen, to create the correct temperature-time history for each individual droplet without disturbing the overall deposition process. To ensure oxide free deposits microcasting is typically performed in an inert chamber or with a shrouded microcasting device.
4.2.1. Comparison of Substrate Heat Transfer The goal of the microcasting process is to achieve superior bonding conditions, which locally require increased heat flux to achieve remelting, while maintaining the low, overall heat input of spray processes. To show, that microcasting is able to produce significantly higher energy flow per coated surface area, and still maintains a low total energy flow, typical scenarios are compared for microcasting, welding and spraying. the results are summarized in Table 4.1. Using parameter set 6 and the measurements for mild steel droplets from Table 4.2 the following would be typical properties for microcasting. At a plasma current of 68 A and a voltage of 29.2 V 1.2 drops where deposited per second with a mass of 0.1912 g/drop and a droplet diameter of 3.6 mm. The average thermal energy stored in each droplet was measured to be 392.7 J/drop, which
49
results in an average droplet temperature of 2596 K (see Table 4.5). Assuming that radiation can be neglected [37], the total energy flux into the substrate results to 471.3 W, for a deposition rate of 826 g/h. Assuming further, that the microcast droplets flatten to a height of approximately half their original diameter and keep roughly the shape of half a sphere (which is roughly consistent with experimental observations), the diameter of that sphere is 4.53 mm and the area covered by the droplet is 16.12 mm2. This results in a density of the thermal energy of 24.37 J/mm2. For a welding scenario, the same current and voltage can be assumed for the arc, as well as the same amount of deposited material and coverage, because of the similarity with the microcasting process. The heat transfer efficiency, i.e the amount of arc power that is transferred into the substrate is between 0.7 and 1.0 for most arc welding processes [38]. Assuming a efficiency of 0.75 the total power transferred into the substrate results to 1489 W. This is 3.16 times the amount that is transferred by microcasting the same deposit. In terms of heat transfer efficiency, microcasting would have an efficiency of approximately 0.25. For spraying, investigations in [36] have shown, that the droplets are not superheated upon their impact with the substrate. Assuming fully molten droplets at melting temperature the graph in Figure 4.12 b) shows an energy of 1427 J/g. Typical deposition rates for arc and plasma spray deposition are in the area of 1 kg/h. To compare the results for equal deposition rates, 826 g/h are used for this example. The total energy flux into the substrate is therefore 327.4 W. Assuming rather large droplets for spraying with a diameter of 100 µm, each droplet has a mass of 4.12 µg. This results in a thermal energy of 5.87 mJ/drop. Observation of sprayed microstructures in [36] and simulations of typical spray droplets in [39] show, that for the impacting droplets a flattening degree of 3, i.e. the diameter of the flattened droplet is 3 times the original droplet diameter, can be assumed. This leads to a covered substrate area of 7.07·10-2 mm2 for each droplet, and an energy density of 83 mJ/mm2.
Process
total energy flux [W]
thermal energy per drop [J]
energy per covered area [J/mm2]
Arc-Welding
1489
-
-
Microcasting
471.3
392.7
24.37
Arc-, Plasmaspraying
327.4
5.87·10-3
8.31·10-2
Table 4.1: Comparison of Energy Content and Flux In comparison to arc or plasma spraying, microcasting makes more efficient use of the thermal energy to create remelting conditions. While the total thermal energy of spraying is only 31% less than the one for microcasting, the available thermal energy per covered area is 293 times greater for microcasting droplets than for spraying droplets.
50
4.3. The Microcasting Apparatus For low melting materials, such as waxes, solders or tin, mechanisms have been developed to selectively deposit uniform, discrete droplets. They normally use a heated supply of liquid material. One setup described in [40] uses a principle similar to the inkjet printing head to squirt picoliter droplets from a tiny tube by squeezing it with a piezo-electric transducer. Other approaches break up the stream from a tiny nozzle by charging the droplets and pulling them from the nozzle by imposing regular vibrations through the charging mechanism [41]. Processes for metals with high melting points, that use a supply of liquid metal are typically not very controllable. Also, storing and using a liquid supply for metals with high melting points is very difficult and costly. Containers must be made from or lined out with refractory material and need constant repair. Ceramic nozzles show excessive wear and need to be constantly replaced. Controlling deposition in form of discrete droplets is almost impossible, since nozzle openings and valve mechanisms typically “freeze” for small flow rates. Valves used for liquid, high-temperature metals are normally ceramic rods with one conically shaped end, which is inserted into a ceramic opening or tube. Accurate flow is very difficult to control with such a setup, which is further complicated by excessive wear of the ceramic parts, which are submerged in the liquid metal. Because of the problems with storing and dispensing liquid, high-temperature metals, a different approach has been chosen for the microcasting process. The microcasting apparatus consists mainly of conventional welding equipment, which is configured in a novel, non-transferred arrangement, i.e. the electric arc is not transferred to the substrate.
4.3.1. Early Versions of the Microcasting Device Early experiments with microcasting were conducted with MIG and TIG welding torches. Instead of transferring the welding arc to the substrate, the arc was transferred to a water-cooled copper block or a rotating, water-cooled copper cylinder [42]. Figure 4.2 a) depicts a setup using a MIG welding torch. The MIG torch is pointed under an angel of 45° at the side of the water-cooled copper electrode (see also Figure A.8 in the Appendix). As the electrode wire is fed through the MIG torch, the arc melts the wire, and a droplet forms on the end of the wire. When the droplet is big enough to overcome the surface tension, it free-falls to the substrate. Shielding gas is supplied through the MIG torch. Figure 4.2 b) shows a similar setup with a TIG welding torch. In this case the arc is established between a non-consumable tungsten electrode and the water-cooled copper electrode. The uncharged wire is fed externally into the arc. By traversing the microcasting setup relative to the substrate, lines of metal can be deposited. By traversing these setups relative to the substrate it was possible to produce layers of deposited metal. However, there were several problems with the early versions of the microcasting devices. The impact of the shielding gas stream and the electric arc on the molten droplets caused the droplets in some cases to be blown onto the cooled electrode, were they collected because of a bigger surface for adhesion and then fell to the target in form of a bigger drop. In case of the rotating electrode, some droplets were dragged along the rotation more than others, some droplets even
51
Rotating water-cooled copper electrode
Power supply
Power supply
-
+
Feed wire
Feed wire
+
TIG torch
MIG torch
water-cooled copper block
Shield gas Substrate
Electric arc Droplet
Tungsten electrode Electric arc Droplet Shield gas
Substrate
a) MIG Microcaster
b) TIG Microcaster
Figure 4.2: MIG and TIG Microcaster with water-cooled Copper Electrode got stuck to the electrode. Other problems included carbon buildup on the copper electrode, wear of the copper electrode due to the arc transfer and adhering liquid droplets, and condensation due to the cooling. As a result, the transferred arc was not always stable, and the droplet trajectories strayed quite significantly (up to 5mm at the target). Due to the rapid solidification of the droplets on the target, droplets miss-hitting to the side are not able to flow out, droplets miss-hitting in the traverse direction fall on top of a previously deposited, solidified drop and cannot slide off to the substrate, thus causing voids in the deposit (see Figure 4.8). The TIG setup, in general, proved to be more unstable than the MIG version. Another problem with the MIG and TIG microcaster was an insufficient droplet temperature range because of extreme wear of the copper electrode due to high plasma currents and increased instability. Droplets with insufficient temperature, i.e. low energy content, show undercuts and do not completely remelt the surface of the substrate. As a result, voids are created at the seam of deposited lines in the lower sections of each layer and the deposit delaminates from the substrate.
4.3.2. The Plasma-Arc Microcaster To improve the performance of the microcasting apparatus a different setup is chosen. To create the plasma, a plasma-arc welding torch is used. Plasma-arc welding (PAW) is in principle similar to TIG welding (Figure 4.3). It uses a non-consumable tungsten electrode and two separate streams of gas, one that serves as the plasma gas for the arc, the other is used for shielding. The plasma gas surrounds the electrode and passes through an orifice, which constricts the arc to a narrow shape and forms a fast jet of hot, ionized gas. The shield gas is supplied between the outside of the orifice body of the torch and an outer shield cup. The plasma velocities and arc temperatures (typical 16,700°C) are higher than for TIG or MIG welding. 4.3.2.1. Configuration The principle of the setup for the plasma-arc microcaster is shown in Figure 4.4 The plasma welding torch is positioned horizontally above the substrate. The metal wire is positioned perpendicu-
52
Tungsten electrode
Outer shield cup
Plasma gas Shielding gas Constricted arc
Orifice
Substrate
Figure 4.3: Plasma-Arc Welding Torch lar to the plasma welding torch and fed towards the substrate. The electric arc is directly transferred between the plasma torch and the feedstock wire. The power supply used for creating the arc is a standard plasma welding power source. The negative terminal of the power source is connected to the plasma torch, the positive terminal is connected to the feed-wire through a welding contact tip. Plasma gas and primary shielding gas (both typically argon) are supplied from the plasma torch, no gas is being supplied from the direction of the wire. When the electric arc is established, and the wire is fed, the tip of the wire is melted and the liquid metal accumulates to form a liquid droplet hanging from the end of the wire. Once the droplet reaches enough mass, so the gravitational force overcomes the surface tension holding the droplet to the wire, the droplet free-falls to the substrate, and a new droplet starts forming. Power supply
+
Feed wire
-
Electric arc
Shield gas Droplet
PAW torch Substrate
Figure 4.4: Plasma-Arc Microcaster The described configuration of the plasma-arc microcaster is built using commercially available welding equipment (see Figure A.9 and A.10 in the Appendix). The plasma torch is a ‘Thermal Dynamics’ PWH/M-4A welding torch with a maximum current rating of 200 A. For controlling the plasma and shield gases and to initiate the arc a ‘Thermal Dynamics’ WC100B plasma welding console is used. The arc power is provided by a ‘Thermal Dynamics’ PS30A plasma welding power supply. To guide the metal wire and to connect it to the positive terminal of the power supply, a ‘M.K. King Cobra’ MIG welding torch is used. The gas cup, which is normally provided
53
with this torch, has been removed. The wire feed is controlled by a ‘M.K. Cobramatic II’ wire feeder, which is connected to the MIG torch and can provide feedrates up to 0.42 m/s. Both the plasma torch and the MIG torch are water-cooled with separate coolant recirculators. The plasma and shield gas is supplied to the plasma console from bottles through a ‘Victor’ 250C two stage regulator. An aluminum bracket is used to hold the two torches at the correct distance and orientation from each other. The bracket is mounted in a tool-changing mechanism, and can be picked up by the deposition robot. The power supply, the plasma console and the wire feeder are controlled by a custom interface, which enables manual or robotic switching of gasflow, arc power and wire feed, as well as control of the settings for current and wire feedrate. 4.3.2.2. Typical Microcaster Adjustments To generate droplets with different characteristics, or to create droplets from different materials, typically only two parameters are changed: The plasma current and the wire feedrate. Other parameters, such as gas flowrates, certain distances and adjustments on the plasma torch are left unchanged. The following describes those settings, as they were used for all experiments with the plasma-arc microcaster. Two settings on the plasma torch influence the shape and velocity of the plasma column. By moving the electrode, the electrode setback, i.e. the distance of the electrode tip to the front face of the torch tip, can be modified. The size of the orifice is changed by replacing the tip of the torch. The electrode setback is adjusted to 2.85 mm (i.e. gauge setting 3), and a 3.175 mm (0.125 in) diameter orifice is used.
MIG torch electrode tip Orifice
Tungsten Electrode
∆z droplet offset Wire tip distance
Orifice diameter
PAW torch
Wire
Droplet
Standoff distance
Plasma tip distance
Electrode setback Deposited Droplet
a) PAW Orifice and Electrode
∆x droplet offset
Substrate
b) Microcaster Distances
Figure 4.5: Electrode Gauging and Microcaster Distances The two torches in the microcaster are aligned, so that the axis of the tungsten electrode of the plasma torch and the axis of the wire intersect. The distance of the tip of the plasma torch to the intersection point characterizes the length of the arc. This distance has to stay below a certain
54
limit to ensure reliable starting conditions for the arc and to prevent breakup of the arc due to excessive voltages, a certain minimum is required to prevent spatter from liquid droplets into the plasma torch. The distance from the tip of the MIG torch, which charges and guides the wire, should be kept low, to limit deviations of the position of the wire due to wire curvature and curl, but has be large enough to prevent a transfer of the arc to the exposed copper tip. Depending on the operating conditions (wire feedrate and plasma current) the droplet forms at a different distance above the axis of the tungsten electrode. Typically, the plasma tip distance is adjusted to 13 mm, the wire tip distance to 17 mm. To increase the stability of the position of the wire, smaller tip openings than recommended for usual welding operations were used with the MIG torch. For 0.9 mm (0.035 in) diameter wire, tips with a 1.0 mm (0.040 in) opening (‘MK 040’) are used. For the plasma and shield gas, welding grade argon is supplied at a pressure of 250 kPa to the plasma console. The flowrates are set to 1.15 l/min (liters per minute) for the plasma gas, and 11.5 l/min for the shield gas. 4.3.2.3. Droplet Formation The mechanism for droplet generation is a balance between the surface tension, which holds the liquid droplet on the wire, gravity, and vibration and impact forces from the plasma stream. As the wire gets fed into the arc it melts and the liquid metal accumulates at the end of the wire. As the liquid droplet grows, it gains weight, and as soon as the gravitational forces and the forces generated by the plasma drag and vibration overcome the surface tension, the droplet falls to the substrate (see Figure A.11 in the Appendix). Factors influencing the size of each droplet are the temperature of the droplet, which effects the surface tension, the plasma current and gas velocities, the location where the droplet melts in the plasma (closer to the edges or the center) and the diameter of the wire, i.e. the surface area to which the droplet adheres. As the temperature of the droplets is increased, the viscosity, and therefore the surface tension of the droplets decreases. Less droplet mass and drag or vibration forces are necessary to pull the droplets off the wire. As a result, with increasing plasma currents and temperature the size of the droplets decreases and the rate, at which the droplets are generated increases (constant mass flow). The droplet temperature can be regulated through the plasma current and wire feedrate. For constant wire feed rate, increasing the plasma current causes an increase in droplet temperature. For constant plasma current, a decrease in the wire feedrate results in increased droplet temperature. Changing these operating conditions also results in a change of the location were the wire melts with respect to the axis of the tungsten electrode. For low plasma currents or high feedrates, i.e colder droplets, the wire melts below the axis of the tungsten electrode Figure 4.6 a). If the lower limits for the current and the upper limits for the feedrate are reached, the wire is not completely melted, and leaves the arc at its lower end still in a wire shape. For increasing plasma current and/ or decreasing wire feedrates, the position at which the wire melts and the liquid droplet accumulates starts moving up along the axis of the wire. For high currents and low feed rates, the droplet seems to be “floating” on top of the arc Figure 4.6 b). In this area, increasing the current or decreasing the feed rate produces little change. When reaching the upper limits for the plasma current and the lower limits for the wire feed rate, the arc starts to break up periodically, and
55
might transfer to the copper tip of the MIG welding torch.
b) High current Low wire feed
a) Low current High wire feed
Figure 4.6: Droplet Melting and Trajectories The standoff distance, i.e the distance between the intersection of the wire-axis and the axis of the tungsten electrode from the plasma torch and the substrate surface, plays an important role in the amount of flattening of the droplet upon impact on the substrate. Droplets with too little velocity upon impact cannot flatten out and stay in a shape, which is close to the original shape of the droplet. This causes undercuts in the lower portion of the deposit which leads to voids in the lower half of a deposited layer. Droplets with too much velocity upon impact break up into many little splats, which get scattered on the substrate and do not have enough energy to reliably remelt the substrate. The random deposition is further a cause of voids in the deposit. To achieve void-free deposits, with sufficient energy to remelt the underlying material, flattening of the droplets to a height of half the diameter is desired. With typical droplet temperatures of approximately 2500 K (for steel) the optimal standoff distance is approximately 7.5 cm. At distances above 12.5 cm droplets started breaking up in the experiment, at distances closer than 5 cm the forming of droplet undercuts, which lead to voids, was observed (see Figure 4.8). Typical droplets for low-carbon steel, stainless steel or copper have temperatures ranging from the melting point to approximately a few hundred degrees below the boiling point. The range for typical droplet diameters is between 3 and 4 mm. While the deposition frequency can be regulated from practically zero to a continuous stream, useful values are between 1 and 4 droplets per second. 4.3.2.4. Droplet Trajectories The trajectory of the droplet, which is primarily caused by gravity pulling the droplet off the wire, is effected by the impact of the plasma. A horizontal velocity component is added to the trajectory of the droplet, which is dependent on the current and feedrate settings. The effect of the impact depends on where the droplet accumulates with respect to the axis of the tungsten electrode, i.e., at what angle the plasma hits the droplet. Near the axis, the effect is strongest and declines below or above the axis. Above the axis, however, increased plasma currents add to the strength of the impact. For the copper droplets produced with parameter set no. 4 from Table 4.4, the offset from the wire axis has been measured at different standoff distances. The trajectory of the droplets approxi-
56
mately follows a parabolic path and can be estimated for any standoff distance. The offset ∆x from the wire axis is approximately proportional to the square root of the standoff distance d. The factor k has to be calculated for each individual parameter setting. In this example k=1.9796 was calculated from the measurement points using a least squares method. ∆x = k d
(4.1)
0
k=1.9796
-10 -20 -30
d [mm]
-40 -50 -60 -70 -80 -90 100
∆x [mm]
Figure 4.7: Droplet Offset vs. Standoff Distance for Copper (for set4)
Figure 4.7 shows the calculated trajectory of the copper droplets and the actual measurement points (marked ‘x’). For other materials and standoff distances, the droplet offset can be calculated from measurements at a given standoff. In Chapter 4.4 droplet offsets have been measured for different materials and parameter settings. Repeatable and stable trajectories are important for the creation of void-free deposits, which rely on the accurate positioning of each droplet. Several factors influence the ability to create repeatable and stable droplet trajectories. Feeding the wire vertically from the top ensures the smallest possible area to which the droplet can adhere. Experiments with horizontally fed wire have shown, that due to the increased area available for adhesion and due to the tendency of the droplet to slightly move away from the tip of the wire, the droplet trajectories became more instable, i.e. the point at which the droplet hits the substrate at a given standoff distance deviated from droplet to droplet. The position of the wire in the plasma stream has to be adjusted, so that the wire is positioned exactly in the center of the stream, i.e. the axis of the wire has to intersect with the axis of the tungsten electrode. Small deviations from the center cause strong vibrations of the droplets, which are then thrown off the wire in unpredictable trajectories. With respect to this, it is very important, that the wire is fed straight (without any curvature) and without curl, which is sometimes typical for ordinary welding electrodes. The adjustment of the plasma torch (electrode setback and orifice opening) influence the size and velocity of the plasma column, and therefore have an influence on the trajectory of the droplet. Due to wear of the electrode and the orifice, and possible buildup of tiny spatter on the tip of the torch, the electrode and torch tip need frequent cleaning and readjustment to ensure repeatable droplet trajectories, which are necessary for void-free deposits.
57
4.3.2.5. Deposition Voids Layers of microcast material are deposited by moving the microcaster relative to the substrate, so that subsequent droplets are deposited at a certain distance d from each other. Due to the rapid solidification of the droplets on the substrate, the droplets do not tend to flow and accurate positioning of the droplets is necessary. The speed, at which the microcaster is moved relative to the substrate, the path spacing and the standoff distance are crucial for the creation of void-free deposits. Several different mechanisms can cause the generation of voids in the deposit. Ideally, as shown in Figure 4.8 a), the droplets are deposited at a distance d from each other, so they can slide off the previous droplet and adhere to the substrate, and the previous droplet. Figure 4.8 b) shows droplets deposited with incorrect spacing. Droplets with a spacing greater than the ideal distance, do not adhere to each other, and create gaps, which form voids in the upper regions of the deposited layer. Droplets being deposited too close too each other, will not slide off the previous droplet, due to the rapid solidification, and adhere to the previous droplet forming an overhang. With the next droplet sliding to the substrate, and not being able to flow into the gap, voids are created at the lower portion of the deposit. The correct distance d between subsequent droplets depends on the diameter and temperature of the droplet and the standoff distance of the microcaster (due to droplet flattening). The path spacing (for a serpentine path) and the speed, at which the microcaster is moved relative to the substrate have to be set accordingly (speed = d / drop rate). Bad wetting conditions, combined with insufficient heat transfer into the substrate can have a similar effect as droplets with incorrect spacing. If a droplet cannot cool fast enough, and the next droplet hits adjacent to it, the liquid cores of the droplets are combined, and the surface tension tries to create one bigger sized droplet. With the first droplet already adhering slightly to the substrate, the material of the second droplet gets moved towards the first droplet. This causes an incorrect spacing for the droplet following the second one, and ultimately voids are created.
Void d
a) Correct droplet spacing
Void d
Void Undercut
c) Droplets with undercuts
b) Incorrect droplet spacing
d) Droplet Miss-hit (top view)
Figure 4.8: Deposition Voids Another source for voids in a deposited layer are droplets, which show undercuts at the lower portion of their cross-section Figure 4.8 c). Low droplet temperature, not enough standoff distance and poor wetting conditions due to substrate oxide layers can cause such droplet shapes. As a result, voids are formed between the individual droplets on the lower portion of the deposited layer. Finally, miss-hitting droplets, due to not fully repeatable trajectories, create voids as shown in Figure 4.8 d). Variations in the droplet trajectories can be caused by turbulences in the plasma
58
gas stream of the microcaster, instabilities in the wire feedrates, miss-adjustment of the microcaster or wear of the microcaster electrode and orifice. The voids are generated on the one hand by missing material and on the other hand by other droplets adhering to the stray droplet and forming overhangs. The creation of layers from microcast droplets with serpentine trajectories leaves a deposit with a wavy top surface. To prevent the creation of voids between the individual layers of a multilayer structure, the top portion of the deposit is removed, to provide a flat, oxide-free surface for the next layer. Structures with macroscopic voids show a significant decrease in material properties, especially strength. While a variety of measures can be taken to prevent the creation of voids in microcast deposits, the potential presence of voids cannot be completely eliminated. Visual inspection of each deposit is therefore necessary, to ensure proper material conditions. Figure 4.9 shows a digitized black and white image of one layer of an injection molding tool made from mild steel for an automotive connector. Typically, the inspection would be done immediately after the microcast surface was planed and before the contours of layer are shaped. In this particular case, the layer was inspected after the contours were shaped. The cavities on the inside were filled with a support material made from ceramic powder, which was bound with sodium silicate, to enhance the contrast. Several voids are visible on the top surface with very good contrast between the machined surface of the layer, and the rough surface of the voids. The inspection process can be easily automated, by thresholding the digitized picture and creating an image, which shows problem areas as black pixels. Larger dark areas can be identified as voids and be compared to the original crosssection of the layer, which is available from the CAD model. Process instructions can then be calculated to drill out the area around the voids, if necessary, and locally redeposit some material.
Voids Voids
Figure 4.9: Macroscopic Voids in a Layer of Mild Steel
59
4.3.2.6. Shrouding While the droplets start forming on the wire, they are protected from the surrounding air by the shield gas from the plasma torch. Shortly after they fall off the wire, they leave the inert gas stream from the plasma torch, and during their fall to the surface of the substrate they are exposed to the air environment for approximately one tenth of a second. Considering the temperature of the droplet this is enough to cause significant oxidation of the droplet surface, resulting in bad wetting conditions, ultimately causing bad bonding to the substrate. Oxidation of the droplets also produces deposits with slight undercuts, which result in voids between deposited lines on the lower portion of each layer. Cutting of oxidized materials is almost impossible, due to the hardness and its abrasive nature. In general, oxidation of the droplets does not produce desired material properties and has to be avoided. Inert chambers are very expensive in their use and frequent evacuation and backfilling with inert gas makes them uneconomic for a layered deposition process. Another rather inexpensive technique is the use of adequate shrouding, with a sufficiently large flow of inert gas, which displaces the air in the vicinity of the trajectory of the droplet. Special consideration has to be given to provide a laminar flow of inert gas, that does not disturb the arc or the stability of the droplet. A recently developed, still proprietary shrouding technique1 was successfully used to protect the droplets from oxidation. Modifications to the initially cylindrical shape led to a shape with beveled sides. Openings in the sides, to allow for bringing in the plasma torch, to prevent the plasma from heating the opposite side, and to allow for the welding fumes to escape without being deposited onto the surface of the substrate ultimately led to the use of two “half-cone” shaped diffusers, which are positioned to both sides of the microcaster. Figure A.10 in the Appendix shows a microcaster with a mounted shroud.
Porous metal
Shroud
Shroud gas supply
Shroud gas
Figure 4.10: Shrouded Microcaster Inert gas supplied to the diffusers flows through the porous metal on the inside and forms a laminar curtain along the microcaster. The inert atmosphere can effectively provide shielding to approximately one diffuser-diameter from the top and bottom of the device. In the experimental setup nitrogen was used for shielding. Typically, a flowrate of 7.87 l/s (liters per second) was 1. developed by Praxair, Inc.; U. S. Patent 4,823,680.
60
used. With a standoff distance of 7.5 cm the oxygen level at the substrate was measured to be below 20 ppm. This is a value that typically can be reached with a vacuum created by rotary pumps. Normally nitrogen is supplied in liquid form. To generate the typical flowrate of 7.87 l/s of gaseous nitrogen, one liter of liquid nitrogen is required per minute. 4.3.2.7. Typical Microcasting Operation Sequence A typical microcasting operation follows a specified sequence of events, which is described in the following. To initiate the plasma arc between the plasma torch and the wire, a pilot arc is used. The pilot arc is initiated between the tungsten electrode and the tip of the plasma torch through a high voltage discharge. The plasma is blown through the orifice, thus creating an ionized path to the wire. The typical operation of the microcaster starts by purging the plasma and shield gas lines of the plasma torch. The pilot arc can then be established, and the gas through the shroud is turned on. For each segment of the path, i.e. each individual “island” of the desired deposit, the microcaster is positioned above the starting point of the path. After waiting for a few seconds to displace the air through the inert shroud gas, the main arc power and the wire feed are initiated simultaneously. After a dwell time (typically 1s) to wait for the first droplet being deposited, the microcaster is moved along the specified path, with a material and droplet dependent speed, standoff and distance between serpentine lines of the path. At the end of a path segment, another dwell time (typically 1s) is added to enable the last droplet to fall off, before the main arc and the wire feed are switched off simultaneously, while shrouding gas is still being supplied for a specified amount of time. With the pilot arc still being initiated, the microcaster can be positioned for depositing the next segment of the path. After the deposit has been completed, the pilot arc and all gas flows are switched off.
4.4. Microcasting Parameters For different materials, plasma currents and wire feedrates parameters characterizing the droplets and the microcasting process have been evaluated. A variety of combinations of plasma current and wire feed rate have been set. The plasma voltage, drop rate, droplet mass, droplet diameter and the offsets were measured. All electrode wires had a diameter of 0.9 mm (0.035 in). The microcaster was adjusted according to the settings described in 4.3.2.2. The droplet offset from the wire axis was measured at a standoff distance of 7.5 cm. Parameters for microcasting operation: • standoff: distance between the top of the substrate and the intersection point of the axis of the tungsten electrode and the wire axis • speed: speed at which the microcaster is moved to create a deposit • path spacing: distance between two adjacent deposited lines • droplet offsets: ∆x: distance in the x and y direction at which the droplet hits the substrate measured from the axis of the wire. Depends on standoff, current and feedrate.
61
∆z: distance in the z-direction at which the droplet accumulated on the wire measured from the axis of the tungsten electrode. • feedrate: rate at which the wire is fed into the arc • current: main plasma arc current • pre- and postdwell: dwell time at the beginning and end of a path segment to allow for the first and last drop to be deposited • pre- and postgas: time before and after the main arc is initiated
4.4.1. ER70S-6 Mild Steel Material specifications: Hobart HB-28 composition: C 0.09%, Mn 1.55%, P 0.011%, S 0.015%, Si 0.84% Fe 97.894% tensile strength: 616.4 MPa yield point: 481.3 MPa elongation: 27% in 50.8 mm density (measured): 7.76 g/cm3 diameter: 0.9 mm (0.035 in) melting point (iron): 1809 K vaporization point (iron): 3160 K The droplet parameters for ER70S-6 mild steel are listed in Table 4.2.
4.4.2. 308 Stainless Steel Material specifications: McKay ER308 composition: C 0.04%, Mn 1.9%,Si 0.45%,Cr 20.3%, Ni 9.7% Fe 67.61% tensile strength: 579.2 MPa yield point:399.9 MPa elongation: 35% in 50 mm elasticity modulus (of wrought and annealed 308): 193 GPa density (measured): 7.90 g/cm3 diameter: 0.9 mm (0.035 in) melting point: 1683K The droplet parameters for 308 stainless steel are listed in Table 4.3.
62
Set
wire feed rate [cm/s]
plasma current [A]
arc voltage [V]
drop rate [drops/s]
mass per droplet [g/drop]
droplet diameter [mm]
∆x offset [mm]
k for eq. (4.1)
∆z offset [mm]
1
5.33
43
24.4
0.89
0.282
4.11
0
0
-3
2
5.33
51
25.1
0.86
0.310
4.24
3
0.346
-1.5
3
5.33
59
26.6
0.94
0.280
4.10
6.5
0.751
0
4
5.33
61
27.1
0.98
0.219
3.78
13
1.501
1
5
5.33
66
28.1
1.25
0.179
3.53
17
1.963
1.5
6
5.33
68
29.2
1.20
0.191
3.61
19
2.149
3
7
5.33
69
31.3
1.64
0.155
3.37
22
2.540
5
8
6.01
49
25.1
0.99
0.300
4.20
0
0
-3
9
6.01
52
25.0
1.11
0.274
4.07
5
0.577
-2
10
6.01
60
26.4
1.05
0.308
4.23
13
1.501
0
11
6.01
66
27.8
1.19
0.255
3.98
16
1.848
1
12
6.01
74
29.0
2.17
0.135
3.22
19
2.149
2
13
6.01
78
31.8
2.27
0.136
3.23
22
2.540
4
14
7.11
55
26.0
1.18
0.295
4.17
2
0.231
-3
15
7.11
59
26.0
1.28
0.284
4.12
5
0.577
-2
16
7.11
68
27.1
1.27
0.300
4.20
7
0.808
-0.5
17
7.11
75
28.6
1.64
0.225
3.82
15
1.732
1
18
7.11
82
30.4
3.03
0.115
3.05
22
2.540
2
19
7.11
86
32.6
2.86
0.113
3.04
23
2.656
3
20
7.96
60
27.0
1.19
0.300
4.20
1
0.115
-4
21
7.96
68
27.1
1.52
0.265
4.03
5
0.577
-2
22
7.96
75
28.6
1.41
0.287
4.14
10
1.155
-1
23
7.96
82
29.7
2.17
0.188
3.59
17
1.963
0
24
7.96
89
31.4
3.13
0.121
3.11
22
2.540
1
25
7.96
93
32.7
3.85
0.083
2.73
24
2.771
2
26
7.96
98
35.2
3.57
0.097
2.89
24
2.771
4
27
7.96
103
37.5
2.94
0.129
3.17
25
2.887
5
Table 4.2: Microcasting parameters for ER70S-6 mild steel
63
Set
wire feed rate [cm/s]
plasma current [A]
arc voltage [V]
drop rate [drops/s]
mass per droplet [g/drop]
droplet diameter [mm]
∆x-offset [mm]
k for eq. (4.1)
∆z-offset [mm]
1
5.84
64
26.3
2.02
0.137
3.21
15
1.732
0
2
5.84
66
28.4
2.24
0.121
3.08
18
2.078
3
3
5.84
70
30.3
1.68
0.159
3.37
19
2.194
4
4
5.84
72
32.0
1.74
0.139
3.22
14
1.617
5
5
5.84
75
33.9
2.21
0.115
3.03
14
1.617
5
6
6.86
62
25.2
1.59
0.207
3.69
2
0.231
-1
7
6.86
69
27.1
3.09
0.103
2.92
17
1.963
0
8
6.86
73
28.7
3.29
0.101
2.90
23.5
2.714
1.5
9
6.86
75
30.4
2.53
0.117
3.04
19
2.194
3
10
6.86
79
32.3
2.43
0.131
3.16
19
2.194
3.5
11
6.86
82
33.9
2.67
0.114
3.02
14
1.612
5
12
6.86
83
35.3
3.43
0.094
2.83
15
1.732
5
Table 4.3: Microcasting parameters for 308 Stainless Steel
4.4.3. Deoxidized Copper Material specifications: Natweld DEOX CU composition (of typical de-oxidized copper according to AWS A5.7 specification: Cu 98.0% min., Sn 1.0% max., Mn 0.5% max., Si 0.5% max., P 0.15% max., Al 0.01% max., Pb 0.02% max., others 0.05% max. density: 8.83 g/cm3 diameter: 0.9 mm (0.035 in) melting point (copper): 1356.45 K vaporization point (copper): 2855 K The droplet parameters for de-oxidized copper are listed in Table 4.4.
64
Set
wire feed rate [cm/s]
plasma current [A]
arc voltage [V]
drop rate [drops/s]
mass per droplet [g/drop]
droplet diameter [mm]
∆x-offset [mm]
k for eq. (4.1)
∆z-offset [mm]
1
5.84
50
22.3
1.71
0.189
3.44
5
0.577
0
2
5.84
54
23.9
2.45
0.137
3.09
10
1.155
2
3
5.84
57
24.9
2.33
0.152
3.20
16
1.848
4
4
5.84
58
26.9
2.11
0.150
3.19
17
1.963
5
5
5.84
59
28.6
2.12
0.158
3.24
18
2.078
7
6
6.86
55
22.9
1.68
0.220
3.62
4
0.462
-1
7
6.86
59
24.1
2.58
0.154
3.21
10
1.155
0
8
6.86
63
24.7
3.48
0.111
2.88
12
1.386
2
9
6.86
68
26.0
3.41
0.109
2.86
20
2.309
5
10
6.86
70
27.8
3.27
0.119
2.95
20
2.309
7
Table 4.4: Microcasting parameters for Deoxidized Copper
4.5. Droplet Temperatures The droplet temperatures have an important influence on the quality of the deposited material and on the interlayer bonding. Metallurgical bonding among the individual droplets and between the droplets and the substrate or underlying layers must be achieved. Sufficient superheating of the liquid droplets is necessary, to enable remelting of substrates or previously deposited droplets, which have temperatures well below their melting point. On the other hand, it is necessary, to keep the amount of superheat of the droplets low, to minimize the depth of remelting. It is therefore necessary, that the microcasting process is able to produce droplets with a wide range of temperatures. Different conditions for the deposition, such as underlying materials with lower or higher melting points, different geometries influencing the heat transfer, or substrates preheated to different temperatures, require the deposition of the droplets with different temperatures. To establish a relation between the microcasting parameters and the droplet temperatures for different materials, the droplet temperatures have been measured for a variety of settings. To evaluate the temperatures of microcast droplets, two methods were used. To obtain the average temperature of a series of drops for many different parameter settings, calorimetric measurements were performed. For a few selected parameter settings, the temperature and cooling rates of individual droplets were measured with high temperature thermocouples.
65
4.5.1. Calorimetric Evaluation of Droplet Temperatures Using a calorimetric method, the average temperature of a series of droplets can be measured. The setup depicted in Figure 4.11 used an insulated ceramic beaker filled with water as a calorimeter. The beaker was placed underneath the microcaster, so that the droplets would fall into the water. A cover with a small hole, to allow the droplets to pass through, was placed on top of the beaker. This prevents water from splashing out and minimizes the influence of radiation from the microcaster on the measurement results. A shutter between the calorimeter and the microcaster was used, to eliminate the possibility of different droplet characteristics during start-up and shutdown of the microcaster. The temperature of the water was measured before and after the droplets were deposited into the water with a type K thermocouple (NiCr, NiAl). Enough time was allowed to establish a uniform water temperature before the readings were taken.
Microcaster
Hot droplet (TD) Cover
Ceramic Beaker Thermocouple (Tini, Tfin)
Water (mH2O) Cold droplets (mD)
Insulation
Figure 4.11: Calorimetric Measurement of Droplet Temperatures The average temperature of the droplets can be calculated using the temperature change of the calorimeter water. The gain of enthalpy of the water ∆HH2O is approximately equal to the enthalpy loss of the microcaster droplets ∆HD. ∆H H2O = – ∆H D
(4.2)
The enthalpy gain of the water is proportional to the mass of the water mH2O, the specific heat of the water cp,H2O and the difference between the water temperature Tini before, and Tfin after the droplets have been deposited. ∆H H2O = c p,H2O ⋅ m H2O ⋅ ( T fin – T ini )
66
(4.3)
To calculate the enthalpy change of the droplets, the temperature dependence of the specific heat of the material has to be taken into account. The change in droplet enthalpy is proportional to the mass of the droplets mD multiplied by the integral over the specific heat cp,D from the final temperature of the water Tfin to the initial droplet temperature TD and the sum of the latent heat from fusion and phase transformations. TD
– ∆H D = m D ⋅
∫ cp,D ( T ) dT + ∆hfus + ∑ ∆htrans
= m D ⋅ [ h D ( T D ) – hD ( T fin ) ]
(4.4)
T
Numeric evaluation of the integral over the specific heat and adding the latent heat of fusion and transformation leads to the specific enthalpy hD(T) relative to a certain temperature (250K or 273K in the examples), which can be used to evaluate the initial temperature of the droplets. ∆H H2O h D ( T D ) = ---------------- + h D ( T fin ) mD
(4.5)
For the calorimetric evaluation of the droplet temperatures, measurements of the specific heat over a wide temperature range (room temperature to evaporation) and values for the latent heat of fusion and phase transformations are required. For unalloyed metals, these values can be found in [43] or in [44], which also covers a variety of different alloys. However, not for all materials used in the microcasting process, the exact compositions could be found. Therefore, the closest available match was used, and the results obtained from the calorimetric evaluation were compared with direct thermocouple measurements of the droplet temperatures. Close matches of the temperatures within a few percent could be obtained. This indicates, that the values from the calorimetric measurements are satisfactory for the purpose of evaluating the droplet temperatures. It also suggests, that neglecting the additional energy from a small amount of evaporated water and not correcting for the enthalpy gain of the water due to the remaining radiation from the plasma of the microcaster are satisfactory assumptions for calculating the droplet temperatures. For each experiment the calorimeter was typically filled with 350 to 430 g of water. The initial temperature of the water ranged from 283 to 300 K. The amount of droplets deposited into the water ranged between 5 and 10 g. The standoff distance between the water level and the microcaster was 7.5 cm. 4.5.1.1. ER70S-6 Mild Steel ER70S-6 is a low carbon steel, with an iron content of 97.9%. For the specific heat, the values plotted by Touloukian in [44] for iron were used for the temperature range from 250K to the melting point at 1809K. The specific heat for the liquid state of the material and the latent heat for fusion and phase transformation was used as compiled by Allard in [43]. Using a molecular weight of 55.847 g/mol the specific heat of liquid iron in the range from 1812K to 3000K is reported according to c p,Fe [ J ⁄ gK ] = 0.73218 + 2.9903 ⋅ 10
–5
⋅T
67
with
1812K ≤ T ≤ 3000K
(4.6)
The latent heat of fusion is reported as ∆hfus= 247.10 J/g at 1809K. The latent heat for the α−γ transformation is ∆hα-γ = 16.83 J/g at 1184K and for the γ-δ transformation ∆hγ−δ = 19.52 J/g at 1665K. The values for the specific heat have been combined in Figure 4.12 a). Figure 4.12 b) shows the specific enthalpy relative to 250K, including the latent heat of fusion and phase transformations. Specific Heat
Specific Enthalpy relative to 250K
.4
500
.3 .2
000
hrel. to 250K [J/g]
cp [J/gK]
.1 1 .9 .8
000
α−γ trans.
.7 .6
latent heat of fusion
500
γ−δ trans.
500
.5 .4
0
T [K]
T [K]
a)
b)
Figure 4.12: Specific Heat and Enthalpy (relative to 250K) of Iron Table 4.5 shows the measurement results and the droplet temperatures evaluated with the specific enthalpy from Figure 4.12 b) for ER70S-6 mild steel. The microcasting parameter setting correspond with Table 4.2. The relation of the droplet temperatures to the arc power for different wire feedrates is shown in Figure 4.13 a). Once sufficient power levels for completely molten droplets are reached, the temperature of the droplets rises sharply. Near the upper end of the temperature range the slope of the curves declines. This is an indication for vaporization of parts of the metal. The upper limit for the droplet temperatures is approximately 2800 K. Due to heat loss of the droplets during the flight, this is less than the vaporization point of the metal (3160 K for iron). Figure 4.13 b) shows the relation between droplet temperatures and droplet diameter for different wire feedrates. In general, the diameters decline with higher temperatures due to decreased surface tension of the liquid droplet. In addition, vibrations and the impact caused by the plasma allow the droplets to fall off the wire before the accumulated weight is able to overcome the surface tension. This is indicated by smaller droplet diameters for droplets generated with the same temperature at higher wire feedrates, which means higher plasma currents. The curve for droplets generated at a feedrate of 7.96 cm/s further suggests, that due to the impact of the plasma current, the location where the droplet is formed in the plasma (∆z-offset), plays a role for the droplet diameter. For lower temperatures the droplet forms at a ∆z-offset below zero. At the minimum of the droplet diameter, the droplet forms approximately in the “center” of the plasma, and the impact of the plasma current is strongest. If the arc power (and with it the droplet temperature) is further increased, the droplet formation rises higher into the upper regions of the plasma. Even
68
Droplet Temperature
Droplet Diameter
3000
4.5
feedrate: 5.33 cm/s 6.01 cm/s 7.11 cm/s 7.96 cm/s
2800 4
2400
dD [mm]
TD [K]
2600
2200
2000
3.5
feedrate: 5.33 cm/s 6.01 cm/s 7.11 cm/s 7.96 cm/s
1800
3
1600
2.5
P [W]
T [K]
a)
b)
Figure 4.13: Droplet Temperature and Droplet Diameter of ER70S-6
Set
arc power [W]
Tfin [K]
∆h [J/g]
TD [K]
Set
arc power [W]
Tfin [K]
∆h [J/g]
TD [K]
1
1049
288.6
1436
1845
14
1430
288.2
1232
1812
2
1280
290.2
1322
1812
15
1534
292.6
1385
1813
3
1569
291.3
1804
2310
16
1843
293.2
1504
1934
4
1653
291.8
1901
2431
17
2145
292.2
1789
2292
5
1855
292.1
1954
2497
18
2493
291.2
2025
2585
6
1986
308.6
2026
2596
19
2804
293.1
2094
2670
7
2160
291.0
2141
2727
20
1620
288.3
1233
1812
8
1230
292.0
1225
1812
21
1842
290.6
1371
1813
9
1300
289.5
1318
1813
22
2145
292.9
1446
1860
10
1584
291.5
1441
1853
23
2435
294.7
1838
2354
11
1835
292.9
1736
2263
24
2795
294.5
1900
2431
12
2146
291.7
2037
2599
25
3041
295.9
2000
2557
13
2480
290.5
2049
2615
26
3450
297.5
2108
2690
27
3863
294.5
2204
2783
Table 4.5: Droplet Temperatures of ER70S-6 Mild Steel with higher currents and temperatures the droplet diameter increases due to decreased impact and vibrations from the plasma.
69
4.5.1.2. 308 Stainless Steel The 308 grade is a stainless steel with a high chromium content (20.3%). Values for the specific heat over the required temperature range could not be found in the available literature. Touloukian [44] reports values for 304 stainless steel with nominal composition (65-71 Fe 18-20 Cr, 10-14 Ni, etc.), which is close to the composition of 308. Unfortunately only three measurement points at 530, 820 and 1080 K are listed, all of which are at 0.670 J/gK. The lack of more detailed information made an exact calorimetric evaluation of the droplet temperatures impossible. To obtain some estimated values for the droplet temperatures an attempt was made to use the specific heat (Figure 4.12) and the latent heat of fusion and phase transformation of pure iron. The specific heat and specific enthalpy used for the evaluation of the droplet temperatures are shown in Figure 4.14. for the lower temperature values, the values reported by Touloukian for 304 stainless steel were used (solid portion of the graphs). The portion of the specific heat and enthalpy used from iron are shown as dashed lines. Specific Heat
Specific Enthalpy relative to 250K
.4
2500
.3 2000 .2
hrel. to 250K [J/g]
cp [J/gK]
.1
1
.9
1500
latent heat of fusion
1000
.8 500 .7
.6
0
T [K]
T [K]
a)
b)
Figure 4.14: Estimated Specific Heat and Enthalpy (relative to 250 K) Table 4.6 shows the measurement results and the droplet temperatures calculated for 308 stainless steel according to the specific enthalpy shown in Figure 4.14 b). The microcaster parameter settings correspond with Table 4.3. Figure 4.15 a) shows the relation between the arc power and the droplet temperature for two different wire feedrates. Over the evaluated range, the temperatures are approximately linear to the arc power. The areas for lower powers, were the temperatures are approaching the melting point of the material, and the areas of high powers, where the initial temperatures of the droplets (i.e. upon falling off the feed wire) approach the vaporization point of the material could not be reliably measured because of instable droplet trajectories in those regions and are therefore not present in the graph. The droplet diameters versus the droplet temperature are plotted in Figure 4.15 b) for two differ-
70
Droplet Temperature
Droplet Diameter
000
feedrate:
.7 800
5.84 cm/s 6.86 cm/s
.6 .5
600
dD [mm]
TD [K]
.4 400
200
.3 .2 .1
feedrate:
3
5.84 cm/s 6.86 cm/s
000
.9 .8
800
P [W]
T [K]
a)
b)
Figure 4.15: Droplet Temperature and Droplet Diameter of 308 Stainless Steel
Set
arc power [W]
Tfin [K]
∆h [J/g]
TD [K]
Set
arc power [W]
Tfin [K]
∆h [J/g]
TD [K]
1
1683
308.3
1846
2306
6
1562
310.4
1629
2035
2
1874
307.4
1787
2231
7
1870
309.8
1813
2265
3
2121
307.4
2048
2555
8
2095
311.1
1760
2200
4
2304
306.2
2082
2596
9
2280
304.3
2033
2534
5
2543
305.0
2222
2768
10
2552
307.9
2029
2532
11
2780
306.5
2283
2844
12
2930
310.9
2058
2571
Table 4.6: Droplet Temperatures of 308 Stainless Steel ent wire feedrates. The diameters follow a similar characteristic as the one observed for the mild steel droplets. For the lower power settings (i.e. lower temperatures), the droplet diameters decline with increasing power. For droplets generated near the center of the plasma (∆z-offset near zero), the droplet diameters reach a local minimum due to increased vibrations and impact from the plasma current. With further increasing power levels, the droplets are formed higher in the plasma stream, and the droplet diameters increase due to a decreased influence of the vibrations and impact, even though the plasma current is increasing. When the droplets reach the upper edge of the plasma stream, the influence of vibrations and plasma impact remain almost constant, and increased plasma currents and temperatures cause declining droplet diameters. 4.5.1.3. Deoxidized Copper For copper, the specific heat and the latent heat of fusion can be found in [43] for the range from
71
298 to 1600 K. Using a molecular weight of 63.546 g/mol, the specific heat for the solid state, is reported as c p,Cu [ J ⁄ g K ] = 0.35628 + 9.8826 ⋅ 10
–5
⋅T
298K ≤ T ≤ 1357K
for
(4.7)
For the liquid state the specific heat remains constant at cp,Cu = 0.49413 J/gK from 1357 to 1600K. The latent heat of fusion at the melting point (1357K) is ∆hfus = 206.15 J/g. The specific heat and the specific enthalpy, relative to 273K, have been plotted in Figure 4.16. For temperatures higher than 1600K, continuation of the value reported for the specific heat of liquid copper has been assumed due to the lack of available measurements. Specific Heat
Specific Enthalpy relative to 273K
0.5
400
200
.48
000
hrel. to 273K [J/g]
cp [J/gK]
.46
.44
.42
800
600
latent heat of fusion
400
0.4
200
.38
0
T [K]
T [K]
a)
b)
Figure 4.16: Specific Heat and Enthalpy (relative to 273 K) of Copper Table 4.7 shows the measurements and droplet temperatures for deoxidized copper, which were obtained with the specific enthalpy shown in Figure 4.16 b). The microcasting parameters correspond with the settings in Table 4.4.
Set
arc power [W]
Tfin [K]
∆h [J/g]
TD [K]
Set
arc power [W]
Tfin [K]
∆h [J/g]
TD [K]
1
1115
304.2
1100
2231
6
1260
302.5
1108
2248
2
1291
304.0
1124
2279
7
1422
302.4
1120
2270
3
1419
297.2
1155
2338
8
1556
304.8
1125
2283
4
1560
301.7
1277
2588
9
1768
302.5
1213
2460
5
1687
300.6
1289
2611
10
1946
295.5
1306
2641
Table 4.7: Droplet Temperatures of Deoxidized Copper The droplet temperatures in Figure 4.17 a) show a similar tendency as the temperatures for mild steel. At high power settings the slope of the temperature declines with increasing power. The ini-
72
tial temperatures of the droplets reach the boiling point and part of the material vaporizes. The temperatures at the standoff distance are slightly lower, due to heat loss during the flight. At low power settings, the droplet temperatures are expected to reach the melting point of the material. The measurements taken for copper droplets deviate from this behavior, and seem to reach their lower temperature limit in the area of 2200K. A possible explanation for this behavior could be the slightly instable operating conditions, which were observed with copper at low power settings, and caused the point at which the wire was melting in the plasma to move up and down. Droplet Temperature 2700
Droplet Diameter
feedrate:
feedrate:
5.84 cm/s 6.86 cm/s
5.84 cm/s 6.86 cm/s
3.6
2600 3.4
TD [K]
dD [mm]
2500
2400
3.2
3 2300 2.8 2200 2.6
P [W]
T [K]
a)
b)
Figure 4.17: Droplet Temperature and Droplet Diameter of Copper (DEOX-CU) The droplet diameters for copper in Figure 4.17 b) show the typical decrease for lower power settings, and then remain approximately constant for each different wire feedrate for increasing power levels.
4.5.2. Thermocouple Measurements of Droplet Temperatures The temperature and cooling rates of individual droplets can be measured with high temperature thermocouples. The droplets are directly deposited onto the exposed thermocouple junction, so that the thermocouple is being embedded by the droplet. A shutter is used for the microcaster to allow only one droplet to hit the thermocouple. The experimental setup in Figure 4.18 shows two different arrangements for feeding the thermocouple to the droplet. In Figure 4.18 a) a hole with a diameter of 3.2 mm is drilled into a 12.5 mm thick steel plate (1020 steel). The two thermocouple wires are brought through a ceramic spacer, which is inserted into the hole. The thermocouple junction is formed approximately 1 mm above the ceramic spacer. In the setup shown in Figure 4.18 b) the thermocouple is brought in horizontally and kept approximately 1 mm above the surface of the steel plate. A ceramic spacer is used to insulate the thermocouple wires from each other and the steel plate. To measure the high temperatures of the droplets, uninsulated type C thermocouples (Tungsten -
73
Hot droplet (TD)
Hot droplet (TD)
Solidified droplet
Solidified droplet Thermocouple (Tm)
Ceramic spacer
Thermocouple (Tm)
Ceramic spacer
a)
b)
Figure 4.18: Thermocouple Measurement of Droplet Temperatures 5% Rhenium vs. Tungsten - 26% Rhenium) with a wire diameter of 76 µm were used. The useful temperature range for type C thermocouples is between 273 and 2593 K. The thermocouple wires are connected to a “Data Translation DT2805” data acquisition board, which is equipped with an integrated circuit to measure the terminal temperature. Software compensation is used to compensate for the temperature of the reference junction. To obtain the thermocouple voltage vTC, the temperature of the reference junction Tref is converted into the equivalent reference junction voltage vref using the conversion tables or functions of the connected thermocouple, and added to the voltage measured on the terminals. v TC = v m + v ref ( T ref )
(4.8)
The reference junction voltage and the thermocouple temperature are calculated using fifth order polynomials, which were least-squares fitted from tables for type C thermocouples in [45]. The reference junction voltage vref in mV can be obtained from the reference junction temperature Tref measured in K with v ref = – 2.5085131 + 4.0390259 ⋅ 10 – 1.4792146 ⋅ 10
–8
3
–3
⋅ T ref + 2.256059 ⋅ 10
⋅ T ref + 4.2759679 ⋅ 10
– 12
4
–5
2
⋅ T ref –
⋅ T ref + – 4.9446064 ⋅ 10
– 16
5
⋅ T ref
(4.9)
and the thermocouple temperature TTC in K results from 2
3
T TC = 271.30065 + 72.820541 ⋅ v TC – 2.8527034 ⋅ v TC + 0.16636819 ⋅ v TC – – 4.3626406 ⋅ 10
–3
4
⋅ v TC + 4.7179699 ⋅ 10
–5
5
⋅ v TC
(4.10)
The cooling rates of each individual droplet depend mainly on the heat transfer conditions to the substrate material. The surface conditions of the substrate, and the size of the interface area play an important role. Thin oxide layers on the surface result in bad heat transfer coefficients, which lead to slow cooling of the droplets and ultimately to bad adhesion and lack of remelting. Further, surface oxide layers impair the wetting conditions and droplets may bounce off. To remove the oxide layer and to achieve optimal surface contact, the substrates have been grit-blasted. Cooling rates for droplets measured with the method shown in Figure 4.18 a) do not reflect the normal conditions for the heat transfer. Due to the ceramic insulator, most of the heat transfer from the
74
droplet into the substrate is limited to a portion of the usual interface area, and the measured cooling rates will be slower. With a droplet diameter of 3 mm, which flattens to roughly 4.5 mm and the ceramic spacer with 3.2 mm diameter, half of the interface area is between the droplet and the metal substrate. The cooling rates for droplets landing entirely on a metal substrate can therefore be assumed to be faster in the order of a factor 2. The influence of the response time of the thermocouples in the measurements can be estimated by modelling the droplet temperature according to a function with exponential decay. The thermocouple can be modeled as a first order lag. In Laplace space this leads to a measured temperature Tm τD T m ( s ) = T D -----------------------------------------------( 1 + sτ D ) ( 1 + sτ TC )
(4.11)
with an actual droplet temperature TD, a droplet cooling time constant of τD and a thermocouple response time of τTC1. As a function of time, the response of a thermocouple measuring an impacting and cooling droplet is t
t
τ D – τ-----D – ------τ TC T m ( t ) = T D ( 0 ) -------------------- e – e τ D – τ TC
(4.12)
Compared to the actual droplet temperature, the measured temperature underestimates the temperature peak, and the maximum temperature measured appears delayed. Past the maximum the measured temperature follows the actual temperature closely for thermocouples with reasonably fast response times. The maximum of the measured temperature from (4.12) is delayed by tdel with respect to the impact of the droplet. – x ⋅ ln ( x ) t del = τ D ⋅ ----------------------1–x
τ TC x = ------τD
with
(4.13)
Evaluating the measured temperature at tdel leads to a factor relating the measured peak temperature to the actual droplet temperature upon impact. x
1
--------------------- T TC ( t del ) 1 - 1 – x 1–x ---------------------- = ---------–x x 1 – x TD ( 0 )
(4.14)
The cooling time of the droplet can be estimated from the decay of the measured temperature after the peak for reasonably fast thermocouples (τD > 10 τTC). With a known response time of the thermocouple, the actual droplet temperature can be calculated from the measurements. To get an estimate of the response time of the thermocouples, data from [45], interpolated with a cubic polynomial, suggest a response time of 14.5 ms for 76 µm diameter thermocouple wire in still water. For droplets cooling with a time constant of 1 s this would lead to a 63 ms delay for the 1. The cooling time and the response time are the times to reach a temperature change of 63.2%.
75
peak temperature. However, the measurements do not show this delay, and for a sample interval of 10 ms the peak temperature is reached within one sample. This suggests, that the response time for the used thermocouples, which were embedded in the liquid metal droplet, are faster than the previous estimate for still water. Using the sample rate of 10 ms as the delay time tdel, solving (4.13) leads to thermocouple response times of approximately 1.5 ms (droplet time constant assumed at 1s). The measured peak temperature would therefore be 99% of the actual droplet temperature. To verify the results of the measurements, the peak temperatures Tpeak of the thermocouple measurements are compared with the results of the calorimetric measurements Tcal. Due to very high cooling rates at the time of impact, the peak temperatures shown by the thermocouples are slightly lower than the temperatures of the droplets evaluated through calorimetry. The figures in the following paragraphs show the temperature curves which were measured for different materials. The temperature and curve expected from the values obtained by the calorimetry have been indicated in the graphs (red curves). The time constant τD of the exponential decay of the droplet temperature has been measured as the time from the peak temperature to the temperature, which is 63.2% of the initial temperature step below the peak temperature. τD,50 is the time constant, which has been measured from the point where the droplet temperature Tpeak,50 has dropped to half of the step temperature to a total temperature drop of 81.6% (i.e., 63.2% drop from the half step temperature). The droplet cooling rates have been calculated by dividing the difference of two samples by the sample interval. RD is the initial cooling rate measured at the temperature peak, RD,50 is the cooling rate at a temperature drop of 50% of the step temperature. 4.5.2.1. ER70S-6 Mild Steel For three different parameter setting for mild steel from Table 4.2 the temperature curves measured with the arrangement from Figure 4.18 a) are shown in Figure 4.19. Table 4.8 summarizes the evaluation of the droplet temperature curves. Comparing the values for curve a) and b) shows, that while the cooling time constants are within a close range, the initial cooing rate is much higher for curve a). Since the sample interval was shorter for curve a), this suggests, that the initial cooling rates experienced by the droplets are extremely high. Comparing the results of curve b) and c) shows, that the cooling time constants as well as the initial cooling rates for curve c) are approximately twice as fast than for curve b). This was most likely caused by a bigger interface area to the metal substrate.
Fig.
Par. set
Tcal [K]
Tpeak [K]
Tpeak,50 [K]
tsample [s]
τD [s]
τD,50 [s]
RD [K/s]
RD,50 [K/s]
a)
6
2596
2463
1391
0.01
0.64
0.73
15450
1468
b)
5
2497
2235
1273
0.05
0.76
0.82
> 4780
1380
c)
12
2599
2435
1363
0.05
0.29
0.30
8600
1282
Table 4.8: Evaluation of Droplet Cooling Curves for ER70S-6 Mild Steel Comparing the measured peak temperature and the temperature from the calorimetry experiments
76
2596 500
2463
expected curve
500
2497
2235 000
000
T [K]
T [K]
solidification
500
expected curve
solidification
500
000
000
500
500
t [s]
t [s]
a)
b)
2599 500
2435
expected curve
T [K]
000
500
000
500
t [s]
c)
Figure 4.19: Temperature Curves of ER70S-6 Mild Steel Droplets shows, that the temperature difference is lower than the initial cooling rate for curve a) and c) during the first sample period. In case of curve b) the temperature difference is 9.6% over the cooling rate. Due to the slower sampling time, the initial cooling rate, which drops rapidly immediately after the impact of the droplet, has been underestimated. Considering the initial cooling rate, the temperatures obtained from the thermocouple measurement match the temperatures from the calorimetric evaluation, and seem to be within 5% of the actual temperature of the droplet. 4.5.2.2. 308 Stainless Steel Using parameter set no. 3 from Table 4.3 two measurements have been taken for the same set. The resulting cooling curves are shown in Figure 4.20. The setup according to Figure 4.18 a) was used with a sampling time of 10 ms. The results for the cooling times and rates are shown in Table 4.9. The difference in the time con-
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2555 500
2424
expected curve
500
2555 expected curve 2285
000
T [K]
T [K]
000
500
500
000
000
500
500
t [s]
t [s]
a)
b)
Figure 4.20: Temperature Curves of a 308 Stainless Steel Droplet
stants between curve a) and b) are caused by different contact areas to the metal substrate, or different droplet size or flattening. Judging from the cooling time constants and the cooling rate at half the step temperature, the initial cooling rate for curve b) has been underestimated. This is a result from the peak temperature, which has been measured too low due to the fast cooling rates.
Fig.
Par. set
Tcal [K]
Tpeak [K]
Tpeak,50 [K]
tsample [s]
τD [s]
τD,50 [s]
RD [K/s]
RD,50 [K/s]
a)
3
2555
2424
1093
0.01
0.33
0.62
18300
2700
b)
3
2555
2285
1022
0.01
0.17
0.41
> 18900
5200
Table 4.9: Evaluation of Droplet Cooling Curves for 308 Stainless Steel For curve a) the difference between the thermocouple peak temperature and the temperature from the calorimetric evaluation is smaller than the amount of cooling during the initial sample interval. For curve b) the temperature difference is 42% higher than the initial cooling, which was to be expected from the underestimation of the initial cooling rate. Considering, that the initial cooling rate is probably close to twice as high as estimated, the calorimetric and thermocouple measurements deliver matching results, which are within 5% of the actual temperature of the droplets. 4.5.2.3. Deoxidized Copper The temperature measurement for the copper droplet in Figure 4.22 a) was taken with the setup depicted in Figure 4.18 b) and parameter set 4 from Table 4.4. The fast cooling rates, which were experienced made it necessary to repeat the measurement under different conditions. The setup which was used to improve upon the measurement of the peak temperature and the temperature curve is shown in Figure 4.21. To reduce the thermal conduction into the substrate, a column of copper droplets was deposited. The thermocouple was then positioned on top of the column, and a
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series of droplets was deposited onto the column. The temperature measurements taken for the first 6 droplets are shown in Figure 4.22 b). Hot droplet stream (TD)
Thermocouple (Tm)
Solidified droplets
Figure 4.21: Measurement of Droplet Stream Temperatures
2588 2500
2500
2505 2369
expected curve 2000
2000
1892
1853
1754 1444
T [K]
T [K]
solidification 1500
1500
1000
1000
500
500
t [s]
t [s]
a)
b)
1371
Figure 4.22: Temperature Curves of a Single and Multiple Copper Droplets The copper droplet in curve a), with it’s full interface area adhering to the metal substrate, experienced extremely high cooling rates, which caused the measured peak temperature to be 28% below the value of the calorimetric experiments. This also leads to a value for the initial cooling rate, which is greatly underestimated from the actual value. To evaluate the cooling time and rates for the first droplet of curve b), a steady state temperature of 900 K has been assumed from Figure 4.22 b). The measured cooling rates are much lower than for curve a). Compared with the droplet temperature from the calorimetric measurement, the peak temperature of curve b) is within 3.5%. This leads to the conclusion, that the droplet temperatures for copper have been evaluated with sufficient accuracy, an dare probably within 5% of the actual value. Figure 4.22 b) further shows, that for successive droplets, due to the increase in distance from the thermocouple, the measured peak temperature decreases. After building up several drops, the
79
Fig.
Par. set
Tcal [K]
Tpeak [K]
Tpeak,50 [K]
tsample [s]
τD [s]
τD,50 [s]
RD [K/s]
RD,50 [K/s]
a)
4
2588
1853
801
0.01
0.085
0.12
>> 19000
9800
b)
4
2588
2505
~ 1429
0.01
~ 0.10
~ 0.14
12000
4500
Table 4.10: Evaluation of Droplet Cooling Curves for Deoxidized Copper temperature will reach a steady maximum (here approximately 1350 K) and then start to decrease slowly.
4.5.3. Conclusions The calorimetric and thermocouple measurements of the droplet temperatures have shown, that the microcasting process is able to produce droplets with a wide range of temperatures, which range from the melting point to approximately 90% of the vaporization point of the materials. The temperatures for ER70S-6 mild steel droplets range from 1800 to roughly 2750 K. The droplet diameters are typically between 3 and 4 mm. The most commonly used parameter set for mild steel is set no. 6 from Table 4.2. For a wire feedrate of 5.33 cm/s and a plasma current of 68 A the arc power was measured to be 1986 W. The droplets, which are deposited at a frequency of 1.20 drops/s, weigh 0.191 g/drop and have a diameter of 3.61 mm. The droplet temperature at a standoff distance of 7.5 cm was measured to be 2596 K. For 308 stainless steel good estimates for the droplet temperatures could be obtained, despite the lack of exact values for the specific heat and latent heat of fusion and phase transformations. the temperature for the droplets ranged from 2000 to 2800 K and the droplet diameter was found between 2.9 and 3.4 mm. For building some of the sample artifacts parameter set no. 3 from Table 4.3 was used. At a wire feedrate of 5.84 cm/s and a plasma current of 70 A, the arc power reached 2121 W. The droplets have a weight of 0.159 g/drop and are deposited at a frequency of 1.68 drops/s. The droplet diameter is 3.37 mm and the temperature at a standoff distance of 7.5 cm was 2555K. The temperature range for copper droplets was found to be between 2250 and 2650K. The droplet diameters are as large as 3.4 to 3.6 mm for low power settings, and remains approximately constant for a given wire feedrate at higher settings. The values for the diameters were found to be between approximately 3.2 and 2.9 mm for the feedrates used. The parameters typically used for depositing copper is set no. 4 from Table 4.4. A wire feedrate of 5.84 cm/s and a plasma current of 58 A produce an arc power of 1560 W and droplets with a diameter of 3.19 mm and a weight of 0.150 g at a frequency of 2.11 drops/s. The droplet temperature is 2588 K.
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4.6. Microstructure of Microcast Droplets To investigate the microstructure of droplets deposited with the microcasting process, metallographic samples were prepared. One droplet of low carbon steel (ER70S-6) with a temperature of 2596 K was deposited on a room temperature substrate (1020 steel). The conditions of droplet solidification, the effects of the liquid droplet on the substrate surface and substrate remelting can be examined. Estimates of the substrate temperature and substrate and droplet cooling rates can be made using cooling transformation diagrams [46].
4.6.1. Single Microcast Droplet
500 µm
Figure 4.23: Cross-Section through a Microcast Mild Steel Droplet (20x)
Figure 4.23 shows the entire cross-section through a single droplet and the substrate composed from multiple images at a magnification of 20x. Upon impact, the droplet flattened to approximately half of its original diameter (3.6 mm). The diameter of the splat is 6 mm at a height of 1.9 mm. The droplet shows a structure of columnar martensite with the exception of the top 300 µm and a thin skin along the smooth surface of the droplet, which show an equiaxed structure. The columnar martensite is a result of the rapid quenching of the droplet. The columns indicate that the heat from the droplet was extracted through the substrate by metallic conduction from the bottom upwards. The martensitic structure indicates high solidification rates, which cool the droplet down to the temperature Mf in front of the nose of the time-transformation-temperature (TTT) diagram (Figure 4.24). According to the TTT curve for 0.2% plain carbon steel, the droplet cools to below 850 K in less than 1 s. This is in accordance with the previously measured droplet cooling curves. The equiaxed structure on the top and along the surface of the droplet is caused by heat loss through radiation and solidification from the top. Since this process is much slower than the rapid heat extraction by the substrate, the equiaxed zone does not extend very far into the droplet.
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Figure 4.24: Time-Transformation-Temperature Diagram of a 0.2% Carbon Steel The bottom of the droplet shows several voids, caused by entrapped air or shrouding gas. Increased substrate temperatures, allowing better flow of the impinging, liquid droplet, could possibly minimize or prevent the creation of voids. As can be seen from the wavy surface of the substrate and the void in the center of the interface, which extends into the substrate, remelting of the substrate occurs. A portion of the waviness of the substrate, however, is a result of the grit-blasting operation to remove the surface oxides before the droplet deposition. On the edges, delamination between the droplet and the substrate is visible. In those areas the heat input from the droplet was insufficient to remelt the substrate. Residual stresses from the cooling of the droplet tend to warp the bottom of the droplet upwards, and cracks extend into the droplet/substrate interface. For the investigation of the microstructure of the interface zone between the droplet and the substrate bigger magnification was used. Figure 4.25 shows the interface zone at a magnification of 100x. On the bottom, the unaffected substrate is visible, showing a structure of perlite (dark areas) in a ferritic matrix [47]. In the heat-affected zone, which extends between 300 µm on the left side and 600 µm on the right side of the droplet into the substrate, the coarse ferrite-perlite structure changes into a finer structure. This is an indication, that the substrate temperature reached the area between 800 and 1000 K or above, and slower cooling rates bring it into the nose of the TTT diagram (i.e., 1 s at 800 K is sufficient). The resulting structure of the heat-affected zone is some form of bainite plus perlite.
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columnar martensite
droplet droplet/substrate interface
martensite
bainite + perlite
heat affected zone
ferrite+ perlite
unaffected substrate
50 µm
Figure 4.25: Droplet/Substrate Interface Zone, Mild Steel, 100x
The upper portion of the heat-affected zone, approximately 100 µm away from the droplet/substrate interface on the left, and 250 mm on the right side of the droplet, shows a structure of martensitic cross-hatches, which indicates, that the region was heated above the austenitizing temperature (1000 K) and quickly cooled in front of the TTT nose. The exact thickness of the remelted zone is difficult to estimate. Changes in the martensitic structure towards the interface, the waviness of the interface, and voids extending into the substrate, seem to indicate a remelting depth between 30 and 75 µm.
4.6.2. Multiple Microcast Droplets To investigate the thermal history of droplets in layered deposits, two rows of droplets have been deposited on a steel substrate. A cross-section through the two droplets is shown in Figure 4.26. The top droplet shows a similar structure as the droplet in the previous case. An approximately 300 µm thick layer on the top shows an equiaxed structure, indicating solidification from the top by radiation, the columnar, martensitic structure of the remaining areas of the droplet are a sign of rapid heat extraction from the bottom. The interface zone between the two droplets has been examined at bigger magnification and shows a wavy surface with a roughness of approximately 120 µm. The surface of the top droplet, however, is smooth, therefore remelting of the first droplet can be assumed. The area underneath the interface zone shows a fine grain structure, which indicates that the originally columnar structure was affected by the heat from the second droplet. The heat affected zone reaches approximately to the middle of the droplet. In the lower portion of the droplet the fine grain structure changes back into the original columnar structure.
83
smooth top surface equiaxed zone
columnar martensite
second droplet
droplet interface
heat affected zone
first droplet
columnar martensite
unaffected zone
500 µm
Figure 4.26: Cross-Section of Two Microcast Mild Steel Droplets
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4.7. Deposition of Dissimilar Materials Joining dissimilar materials with high demands on the material quality and interface zone is a difficult problem in manufacturing technology. Conventional welding processes typically cause penetration and alloying of the materials, or are incapable of readily joining certain combinations of materials. Thermal spray technologies cannot produce thick coatings and achieve only limited material quality. Other processes, such as sputtering or plating achieve only slow deposition rates, and are also limited in the thickness and quality of the deposit. As an alternative, deposition of dissimilar materials with the microcasting process is possible. Several examples have shown the feasibility of joining different materials. Deposits with high material quality can be produced, while the geometry of the interface zone is well preserved. In conjunction with intermediate shaping operations, the creation of embedded or coated structures is possible. This opens a wide area of applications, such as coatings with considerable thickness, laminated structures or the use of different materials to optimize the properties of a structure. The deposition of dissimilar materials is used in the SDM process to create parts embedded in a support structure. Detailed investigations into the interface strength have not been conducted yet, since the support structures in the SDM process are removed after completion of the part. For the application of support structures, the main concern is remelting of substrate areas by dissimilar droplets, which is detailed in Chapter 5.4.4. While for support materials remelting of the substrate has to be kept minimal, or has to be avoided, to preserve the surface finish of the part, for joining dissimilar materials a controlled amount of remelting is desired to achieve the required interfacial strength.
Figure 4.27: Stainless Steel Layer on Mild Steel Substrate Figure 4.27 shows a 5 mm thick stainless steel (308) layer, which was deposited by microcasting onto a mild steel substrate (1020). The steel substrate has been etched to increase the contrast between the materials. Figure 4.28 shows a 1.5 mm thick copper layer, which is embedded in stainless steel. To create this sample, first a stainless steel layer was deposited through microcasting onto the mild steel substrate, and planed to achieve a flat surface. The copper was then microcast onto the flat layer
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Figure 4.28: Copper Embedded in Stainless Steel of stainless steel, planed and the sides of the copper were cut. Finally, another layer of stainless steel was deposited on top of the structure. The top surface of the last layer has not been altered, and show a structure typical for microcast deposits. Combinations, such as precise copper structures, embedded in a structure of stainless steel, thus combining the superior electrical properties of copper with the mechanical properties of (stainless) steels, will certainly have a wide range of applications. It is the intent of the author, to investigate potential applications for these combinations in the future.
4.8. Conclusions and Future Improvements The microcasting process described in this chapter has shown the ability to produce liquid metal droplets with a wide range of temperatures, spanning from the melting temperature to only a few hundred degrees below the vaporization point. Deposits made from microcast droplets have shown high material quality with metallurgical interlayer bonding, while keeping remelting of underlying material to a minimum. The deposition of dissimilar materials, with remelting and mixing limited to a superficial layer, lead to the successful creation of high quality multi-material structures, with distinct boundaries between the materials. The described setup of the microcasting process was used in the SDM process to produce a variety of parts. However, several aspects of the process can be improved. In the current setup, only the droplet temperature can be controlled through the plasma current and the wire feedrate, the droplet diameter is given by the operating conditions and cannot be adjusted independently. While the droplet temperature is crucial wether remelting of the substrate will occur or not, the droplet size has an important impact on the remelting depth through the heat content of the droplet. Smaller droplet sizes than the ones currently produced are desired. Mechanical means, such as mechanical, ultrasonic or piezoelectric vibrators, or a high-voltage electric system, with a ring electrode positioned
86
underneath the wire, and a constant or pulsed high voltage between the ring and the wire, could be used, to controllably reduce the droplet size. Due to the high conductivity of the plasma surrounding the droplet, the electric method might be limited to the laser microcaster described in the following section. In addition to adjustments of the electrode and orifice of the microcaster, the droplet trajectories currently depend on the current and wirefeed setting, and therefore vary for different droplet temperatures. Straight droplet trajectories can be accomplished by creating droplets with melting temperature (those droplets melt at the lower end of the arc and fall straight down) and using an induction heater, which is mounted underneath, and through which the droplets fall, to create superheated droplets. Another method is to track the droplet trajectories during the initial period of the flight with two cameras and laser illumination to overcome the brilliant light of the arc, and to adjust the position of the substrate, or to deflect a previously charged droplet with electric fields. This method would further compensate for straying trajectories, which can be caused through turbulences in the plasma and shroud gas stream, wear of the electrode and orifice, or slight misalignment of the plasma torch and the feed wire.
4.8.1. Laser Microcaster To avoid the varying and straying droplet trajectories of the plasma microcaster, instead of the plasma torch, a laser can be used as a heat source. Initial experiments were carried out with a pulsed 1000 W YAG welding laser. The laser beam was fed through a glass-fiber cable, and focused with a conventional welding head onto the feed wire (Figure 4.29 a). With the laser beam defocused to approximately three time the wire diameter and low power settings, droplets could be created, that accumulated on the wire just underneath the laser beam. The droplet trajectories were consistent and pointing straight down. The droplet temperature was just above the melting point of the material, and too low for reaching substrate remelting (Figure 4.29 b). With increasing laser power, the droplet temperature increased, but the wire melted earlier and the growing
b)
CO2 Laser
Glass-fiber Laser beam
T ~ Tm
Wire Shroud gas Shroud
Welding head
P↓
c)
Substrate
Evaporation
a) P↑
Figure 4.29: Laser Microcaster
87
droplet was directly hit by the laser beam. Due to the energy spike of the laser pulse (on-off ratio 1:5), some of the droplet material was evaporated, the developed gas pressures pushed the droplet away and caused irregular droplet trajectories and spattering(Figure 4.29 c). To avoid the violent evaporation of the droplet due to the laser pulses, further experiments have to be carried out with continuous CO2 welding lasers to reduce peak energy densities. The laser beam has to be split and focused onto the wire from more than one side, to decrease the local energy density on the droplet. The use of hot wire feed reduces the necessary laser energy to melt and superheat the droplets. An induction heater underneath the laser beam could provide additional superheat if necessary. The ultimate solution is the development of a laser microcasting head with concentric wire feed. The laser beam needs to be split into a multitude of smaller beams, which can be positioned in a circle around the wire, or a special beam splitter/modifier has to be developed to convert it into a hollow (tube-shaped) beam, allowing for the wire to be fed through. The tube-shaped beam can then be focused conically onto the wire through a lens with a center hole to allow for the wire to pass through (Figure 4.30). With a laser beam more or less focused onto the wire at different energy settings, part of the laser energy could be used to preheat the substrate directly at the impact point of the droplet. Laser beam Concentric wire feed
Beam splitter Tube-shaped laser beam
Lens
Droplet
Figure 4.30: Concentric Laser Microcasting Head
4.8.2. Closed Loop Microcasting Control High quality deposits, with predictable metallurgical properties and with proper interdroplet and interlayer remelting, require real-time control of the deposition parameters. To achieve the required bonding conditions, the substrate temperature has to be constantly measured at (or near to) the impact point before the droplet is deposited, and the microcasting parameters have to be adjusted accordingly. Independent control of droplet temperature, droplet size, drop rate, standoff distance, path spacing and traversing speed is necessary to obtain optimal deposition conditions. Active cooling and heating of the substrate has to be incorporated to allow control of the substrate temperature during the deposition process.
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5. Support Structures Shape Deposition Manufacturing builds parts by decomposing a geometric model of the part into layers and compacts. The individual compacts and layers of the part are subsequently deposited and shaped. During the manufacturing process a support material is necessary to protect shaped layers from the deposition of material for subsequent compacts, to supplement each layer to provide a complete, flat surface for the next layer, and to shape the undercut features on the outside surfaces of the part. Upon completion of the last layer, the support material is removed from the structure.
5.1. Support Structure Geometry In general, SDM is a multi-materialprocess, i.e., a part can consist of different materials. The compacts of the part are generated, so that each compact consists of only one material. The multimaterial structure is created by depositing subsequent compacts of different materials. To handle the support structure during the planning process, a uniform (one-material) part will be built as a two-material part, for multi-material composite parts, one material (the support material) is added to the materials list. The geometry of the support material, which is basically the space not occupied by the part, can be easily generated by subtracting the model of the part from a block, which completely encloses the part. The only restrictions on the geometry of the block are a sufficient distance from the outside surface of the part to effectively protect the part surface, and it may not contain any undercut surfaces on the outside. Typically cubes, cylinders or extruded shapes, are used. The cross-section of extruded support structures can be obtained, by offsetting the boundaries of the projection of the part onto the x-y plane to the outside. More sophisticated support structure geometries take the actual geometry of the part at a given level into consideration, to provide more economical structures in terms of material deposition. At any particular height, the cross-sections of the support structure for parts, which show smaller features towards the top, can be constructed from the offset projection of the portion of the part at and above the particular height. Other considerations, such as providing more structural stability for the embedded part or the type of support material and the deposition method used can influence the geometry of the support material towards different shapes. E.g., for a powdered support structure, a frame needs to be put around the part and support structure, which leads to simple, extrusion type shapes. Cutting of the support material is simplified, since the outside geometry of the support material does not have to be shaped (the support material is removed after the building process is finished). However, the features inside the support structure, such as the borders between compacts due to splitting of the support structure because of precedence loops, or the undercut features of the part which are build through the support structure, need to be shaped.
89
5.2. Material Requirements The SDM process can use a variety of materials and deposition methods for manufacturing parts with different structural properties. For different materials and deposition technologies, different support materials are required, which must be compatible with the materials of the part. Depending on their geometry, parts can be separated into two different groups. Parts with arbitrary shape and no restrictions on their geometry would be one group. The other group are parts which do not contain undercut features, such as dies used for injection molding. Due to less restrictive manufacturing strategies, which can be used for creating parts without undercuts, the requirements for a suitable support material are different for arbitrarily shaped parts and parts without undercuts.
5.2.1. Support for Arbitrary Geometries For parts with arbitrary geometry, the support material must provide protection for the shaped surface of underlying layers, as well as the ability to provide surfaces to shape the undercut features of the part. Penetration of the material of the part into the support material, due to excessive heat transfer and remelting of the surfaces, will cause undercut features with distorted geometries, poor surface quality and unwanted materials properties due to mixing and alloying, and has to be avoided. In the same way, penetration of the support material into the material of the part has to be avoided, to ensure high quality for non-undercut features. Requirements for materials suitable as support structure for arbitrary geometries: • Deposition of the support material may not penetrate into the previously deposited compacts of the part, nor destroy any of the shaped surfaces of the part. The effects of penetration and surface deformation have to be as minimal as possible. • The support material must adhere well enough to the materials of the part, to provide structural strength for the shaping operation and the embedded part. • The support material must be shapeable (machineable), to provide the surfaces for building undercut features. • Deposition of the materials of the part may not penetrate into the support material, or destroy the support material surfaces used to shape undercuts. The effects of penetration and deforming of shaped surfaces have to be as minimal as possible. • The materials of part must adhere well enough to the support structure to provide structural strength for the shaping operations and to prevent warpage due to internal stresses. • The support material must be removable after completion of the building process.
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5.2.2. Support for Parts without Undercuts The second group of geometries, parts without undercuts, are a widely used group of parts, such as industrial injection molding tools. For such parts some of the requirements of the support material for arbitrary geometries are not necessary. In general it is possible to access all surfaces of the part during the shaping process, a limitation in depth is only imposed by the geometry of the cutting tool. Penetration of the material of the part into the support material during the deposition process can be tolerated, by removing and redepositing the penetrated areas of the support material. A wider range of support materials can be found and because different material deposition processes can be used, penetration and surface deformation of the part by the support material during the deposition can be eliminated. Requirements for a material suitable as support structure for geometries without undercuts: • Deposition of the support material may not penetrate into the previously deposited compacts of the part, nor destroy any of the shaped surfaces of the part. The effects of penetration and surface deformation have to be as minimal as possible. • The support material must only protect the surfaces of deeper lying layers of the part. The immediately preceeding layers are still accessible and can be reshaped. • Structural integrity of the support structure is only required to withstand the deposition process. • Deposition of the materials of the part can penetrate into the support structure, to a depth that is accessible by the shaping process. However, minimum penetration is desired. Adhesion of the part materials to the support structure is not required. • The support material must be removable after completion of the building process.
5.2.3. Removal of the Support Structure For removing the support structure from the finished part, special consideration has to be given to the removal method, so that the structural integrity of the part is kept intact and the part surfaces are not destroyed. While machining away the support material is not one of the available options (SDM and the support material would not be the optimal process to manufacture such a part), different solutions are possible, depending on the type of material used. Water-soluble support materials can easily be dissolved, materials with lower melting points than the materials for the part can be melted out. Support structures from bound powders can be cracked by shock waves and are removed with air or water jets. Support structures for high temperature metals, which have to withstand high temperatures during the deposition process, are typically removed by chemical means.
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5.3. Materials Penetration While a small penetration (remelting of underlying material) is desired to achieve high material strength at the interfaces of layers and compacts when the material of the part is deposited onto areas with the material of the part, penetration of the part materials into the support material or vice versa during the deposition processes has to be avoided or kept to an absolute minimum. Figure 5.1 shows penetration of the support material into shaped non-undercut surfaces of the part (a) or penetration of the parts materials into shaped surfaces used for the creation of undercuts (b), resulting in a deteriorated surface quality of the part and in severe cases in the destruction of the geometry of the part. Penetration of material into the support on a flat interface along the edge between the support and the part at the underlying layer causes thin fins (c), which will extend from the part at the interfacial line of the layers. In the same way, penetration of the support material along interfacial lines on the bottom of a layer (d), can cause tiny fissures in the part along the interfacial lines of the layers. On the top edges of previously deposited compacts of the same layer, penetration can cause rounding of the edges and corners, leading to fins (e) and fissures (f). Penetration of the support material into the part over the interfacial area is extremely critical. Inclusions at the interfacial area (g) reduce the interfacial strength, and especially in case of thermal removal of the support material, if the interfacial strength is based on bonding between the material and the support material inclusions, or when the inclusions lead to cracking, the part can delaminate at the interfacial area. The other case, inclusions on interfacial areas of the support material (h), are not critical for geometric reasons, but might impair the integrity of the support material, and cause delamination of smaller portions of the support structure during shaping operations.
g
h
f
e
d
a
c
b
Figure 5.1: Materials Penetration If penetration can not be fully controlled through process parameters, for certain cases geometric measures can be taken to minimize the influence on the surface quality. While there is no solution to avoid the problems caused by the cases shown in Figure 5.1 a) - d), e) - h) can be eliminated by having additional material above the interfacial plane between the layer, on the top of the previous, as well as the current layer. In Chapter 6.2 this process is explained in detail.
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5.4. Support Materials for Different Applications Using the specifications of the previous section, a variety of combinations of part materials with suitable support structures have been identified. Among the possibilities are combinations of different thermally sprayed alloys and a variety of different approaches for high temperature metals deposited through microcasting. Other possible combinations are fast curing plastics (UV-curable resins, two-component epoxies) and support structures from water-soluble materials or low melting point waxes, which can be helpful for the testing and development of some SDM strategies.
5.4.1. Support for Thermally Sprayed Structures Thermal spray processes, in particular arc and plasma spraying, are capable of depositing a variety of different materials with a minimal heat impact on the underlying material. By controlling the spray parameters in ranges with low heat input, materials with higher melting points can be deposited onto materials with lower melting points without destroying the surface and geometry of the low melting material. The material with the lower melting point is only superficially or not remelted at all. A low melting point alloy can therefore be chosen as the support material for structures created through thermal spraying. After the building process is finished, the part is melted out from the support material. However, due to limitations in controlling the process parameters, the approach of spraying materials with different melting temperatures is not suitable for the creation of fully functional parts. In order to deposit materials with high melting temperature onto the low melting alloy the power settings on the spray torch have to be kept extremely low, to prevent melting of the low melting material. In many cases, operating conditions with low power settings are not optimal for the deposition of higher melting materials and result in poor properties of the deposit. The individual, not sufficiently hot droplets form lamellae, which are not melted to each other and cause fair amounts of porosity. Bonding conditions are especially bad at the layer interfaces and can lead to delamination of the structure due to thermal stresses during the melt-out process. Several parts have been built with thermal spraying of different materials. A tin-zinc alloy (91% Sn, 9% Zn) with a melting point of 473 K was used as the support material. Attempts using plasma sprayed materials known as bondcoats, e.g. 95% Ni - 5% Al (Metco 450 powder), or selfbonding, free-machining materials, e.g. 52% Fe - 38% Ni - 10% Al (Metco 452 powder), were unsuccessful. Tough material properties caused by low power settings to prevent remelting of the support material produced excessive wear of milling cutters to the point where the material became uncutable, or caused chipping of the material during the cutting operations. Structures sprayed from zinc layers, with a thin Ni-Al bondcoat to increase adhesion between the individual layers, produced better results. However, the difference in melting points between zinc and the support material is low, and during the melt-out process, some of the zinc structures delaminated at layers with high stress concentrations. Using zinc with a Ni-Al bondcoat and a tin-zinc support structure has basically shown the feasibility of the approach using sprayed material deposition only. To achieve reliable results, without
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the possibility of parts delamination during the melt-out process, compatible combinations of materials, support material and bond-coat have to be identified. Due to the limited physical properties of sprayed structures in general, which make it unlikely for sprayed structures to be used as fully functional parts, the work in this thesis has focused on the development of support materials and material combinations for the microcasting process.
5.4.2. Low Melting Support with Thermal Barrier for High Temperature Metals Materials with high melting points, deposited with the microcasting process, produce a much greater heat input into the substrate than sprayed deposits, and cause a significant temperature rise at the interface zone of the substrate. While support structures with low melting points have the advantage of easy removal by melting at comparably low temperatures (450 - 700 K), they cannot withstand the temperature rise caused by directly impacting droplets. 5.4.2.1. Interface Temperature The interface temperature between the droplet and the substrate at the moment of impact is important for the remelting conditions. At the first instance of the droplet impacting onto the substrate, the droplet and the substrate behave like two semi-infinite half-spaces. The interface temperature Ti of two semi-infinite half-spaces, which are brought into perfect contact, immediately adjusts to a value, which depends on the thermal properties and the initial temperatures of the half-spaces [48].
i
T1 ρ1 c1 λ 1 + T2 ρ 2 c2 λ2 = ------------------------------------------------------------ρ1 c1 λ 1 + ρ 2 c 2 λ 2
(5.1)
T1 and T2 are the initial temperatures, ρ1, c1, λ1, and ρ2, c2, λ2 respectively are the densities, specific heats and thermal conductivities. In case of the two infinite half-spaces, the value of Ti remains constant at the interface, and the temperature profiles to both sides of the interface follow error functions. If the liquid and solid state of a material are brought into contact, the interface temperature adjusts instantaneously to the melting temperature Tm of the material. Solidification of the liquid half-space will occur if the interface temperature Ti is lower than the melting temperature Tm, melting of the solid half-space will occur if the interface temperature Ti is above the melting temperature Tm, and the interface will move [49]. For two materials (one liquid, one solid) with different melting points, four different cases are possible. If the interface temperature is above the melting temperature of the liquid and the solid, the solid will remelt, if the interface temperature is below the melting temperature of the liquid and the solid, the liquid will solidify. For an interface temperature above the melting temperature of the solid, but below the melting temperature of the liquid (the liquid material has a higher melting point than the solid material), the solid will remelt, and the liquid will solidify. Two new interfaces will be created. A melting front moves into the originally solid half-space, a solidification
94
front moves into the originally liquid half-space. For an interface temperature above the melting temperature of the liquid and below the melting temperature of the solid (the melting temperature of the solid is higher than the melting temperature of the liquid) neither solidification nor remelting will occur. The previous conclusions are made for semi-infinite half-spaces, with an infinite supply of thermal energy (i.e., the temperatures at the infinite end of the half-spaces remain constant). Liquid droplets being deposited onto a substrate, have only a finite amount of thermal energy available and the temperature of the droplet will decrease. In case of substrate remelting, the melting front will stop and reverse after a certain time. The interface temperature Ti from (5.1) can be used as an indication, whether remelting will occur. 5.4.2.2. Deposition of Dissimilar Materials To estimate the effect of a liquid, high temperature metal droplet on a low melting substrate, numerical simulations have been performed1. A one-dimensional thermal model has been developed to simulate the droplet and substrate temperatures and the migration of the melting front. Due to the small droplet size, compared with the size of the substrate, the droplet height had to be modified in the numerical model, to match the simulated temperature of mild steel droplets impinging on mild steel substrates with experimental measurements [37]. The one-dimensional model was used to simulate the temperature distribution and the melting front migration of a 100 µm thick stainless steel layer at a temperature of 1723 K being deposited onto a zinc substrate at room temperature (298 K). The results shown in Figure 5.2 indicate, that the zinc, with a melting point of 693 K is remelted to a significant depth, while the lower portion of the steel droplet freezes immediately. Investigations of the temperature values show, that the upper portion of the zinc substrate actually vaporizes. Since evaporation was neglected in the numerical model, the graph in Figure 5.2 shows only the coexistence of 4 different states: the solid portion of the zinc substrate, the liquid zinc interface to the steel droplet, the solid (frozen), lower portion of the steel droplet and the upper liquid portion of the droplet. 1823 K St.St. Droplet on 298 K Zn Substrate 100
Zinc
Melting Front: Stainless Steel
Liquid
50
Solid
0 d [µm] -50
Liquid
-100
Solid
-150 -200 0
0.5
1.0 t [ms]
1.5
2.0
Figure 5.2: Melting Front Simulation of Stainless Steel Droplet on Zinc Sub-
1. The numerical simulations using a one dimensional model where performed by Kevin S. Schmaltz at Carnegie Mellon University.
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Experimental tests of depositing mild steel droplets directly onto substrates with low melting points showed similar results. Superheated steel droplets being deposited onto zinc or tin-zinc alloys (91% Sn, 9% Zn) resulted in evaporation and remelting of the substrate surface. The gases produced by the evaporating material caused the droplets to spatter violently. Deposition of mild steel droplets onto aluminum substrates caused significant remelting of the substrate surface, and alloying of the aluminum with the steel produced uncutable materials. 5.4.2.3. Substrate Temperature during Deposition To estimate the thickness of a metal thermal barrier, which is required to protect the low melting support structure from the superheated droplets, the temperature inside a substrate have been measured during the deposition of a droplet. Type K thermocouples (Ni-Cr and Ni-Al, temperature range 73 - 1523 K) with a wire diameter of 25 µm and the PC based data acquisition system described in Chapter 4.5.2 have been used. The thermocouple temperatures were calculated using the routines provided with the “Data Translation PCLAB” software library, which was used to program the data acquisition card. To insert the thermocouples (Figure 5.3), a 3.2 mm diameter hole was drilled into the back of a 12.5 mm thick mild steel substrate (1020 steel). A ceramic spacer was used to insulate the bare thermocouple wires. The thermocouple junction was touching the substrate at a distance of 1.27 mm from the top surface of the substrate. The droplets were placed to hit directly above the thermocouple. Hot droplet (TD) Solidified Droplet Thermocouple (TS)
Ceramic Spacer
Figure 5.3: Measurement of Substrate Temperature Figure 5.4 a) shows the result of two impacting droplets with a diameter of 3.53 mm and a temperature of 2497 K. The first droplet impacted on the substrate, and rolled about 2 mm before adhering to the substrate. Due to this offset from the center-line of the thermocouple, the temperature peak measured in the substrate at 587 K is slightly lower than for a direct hit. The second droplet adhered at the impact point and registered a substrate temperature of 681 K. In Figure 5.4 b) a deposit of 6 consecutive droplets with a diameter of 3.61 mm and a temperature of 2596 K was placed onto the substrate. The first droplet showed a peak temperature of 574 K in the substrate. The following two droplets were directly impacting on top of the first droplet, thus causing lower peak temperatures due to the increased distance from the thermocouple. When the forth droplet was deposited, it was sliding off the previously deposited droplets and created a peak temperature of 663 K.
96
50
50
681
00
00
50 00
587
00 574 50
T [K]
50
T [K]
663
50
00
00
50
50
00
00
50
50
00
00
50
50
t [s]
t [s]
a)
b)
Figure 5.4: Substrate Temperature Curves (1020 substrate, ER70S-6 droplets)
The measurement of the substrate temperature 1.27 mm below the impacting point of the droplets show, that for droplet temperatures of approximately 2500 K a temperature in the area of 700 K can be expected in materials with a thermal conductivity similar to steel. For the protection of low melting material, a metal thermal barrier coating with a thickness between 1 and 2 mm (in case of a steel coating) is required. The thickness of the coating can be kept lower, if remelting of the underlying support material can be tolerated, and the thermal barrier, rather than the low melting material, is used to potentially shape the deposit. 5.4.2.4. Thermal Barrier Protection of Low Melting Support Structure To protect structures from the effects of high temperatures, typically materials with low thermal conductivity and high melting points are used. Materials, such as various ceramics, can be deposited through plasma spraying. However, the heat of the plasma torch, which is required to melt the ceramic powders, is very high and causes low melting support structures to melt. Metal droplets being deposited onto ceramic substrates, show poor wetting conditions, which result in extreme difficulties to achieve acceptable deposits. For undercut features, the thermal barrier has to becomes a part of the structure, which is unacceptable for functional parts. Further, shaping ceramic material together with metals is extremely difficult and un-economical. To achieve satisfactory material properties of the resulting part and the necessary wetting conditions for the microcast droplets, it is proposed to use the same material as is used for the droplets to thermally spray the barrier protection for the support material. Figure 5.5 depicts the individual manufacturing steps for the creation of a layer (compact) showing both undercut and non-undercut features. The previous layer is encased in the low melting support material (a). More support material is deposited and shaped to provide the surfaces for the undercut features (b). For areas of the part, where non-undercut features are being built on top of non-undercut features of the previous layer, the top surface of the low melting support material can be cut underneath the top of the
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previous layer. The thermal barrier to protect the low melting support material is then deposited (c). To allow high interlayer strength through the metallurgical bonding of the microcast droplets, the thermal barrier is removed from most of the area inside the part (d). The material for the part can then be deposited (e), the remaining features of the layer (compact) can be shaped, and the thermal barrier protecting the support material next to non-undercut features is removed (f). In the areas of undercut features, the thermal barrier remains a portion of the resulting part (g). Due to the different properties of microcast and sprayed materials and different bonding conditions, the interface between the microcast material and the thermal barrier can be a potential source for cracks and cause failure of the parts.
a) previous layer and support cut underneath top of previous layer
e) deposit part material remove thermal barrier
b) create low melting support
f) shape part c) deposit thermal barrier
g) remove support d) shape thermal barrier
Figure 5.5: Low Melting Support Structure with Thermal Barrier
This technique has been used with a tin-zinc alloy (91% tin, 9% zinc, melting point 475 K) as a support material and 1mm thick arc sprayed thermal barriers. The material used for the thermal barrier was the same material as the material deposited on top of the barrier through microcasting. The measurements of the substrate temperature in the previous section indicate, that a thicker barrier should have been used to prevent melting of the support structure. Experiments showed, that slight remelting of the low melting support structure can be tolerated. However, it is necessary, that either the support material is enclosed inside the part itself (e.g. hollow structures, cooling channels), or a frame or a sprayed shell is placed around the support material. Holes to allow for the expanding, liquid support material to escape into areas where it does not interfere with the deposit of the high temperature metal, have to be provided. Cracks in the thermal barrier produced
98
by the thermal shock from the impacting droplets due to the material properties of the sprayed coating, or from the expanding support material due to the lack of openings in the encasing shell or frame, can be a potential problem. Liquid support material can break through the barrier and when directly hit by a droplet, disturbances and voids can be caused in the structure. Examples are shown in Chapter 8.2.1.
5.4.3. Powder Support for Microcast Structures without Undercuts Support materials for parts without undercuts are not used for shaping or to provide support for overhanging features. They are only used to protect surfaces of lower layers from the deposition of the part material. Parts without undercuts and with only one material consist of one compact per layer. The support material, which is the second compact to complete the layer, is deposited after the material of the part has been applied, and does not need to be shaped. The one-compact geometry and the accessibility of parts without undercuts allow the shaping of underlying layers even after subsequent layers have been deposited. This permits the use of support materials which are slightly penetrated during the deposition of the material for the part. One such group of materials are various bonded powders, which can be used similar to the techniques used for sand mold creation. Experiments using ceramic and metal powders have been conducted. Ceramic powders showed poor wetting conditions for microcast metal droplets. Low thermal conductivity in the ceramic structure causes different conditions for the droplet deposition than in metal areas. Droplets deposited on the ceramic support cool very slowly. They stay liquid long enough to flow together to form bigger droplets, causing irregular deposits on the edges of the part and in areas of thin features. Support structures from metal powders, which have thermal properties similar to those of the growing part, provide better conditions for the deposition process and allow the creation of uniform layers over the solid as well as the powered portions of the underlying layer. Typically, powers of the same material as the material for the part were used. A sodium silicate binder was used to form an inorganic bond between the powder particles. Sodium Silicate is a typical binder used for the creation of sand molds used for large cores and castings, where hardness and dimensional control are essential [50]. Sodium silicate is essentially silicic acid containing large quantities of colloidal sodium. Carbon dioxide is used to precipitate the sodium and form a silicate polymer. continued gassing with carbon dioxide gives: Na2O 2SiO2 + 2CO2 + H2O ⇔ 2Na2HCO3 + 2SiO2
(5.2)
This shows, that continuous gassing dehydrates the amorphous silica gel and increases the strength of the support structure. Powder with a particle size between 45 and 150 µm is typically mixed with sodium silicate in a mixture of 80 vol.% powder and 20 vol.% sodium silicate. The sodium silicate is dehydrated by supplying carbon dioxide at a flowrate-density of 0.18 l/min cm2 over a period of 2 minutes. Any remaining water is removed by heating the support structure with
99
a flame torch for approximately 5 minutes. The final temperature of a typical support structure was measured to be 353 K. Residual water in the support structure can instantly evaporate when material is deposited through microcasting. The developing water vapors can disturb the deposition process and cause hollow deposits or voids. To remove the support material after completion of the part, methods employed for removing sand molds and cores in casting applications, can be used. The thermal shock during the deposition process causes fine cracks in the support structure. High velocity air jets or water jets can be used to further break up the support structure and remove the pieces from the part. Figure 5.6 shows the manufacturing steps for creating the support structure and manufacturing a layer. The part is started directly on the substrate, with no support material layer in between. After completion, the substrate is removed and the bottom surface of the part is machined with conventional techniques. For each layer the metal powder is mixed with the sodium silicate and is filled in the empty sections of the layer (a). A frame surrounding the part is provided to hold the powder in place. The frame can either be mounted to the substrate in incremental, thin sections, to match the height of the growing part, or a wall can be deposited with the microcasting process. The powder is then slightly compressed (b) to ensure a uniform and proper density of the material. After dehydrating the sodium silicate with carbon dioxide (c) the support structure is heated to evaporate residual water (d). Excess powder isremoved from the top (e) and the top surface is sanded to create a flat surface of the support material and to remove oxides or powder remains from the top cross-sections of the part. The material for the next layer of the part is deposited, showing penetration into the powder support structure (f). Finally, the new layer is shaped. To remove the results of the penetration, the previous layer(s) are also machined to a depth slightly deeper than the penetration (g). metal powder + sodium silicate
frame
penetration
a) deposit powder pressure
f) deposit part material CO2
remove penetration
b) compress powder
d) heat
c) CO2 gassing
g) shape top layers
e) clean top surface
Figure 5.6: Bound Powder Support Structure
An example of one half of an injection mold, which was built with a metal power support struc-
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ture is shown in Chapter 8.2.3. The material of the part is microcast ER70S-6 mild steel, the powder used as support material was a mild steel powder commonly used for plasma spray applications (Plasmalloy AI-P230). The 80:20 volume percent mixture of powder and sodium silicate translates into a 10:1 weight mixture for the mild steel powder with a particle sizes between 45 and 150 µm. The penetration depth of a deposit of 3.61 mm diameter mild steel droplets at a temperature of 2596 K was between 1.0 and 1.5 mm.
5.4.4. Copper Support for Microcast Steel Structures The support materials discussed in the previous sections, do not have the ability to be used directly for microcast structures with undercuts. Low melting support materials have to be protected with thermally sprayed coatings, bound powders show a significant amount of penetration, and ceramics are difficult to shape and remove. An alternative seems the use of high temperature metals, with melting points comparable to those of the material used for the part and chemical methods for the removal of the support structure. 5.4.4.1. Remelting Conditions for Dissimilar Materials The ideal combination of two metals for the part and the support structure should have the following behavior in terms of substrate remelting: A liquid droplet with a temperature TD should remelt the portions of the substrate with a temperature TS, which are of the same material, while it does not remelt the portions of the substrate (with the same temperature TS), which are of the other material. Rewriting (5.1) for the interface temperature at the instance of contact gives an indication of the required thermal properties. TD – TS T i = T S + ---------------------------------ρ S cS λS 1 + ------------------ρD cD λD
(5.3)
For the purpose of estimation, the specific heat and the density of the liquid and solid state can be assumed to be comparable. According to the Dulong-Petit rule the product of density and specific heat of metals at 300 K is approximately constant [51]. ρ ⋅ c ≈ const
(5.4)
Values reported in [44] for the thermal conductivity λ and the thermal diffusivity α (with α = λ / ρ c) show, that ρ · c is comparable for iron, stainless steel and copper at room temperature and close to the melting temperatures of the materials (Table 5.1). Considering, that the changes in the product of density and specific heat are only small, and that the thermal diffusivities of liquids during pouring are orders of magnitude higher than for solid materials [52], the density and specific heat can be eliminated from (5.3). The remaining material
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ρ · c (at 300 K) [J/cm3K]
ρ · c (at ~Tm) [J/cm3K]
(ρ·c)Tm / (ρ·c)300
Fe
3.96
5.26
1.33
Stainless Steel
3.60
5.08
1.39
Cu
3.40
4.52
1.33
Table 5.1: Product of Density and Specific Heat for Steel and Copper properties influencing the interface temperature are the thermal conductivity of the substrate and the droplet. TD – TS T i ≈ T S + ------------------λS 1 + -----λD
(5.5)
In [49] it is shown, that for remelting of a substrate of similar material, the interface temperature Ti must be bigger than the melting temperature Tm. For a material being deposited onto a dissimilar substrate, the interface temperature Ti must remain below the melting temperature Tm,S of the substrate, to avoid remelting. According to (5.5) it is possible to find two materials, so that each material being deposited at a temperature TD,mat can create remelting on a substrate of the similar material with temperature TS, but does not remelt a substrate with the same temperature TS of the other material. The materials have to have different melting points, and the material with the higher melting temperature has to have a much lower thermal conductivity than the other material. (5.5) shows, that the influence of the temperature difference between the droplet and the substrate TD - TS on the interface temperature Ti becomes smaller with increasing ratios of the thermal conductivity between the substrate and the droplet. This means, that the interface temperature is higher, if a droplet with the same temperature is deposited onto a substrate with lower thermal conductivity than for a substrate with higher thermal conductivity. For the droplet with the higher melting temperature, the interface temperature will be lower, when it is deposited onto the material with the lower melting point (which has the higher thermal conductivity), than when it is deposited onto the similar material. For the droplet with the lower melting temperature, the interface temperature is higher when it is deposited onto the second material (due to the lower thermal conductivity), than when it is deposited onto the similar material. To find two materials with the right thermal properties and melting temperatures, and to find the range of droplet and substrate temperatures, the following conditions can be derived from (5.5). The melting temperatures of the part material and the support material are Tm,obj and Tm,sup, the temperatures of the liquid droplets are TD,obj and TD,sup, the temperatures of the substrates (i.e., for both materials) are TS1 for depositing the part material, and TS2 for depositing the support material, and the thermal conductivities are λS,obj and λS,sup for the solid substrate materials and λD,obj and λD,sup for the liquid droplets. Deposition of the part material onto substrate areas consisting of the part material requires interface temperatures higher than the melting temperature of the part material in order to achieve
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remelting of the substrate. T D,obj λ D,obj + T S1 λ S,obj T m,obj < ---------------------------------------------------------------λ D,obj + λ S,obj
(5.6)
On substrate areas, where the part material is deposited onto the support material, the interface temperature has to be lower than the melting temperature of the support material. T D,obj λ D,obj + T S1 λ S,sup T m,sup > ---------------------------------------------------------------λ D,obj + λ S,sup
(5.7)
Similar relations can be derived for depositing the support material. The droplets have to produce an interface temperature high enough to remelt the substrate areas consisting of support material, T D,sup λ D,sup + T S2 λ S,sup T m,sup < ----------------------------------------------------------------λ D,sup + λ S,sup
(5.8)
and an interface temperature below the melting point for areas with the part material. T D,sup λ D,sup + T S2 λ S,obj T m,obj > ----------------------------------------------------------------λ D,sup + λ S,obj
(5.9)
With kso being the square root of the ratio between the thermal conductivity of the support material and the part material at substrate temperature, and kls the square root of the ratio between the thermal conductivity of the liquid droplet and the solid material at substrate temperature, λ S,sup -, k so = ---------------λ S,obj
λ D,obj k ls,obj = ----------------λ S,obj
and
λ D,sup k ls,sup = ----------------λ S,sup
(5.10)
(5.6) through (5.9) can be used to find the relation between the substrate temperatures TS and the droplet temperatures TD,mat. For the deposition of the part material with a droplet temperature TD,obj, the substrate temperature TS1 has to meet the condition T S1 > ( 1 + k ls,obj )T m,obj – k ls,obj T D,obj
(5.11)
to metallurgically bond the droplet to substrate areas with the part material and 1 T S1 < --- [ ( k ls,obj + k so )T m,sup – k ls,obj T D,obj ] k so
(5.12)
to prevent substrate areas with the support material from remelting. For the deposit of the support material with a droplet temperature TD,sup the substrate temperature has to satisfy
103
T S2 > ( 1 + k ls,sup )T m,sup – k ls,sup T D,sup
(5.13)
for metallurgical bond of the support material and 1 T S2 < k so k ls,sup + ------- T m,obj – k ls,sup T D,sup k so
(5.14)
to avoid penetration into the part. The thermal conductivities of materials in their solid (room temperature) and liquid (melting point) state are different. For materials like copper or iron, the values for the thermal conductivity are smaller at the melting point than the values for room temperature. According to (5.11) and (5.13) significant amounts of superheat and high substrate temperatures would be required to achieve remelting of the substrate for the deposit of similar materials. However, extremely high cooling immediately after the impact of the droplet has been observed, which can be accounted for by artificially raised thermal conductivities of the droplets. A similar behavior is found in usual ingot and shaped casting processes. Experimentally it has been established, that superheat is nearly or completely eliminated soon after pouring due to the effects of convection. In numerical solutions, this is accounted for by assuming artificially high liquid thermal conductivities, which are 10 to 100 times higher than the static value [52]. λ D = υ ⋅ λ D,static
with
10 ≤ υ ≤ 100
(5.15)
With increasing thermal conductivities of the liquid droplets, lower substrate temperatures and less superheat are necessary, to achieve remelting conditions. Since the exact conditions for the artificially high liquid thermal conductivities for the droplets are not known, several factors within the range known from casting are used for the purpose of estimating the required droplet and substrate temperatures. 5.4.4.2. Estimated Temperatures for Stainless Steel and Copper The exceptionally high thermal conductivity of copper, while its melting temperature is lower than the melting point of typical high temperature metals (Fe, Ni, Cr, Mo, Ti), make it an ideal candidate for the support material. Stainless steel on the other hand, shows an extremely low thermal conductivity, but high resistance against chemical corrosion. The properties for 308 stainless steel, which was used for the experiments and to create parts, have been taken from [53] for the calculations. The values reported for the thermal conductivity at 373 and 773 K have been extrapolated to get estimates for the values for 300 and 1600 K. The melting temperature Tm,stst and the thermal conductivities λ300,stst and λ1600,stst are T m,stst = 1683K ,
λ 300,stst = 14.0W ⁄ mK
and λ 1600,stst = 34.83W ⁄ mK .
(5.16)
The extremely low values of the thermal conductivity of stainless steel, and the increase of the thermal conductivity with increasing temperature, result in extremely good remelting conditions
104
for low substrate temperatures and low values of superheat (see (5.5)). Experimentally, this can be observed by excellent wetting and bonding conditions of stainless steel deposits. For copper, the thermal conductivities λ300,Cu at 300 K and λ1200,Cu at 1200 K [48] and the melting temperature Tm,Cu are T m,Cu = 1356K ,
λ 300,Cu = 401W ⁄ mK
and
λ 1200,Cu = 339W ⁄ mK .
(5.17)
With these values, the ratio kso between the square roots of the thermal conductivities of the support material and the object material at substrate temperature (assumed to be room temperature) is k Cu,stst = 5.35 ,
(5.18)
the ratios of the square roots of the thermal conductivities between the droplets and the substrate are listed in Table 5.2 for different factors ν (10, 25, 50 and 100). ν = 10
ν = 25
ν = 50
ν = 100
kls,stst
4.99
7.89
11.15
15.78
kls,Cu
2.91
4.60
6.50
9.19
Table 5.2: Square Root of Thermal Conductivity Ratios Using these values, the ranges for the substrate and droplet temperatures have been evaluated. Figure 5.7 a) shows required substrate temperature TS1 for a deposit of stainless steel droplets with the temperature TD,stst. The lower boundary, expressed by (5.11), is shown in blue color for different thermal conductivities of the liquid droplet. In order to achieve remelting of the stainless steel portions of the substrate, the droplet/substrate temperature pair has to be above this boundary. For typical substrate temperatures at or above room temperature (300K), it can be seen, that regardless of the thermal conductivity of the droplet, remelting can be achieved with droplet temperatures of 2000 K or above. The upper boundary given by (5.12) is shown in red for the different values assumed for the thermal conductivity of the droplet. The droplet/substrate temperature pair has to stay below the upper boundaries in order to prevent remelting of the copper support material. While the droplet temperatures required for remelting the stainless steel are relatively insensitive to the substrate temperature and the thermal conductivity of the droplet (values are approximately within 200°), the remelting conditions for the copper portion of the substrate, show a greater dependency on the thermal conductivity of the droplet. For the higher values of the thermal conductivity, no realistic or only a small range is available for the substrate and droplet temperature, due to the tendency towards remelting which increases faster for the copper than for the stainless steel. In case of the substrate temperature, which has to be kept below 400 K (for kls,stst = 11.15) this might be difficult to accomplish due to heating of the substrate during a continuous deposit. For the smaller values of the thermal conductivity of the droplet, a wider range is available. For a medium value of kls,stst (kls,stst = 7.89) droplet temperatures between 1900 and 2100 K with substrate temperatures from 700 to 300 K will produce the required remelting characteristics. In case of low values for the liquid thermal conductivity, the droplet temperatures up to 2500K
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with room temperature substrates are possible. 000
000
kls,stst 4.99 7.89 11.15 15.78
800 600 400
600 400
TS2 [K]
200
TS1 [K]
kls,Cu 2.91 4.60 6.50 9.19
800
000
200 000
800
800
600
600
400
400
200
200
0
0
TD,stst [K]
TD,Cu [K]
a) Temperature Ranges for Stainless Steel Deposit
b) Temperature Ranges for Cu Deposit
Figure 5.7: Droplet and Substrate Temperature Ranges for Stainless Steel and Copper
Figure 5.7 b) shows the evaluation of (5.13) and (5.14) for impinging copper droplets. The upper boundary (red) for remelting the stainless steel areas of the substrate shows little influence of the thermal conductivity of the liquid copper droplets and the substrate temperature. To avoid remelting, droplet temperatures of 1700 K or below are required. Due to the higher conductivity of the copper substrate, the lower boundary for the temperature of the copper droplets shows a greater dependency on the thermal conductivity of the droplets. For reasonable substrate temperatures (room temperature or above) droplet temperatures ranges with a difference of 50° to 250°, depending on the thermal conductivity of the droplet, are possible. Looking at the relation between the upper and lower boundaries, a very important conclusion concerning the substrate temperatures can be drawn. For the deposition of stainless steel droplets, the available range of droplet temperatures increases with lower substrate temperatures for a given thermal conductivity of the droplet. For the copper droplets, on the other hand, the range of possible droplet temperatures increases with higher substrate temperatures. To reach process conditions, which are more insensitive towards the thermal conductivity of the droplets, which in all essence is not exactly known, and to have a wider range of droplet temperatures available, to allow bigger amounts of superheat for better remelting and wetting conditions to the similar substrate, for stainless steel deposit the substrate temperature should be kept as low as possible, for deposit of copper the temperature should be kept high. In praxis this means, that for the deposition of copper extensive preheating of the substrate is necessary, and for the deposition of stainless steel the substrate needs to be actively cooled. The range of values derived for the substrate and droplet temperatures show, that it is basically possible to deposit two materials with different melting points and thermal conductivities, so that portions of the substrate with similar material are remelted, while the portions with the other material stay solid. The values for the temperatures, however, are only estimates, since the exact
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conditions of the heat transfer between the droplet and substrate are not fully known. Furthermore, the thermal properties of rapidly solidified material, such as material deposited with the microcasting process, have not been established. Non-ideal contact between the droplet and the substrate and surface oxides can lower the temperature at the interface (on the substrate side), making higher droplet temperatures possible and necessary to achieve remelting of the similar substrate areas. The interface temperatures are further only valid at the moment of impact. The temperature after the impact and the depth of remelting are influenced by the size of the droplet. While the droplet size has no influence on the occurrence of remelting under the ideal assumptions, under non-ideal conditions, the droplet size plays an important role in the remelting conditions. During the deposition of the part material and the support material, it is not always desired to keep the temperatures within the area limited by the previous calculations, where the droplets do not remelt the dissimilar material. Due to the rapidly cooling droplets and the small droplet volumes (compared with a typical substrate), the remelting depth for increased droplet temperatures is low (in the order of µm). Further, increased droplet or substrate temperatures, which exceed the upper limit, may be required to achieve sufficient wetting of the droplets for deposits without undercuts and voids. This is especially apparent for the deposit of copper droplets. Experimentally it was determined, that droplet temperatures several hundred degrees above the upper boundary were necessary, to create acceptable copper deposits. Unfortunately, this resulted in small penetration of the copper into stainless steel substrate areas. A small amount of penetration into the dissimilar material, an therefore droplet/substrate temperature pairs above the upper (red) boundaries may be required, to establish a bond, which is strong enough to withstand forces exerted by thermal stresses and during the shaping process. A compromise between penetration depth and bonding strength is necessary. Also, active control of substrate temperatures during the deposition process is necessary, to keep the temperature at a constant value. Without control, heat from the droplets transferred into the substrate, causes the substrate to heat up and changes the bonding conditions. Droplets at the beginning of the deposition might not adhere to similar substrates, due to low initial substrate temperatures, droplets at the end of the deposition might cause excessive penetration into dissimilar substrate material, due to the increased temperature from the heat input from the droplets. 5.4.4.3. Support Material Removal Melting the support material after the building process has been completed is in principle possible, since the melting temperature is lower than the melting temperature of the part material. Due to the comparably high melting point of the support material, which is required to achieve adequate remelting conditions, this is connected with a series of problems. Handling and cleaning the part at extremely high temperatures is difficult, a controlled atmosphere is necessary to prevent oxidation, and heating the part close to its melting point causes distortion. As an alternative the support material can be removed with chemical means. Copper, for example, shows high corrosion rates in nitric acid [54]. High silicon or chromium steels, on the other hand, show outstanding corrosion resistance in nitric acid. Stainless steels with a chromium content of more than 17% (such as 430 or 304) are typically used in applications involving nitric acid. While
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for some stainless steels proper heat treatment is necessary to avoid failure in the material, e.g. rapid cooling for 304 stainless steel, for austenitic grades, such as 304L or 308, satisfactory resistance to nitric acid corrosion can be achieved without quenching. However, rapid cooling conditions are given by the nature of the microcasting process anyway. Corrosion rates for 304L (18% Cr, 8% Ni) and 347 stainless steel are reported to be 2.4 nm/min for 70 - 90% nitric acid at 343 K [53]. Due to the higher chromium content the corrosion rates of 308 stainless steel (19 - 21% Cr, 10 - 12% Ni) can be expected to be even lower. The corrosion rates for microcast copper have been measured in 69 - 71% nitric acid and were 0.225 mm/min at a temperature of 333 K. Considering the thermal properties, the evaluation of the interface temperatures during the deposition process and the corrosion properties in nitric acid, copper and stainless steel seem to be suitable materials for the support structure and the part, respectively. Typically, the support material is removed in 70% nitric acid at a temperature of 333 - 348 K. Another important material to be used with a copper support structure is titanium, which has corrosion rates smaller than 48 pm/min in 70% nitric acid at a temperature of 473 K [53]. A critical issue to be addressed with the corrosive removal of the support material is stress corrosion cracking of the part material. Stress corrosion cracking is caused by the simultaneous presence of tensile stresses, which are inherent to microcast structures, and a specific corrosive medium. While the surface of the metal is virtually un-attacked, stress corrosion cracking causes fine cracks to progress through the material. Typically, the range of environments causing cracking is small for a given material. Stainless steel, for example, will not crack in sulfuric, nitric or acetic acid, or pure water, but cracks in chloride and caustics [54].
5.4.5. Additional Support Material Strategies The four methods outlined in the previous paragraphs are showing a few basic strategies for the creation of support materials. Different support material strategies can be developed by combining and modifying the shown techniques. For example, for sprayed stainless steel structures sprayed copper can be used as a support and can be etched away, thus avoiding the possibility of destroying low melting support structures while using optimal (hot) parameter settings for the sprayed deposit. Another possibility is a modification of the strategy for low melting support structures with a sprayed thermal barrier for high temperature microcast metals. The thermal barrier can be sprayed from copper and is chemically removed from the part, after it has been melted out from the rest of the support structure. In contrast to the earlier strategy, parts created with this method will not contain any sprayed portions, but material selection is limited to corrosion resistant materials. Figure 5.8 depicts a scenario for the creation of a layer using a sprayed copper thermal barrier. The low melting support material is deposited onto the previous layer (a) and shaped to accommodate the sprayed copper shell (b). The copper barrier is sprayed on top of the low melting sup-
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a) previous layer and support
e) deposit part material
shape for external thermal barrier
b) create low melting support
f) shape part
g) remove low melting support
c) deposit Cu thermal barrier
d) shape thermal barrier
h) remove thermal barrier
Figure 5.8: Low Melting Support with Copper Thermal Barrier port structure (c). The copper is removed from the top cross-section of the part and the surfaces necessary for creating undercut features are shaped (d). Then, the material for the part is deposited (e) and the remaining features are shaped (f). After the last layer of the part is finished, the low melting support material can be melted away (g), to leave the part with the sprayed copper barriers attached to the areas with undercut features. In the final step, the sprayed copper is etched away from the part (h).
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6. Manufacturing Strategies Shape Deposition Manufacturing uses a variety of different deposition processes and shaping techniques to manufacture three-dimensional parts. During the geometric decomposition stage of the planning process a solid CAD model of the part is divided into layers and a sequence of manufacturable compacts (see Chapter 2.4). The basic strategies for manufacturing each compact described in Chapter 2.3 solve the geometric problem of creating compacts with a combination of undercut and non-undercut features. While non-undercut surfaces can be created through the deposition of the material followed by a shaping operation, undercut features are built using the surfaces of the adjacent compacts (i.e., the support material for one-material parts, pieces of the part made from a different material or the support material for multi-material parts). For undercut surfaces, the adjacent compacts are deposited first, and the surfaces, which are non-undercut surfaces on the adjacent compacts, can be shaped. The material for the compact containing undercuts is then simply deposited and conforms to the previously shaped surfaces. The strategies in Chapter 2.3 show the theoretical concept of the SDM process. They have to be expanded and adapted to meet the processing requirements of the different deposition and shaping technologies. Certain processes require different preparation or postprocessing steps, such as cleaning, stress relieving, surface roughening or other surface preparation techniques, which have to be included in the process descriptions used by the scheduling module during process planning, to derive the required sequence of processing operations necessary to manufacture a given compact. Several different (rule-based) process descriptions, which were used to build sample parts, are discussed later in this chapter.
6.1. Interlayer Penetration During the manufacturing of the compacts in a specific layer, several effects can cause the creation of layers with inaccurate interlayer boundaries. Penetration of dissimilar materials during the deposition processes and surface abrasion during surface roughening operations (e.g. grit blasting) can cause interlayer boundaries, that lie below the actual level and show a wavy surface. Inaccurate indexing on the shaping station and not accurately calibrated cutting tools (in case different tool sizes are used, or different tools are used for different materials) can cause interlayer boundaries, that can be too high or too low, and in case of indexing problems they can show an inclination in addition.
6.1.1. High Interlayer Boundary Areas, where the interlayer boundaries are higher than the expected level, do not create processing problems, but can cause geometric inaccuracies in the part, since subsequent shaping operations remove material from the previous layer. Figure 6.1 shows an example for two-compact
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layers. An indexing problem or inaccurately calibrated cutting tools cause the interlayer boundary after the material for the first compact has been deposited to be located at a higher level (a). Instead of shaping the compact correctly (b), the material is removed below the interface line (1). For cases, where the compact, or parts of it, are built on top of a dissimilar material the lower portions of the compact can contain segments of a different material than intended for the compact (2). Geometries with inclined contours can encounter discontinuities (3). The previous layer will be modified, and it will appear, that the top portion of its geometry is missing, and the total accumulated height of the part will be low (4). In case the indexing problem is corrected in the next step, the top of the compact or layer will be too low. top of previous layer too high
expected level
a)
expected shape 1
3
4 2
b)
Figure 6.1: Indexing Problem with High Interlayer Boundary
6.1.2. Low Interlayer Boundary Areas of the part, where interlayer boundaries are located lower than the expected level produce greater problems. Figure 6.2 shows two-compact layers, with the material of the first compact of the top layer deposited. Due to an indexing problem, abrasion of the top surface before the deposition process, or penetration during the deposition, the interlayer boundary between the previous layer, and the newly deposited material is lower than the required level (a). The shaping process for the compact now leaves a thin, intermediate layer on top of sections with dissimilar material of the previous layer (b). For a one-material part this means, that intermediate layers of part material, which are connected to the part, extend into the support structure (c). Problems with removing the support structure due to completely enclosed volumes can occur. Also, the fins, which are created by the material extending into the support structure, marking the layer interfaces on the part have to be removed from the part. Even more severe are intermediate layers of support material embedded inside the part itself (d). Removal of the support material causes gaps and voids in the part, and for areas with no contact between the individual compacts of the part material, the structure will delaminate. While indexing and tool inaccuracies can be avoided by careful tool calibration, usage of sharp tools, and measurement of the part position before every shaping operation, penetration during deposition due to required process parameters and abrasion of the surface due to surface roughening cannot be eliminated. For cases, were interlayer boundaries are lower than the expected level, it is inevitable to incorporate a compensation mechanism into the process strategies. Accumula-
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top surface too low
next deposit
expected shape
a) remaining material
b)
fins
c)
delamination
d)
Figure 6.2: Interlayer Penetration
tion of the effects of penetration and varying amounts of abrasion due to different materials have to be taken into account.
6.2. Compensation for Penetration and Surface Abrasion Two different compensation methods can be used, to avoid fin creation and delamination of the part due to thin, intermediate support material layers. While one method uses additional material on top of the exposed surfaces, which is selectively removed before the deposition, the second method builds a multi-compact layer from compacts of different height with step-wise shifted interlayer boundaries. Both methods, however, can potentially result in small deviations from the original shape of the part.
6.2.1. Selective Protection For typical multi-compact layers, only a portion of the surface needs to be exposed for the deposition of the material for the next compact. All other areas can be protected from abrasion and penetration by additional material above the top of the layer or the level of the interlayer interface. During the deposition process, additional material is deposited for each compact, and during the shaping process, the compacts are planed above the top of the layer. The additional height ∆h
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depends on the particular application. For surface roughening operations before depositing each compact with deposition processes that cover the whole part with material, or to compensate for penetration during the deposition process alone, the additional height needs to be the greater of the penetration or the abrasion depth of the most abrasive material. To compensate for abrasion with selective deposition, which covers only the necessary areas with material, enough height on top of the layer is necessary, to account for the accumulated abrasion caused by manufacturing each compact in the layer. When a new layer is started, the previous layer has an additional height of ∆h (Figure 6.3 a). The additional material is removed selectively from the area which is covered by the next compact (b). After surface roughening (c), if necessary, and due to the penetration during the deposition, only the additional, protective material shows contamination with dissimilar materials, and the interlayer boundaries are not affected (d). The top surface of the compact is planed above the actual height of the layer, to provide protection for depositing the remaining compacts of the layer, and for manufacturing the next layer (e). The compact is only shaped along the surfaces adjacent to the next compact. All other surfaces are shaped right before the compact adjacent to them is deposited. The process continues by selectively removing the protective material for the next compact. When all compacts are shaped and a layer is finished, the additional material above the top provides the required protection for the next layer (f). The first compact of the next layer can be started by selectively removing the protective material (b).
top surface of previous layer
protective layer
penetration ∆h
a)
d) selective removal
e)
b)
surface roughening
c)
g)
f)
protective layer exposed area shaped side of removed previous compact
h) exposed area small feature
previous compact
protective layer
material not removed top view
Figure 6.3: Selective Protection of the Interlayer Interface
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Unfortunately, due to the selective removal of the protective material, there are problems with some of the protective material remaining on top of areas prepared for deposition, or too much material being removed, exposing areas requiring protection (Figure 6.3 g and f). First, circular cutting tools with finite diameters used for the removal of the material leave some of the protective material on the inside of concave corners. Due to previous abrasion and penetration this material is already contaminated, or it could be a completely different material, if the underlying material is different from the current compact. Secondly, for accurately shaping some sides of previously deposited, adjacent compacts, or to expose thin features, such as thin walls, the cutting operation has to remove some material from areas which still have to be protected, and these areas will be subjected to abrasion and penetration.
6.2.2. Stepped Compensation protective layers ∆h ∆h · (Ci - 1)
a)
remove one protective layer
e) surface roughening and penetration
deposit
surface roughening and penetration
deposit
b) f) remove one protective layer
∆h · Ci+1
c) plane top
surface roughening and penetration
∆h · (Ci+1 - 1)
deposit
g) d)
Figure 6.4: Stepped Compensation for Interlayer Interface Penetration
Some of the problems caused by the selective removal of the protective material can be overcome by using multiple protective layers and removing the top portions of the protective material com-
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pletely each time a compact is deposited. Figure 6.4 shows an example for layers with three compacts. To start a new layer, additional material is required on the top of the previous layer with a thickness ∆h · (Ci -1), which is equal to the product of the number of compacts Ci of the new layer minus one and the height ∆h required for penetration or surface abrasion (a). After surface preparation and deposition of the material for the first compact (b), the compact is shaped. The areas of the previous layer, which are still exposed, are planed one thickness ∆h deeper than the previous level. The effects of surface abrasion and penetration from the deposition are removed, to avoid contamination of the top of the previous layer by unwanted material (c). On the top, protective layers are added to the compact, to allow for finishing the remaining compacts of the current layer, and to provide the protective material for the next layer. The thickness of the additional material on top of the layer is equal to the number of compacts Ci+1 in the next layer times the thickness of one protective layer ∆h. The sequence of deposition (d), shaping and removing the top portion of the protective material on top of the previous layer (e) is then continued for the remaining compacts of the layer. This will ensure, that no dissimilar materials are entrapped in form of thin, intermediate layers, causing fins or delamination. After the material for the last compact of a layer has been deposited (f) the top of the top protective layer is removed, to provide a clean interface for the next layer (g). The shape of the compacts in the zones of the additional height on top of the actual layers influences the geometric accuracy of the part. While the first compact of each layer uses the full amount of the initial, additional height and its shape from the previous layer, the amounts decline for each subsequent compact. Different approaches can be taken to shape the additional material on top of each compact. If the layer boundary is not necessary for a particular compact from the view of manufacturability of the geometry, i.e. the current compact and the following layer do not have a manufacturing conflict (undercut - non-undercut transition), then the compact, including the additional material, can be shaped according to the original compact geometry with shifted interlayer boundaries. A new compact model can be used for shaping, which extends ∆h · Ci+1 upwards into the next layer, and starts at a height of ∆h · (Ci - j -1) from the bottom of the current layer (with j being the number of the compact in the current layer). Figure 6.5 shows the stepwise shifted compacts for a multi-compact, multi-layer geometry. While this approach is completely accurate in terms of the geometry used for the shaping process, it cannot be used in case of manufacturing conflicts between a compact and the following layer. Alternative approaches, which cause small geometric inaccuracies, include extruding the top surface of the compacts straight upwards, continuing the slope of the top portion of the compact upwards, and using an approximated geometry from the original compact models for cases where undercut - non-undercut transitions occur. original interlayer boundaries
actual interface
Figure 6.5: Stepwise Shifted Interlayer Interfaces of Subsequent Compacts
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6.3. Sequence of Shaping Operations The shaping operations consist of several pathes with different characteristics, to remove additional material and to create the geometry of the current layer and compact. The principles of generating the pathes for shaping a compact have been briefly described in Section 2.4.9. This section describes the sequence in which the different cutting pathes are used in typical SDM manufacturing strategies. After the material for a compact has been deposited, the top surface of the compact is planed to obtain a flat surface at the top height hT of the compact. Additional height x·∆h for protection against abrasion and penetration has to be added to the planing height. If deposition processes are used, that do not selectively deposit the material, but cover the whole area of the part, the path for planing has to include all previously deposited compacts of the layer, to avoid excessive accumulation of material on the top of the layer. For contouring the shape of the compacts, a list of cutting tools with different diameters is specified. Typically, a sequence of cutters, where each subsequent tool has half of the diameter of its predecessor, to ensure that all the material is removed. If more sophisticated cutting strategies are employed, that can identify material left by a cutting operation (e.g. in small slots or concave corners), bigger spacing between the diameters of the cutting tools can be chosen or automatically calculated. To contour the side-band of a compact (or actually the current compact combined with the previous compacts of the layer) the cutting operations are divided into 2D contouring and a 3D contouring step. The 2D contouring shapes the compact with straight walls, to remove most of the material. The depth of the cut is given by the bottom height hB of the compact plus the height necessary for protecting the top surface of the previous layer for manufacturing the remaining compacts of the current layer. Since the shape of the combined compacts will not contain any undercuts the bottom cross-section can be used for deriving the cutter path. A ramp following the contour of the cross-section is added to the path, to avoid plunging straight into the material. The exact geometry of the side-band is then generated in a separate 3D contouring path, which uses the various 3D shaping strategies outlined in Section 2.4.9.2. The top and bottom heights of the geometry necessary to derive the 3D path are given by the top and bottom heights of the compact plus any additional height necessary for protection of the surfaces the top of the current and previous layer. Contouring the object starts with the biggest cutter in the list. A 2D-contouring path with the biggest available cutter is used to shape the outer boundaries of the combined compacts. To remove deposited material from the top of the previous layer in the areas, where subsequent compacts of the current layer have to be deposited, a mow-path with the biggest available cutter is used. The height at which the top surface of the previous layer is planed is the bottom height of the current layer hB plus the additional height from the remaining protective material x·∆h. After the mowpath, shaping continues with 2D-contouring with the remaining cutters, in the order of decreasing diameters. Finally, the three dimensional shape is created using the 3D-contouring path with the smallest cutter.
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To achieve better surface quality in cutting operations, usually rough cuts are performed a small distance away from the surface of the part. The actual surface of a part is created with a final cut. This can be easily accommodated by adding a small distance to the diameter of the cutters for generating the path offset, and the 2D-cut with the smallest cutter is repeated with the correct diameter of the cutter to shape the actual surface. To avoid abrasion of the three dimensionally shaped surfaces of the part during surface preparation for depositing the next compact of a layer, the surface preparation can be done after the rough 2D-cuts, and the final 2D-cut and the 3D-cut will produce surfaces with a high finish immediately before the deposition step.
6.4. Adaptive Compact Geometries The finite diameter of the cutting tools causes surfaces with a curvature larger than the curvature of the tool to be shaped inaccurately, i.e., concave corners will appear to be filleted. For multicompact layers, with concave boundaries between the individual compacts, this problem also results in “fillet-shaped” fins extending from the part, or “fillet-shaped” holes in the part surface, when the combined compacts (= union of current and previous compacts in the layer) have to be shaped. For certain geometries, this can be solved by using adaptive geometries for some of the compacts, where straight surface areas are adjacent to the parting surface of the compacts. The geometry of the compact used for shaping an undercut surface can then be kept slightly bigger (i.e., it can extend into the areas with the straight side-band) for deposition and shaping, until the shaping step before the material filling the concave boundary is deposited. The size of the initial compact is then reduced, to remove unwanted, “fillet-shaped” material. This concept of adaptively changing the geometry of certain compacts is illustrated in Figure 6.6. b) - f): top view inclined surface inclined surface
a) b)
desired geometry
d)
minimal support material
oversized support material
portion of compact removed
adaptive geometry c) not possible
filleted corners
non-adaptive e) geometry
“fillet-shaped” extensions
correct geometry
f) adaptive compact geometry
Figure 6.6: Adaptive Compact Geometry to Avoid “Fillet-Shaped” Extensions
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To create the part shown in Figure 6.6 a), support material is necessary to shape the undercut surface. The minimal size of the support structure (b) leaves no room to adapt the geometry, and results in filleted corners after depositing and shaping the part (c). If the support structure is extended along the straight walls (d) and the geometry of the initial support material compact is kept the same for cutting the part after the deposition of the next compact, “fillet-shaped” extensions will be the result (e). The “fillet-shaped” extensions can be avoided by adapting the geometry of the initial support material compact to a smaller size, and removing portions of the support material (f). Similarly, this concept can also be applied to shaping concave corners of straight side-band areas. In those cases, portions of the adjacent material can be deposited first, to avoid filleting.
6.5. Process Definitions Process definitions are a pre-established sequence of operations, which are necessary to manufacture a compact. The rules of the process definition are used by the process planner, to create the sequence of manufacturing instructions for the part (see Chapter 2.4.7). A variety of different definitions have been established and used for building parts. The operations include the deposition of materials, cutting operations and a variety of preparation and post-treatment procedures. Surface preparation techniques include grit-blasting to roughen surfaces and remove oxide layers, and preheating of the part before the deposition process. Shot-peening to relieve internal, thermal stresses in the deposited layers, and washing, to remove cutting oils or residues from the deposition process are typical post-processing procedures. Several process definitions are shown as examples in the following sections. New process definitions for different applications or using different processes, can be easily created by using and modifying elements from the shown definitions. The following is a list of variables, which are used in the process definitions: hTi ..................top height of layer i hBi...................bottom height of layer i di.....................thickness of layer i ∆h ...................height of protective material per compact L .....................number of layers Ci ....................number of compacts in layer i T .....................number of cutting tools specified to be used for shaping mati,j ...............material of compact j in layer i The following list describes the functions used for creating the process definitions: Plane(h, mat) ..........................plane the part at height h with the parameters for material mat and the tool specified for planing 2D-cut(h, mat, tool) ...............cut a 2D contour of the combined compacts at bottom
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height h with cutting tool tool 3D-cut(hB, hT, mat, tool) ........cut the 3D contour between height hB and hT with cutting tool tool Mow(h, mat) ..........................use mow-cut to plane the remaining surface of the previous layer at height h Deposit(d, mat) ......................deposit material mat to a minimum thickness of d Blast(mat)...............................grit-blast top surface with parameters for material mat Peen(d, mat) ...........................shot-peen material with a thickness d with parameters for material mat Preheat(mat)...........................preheat substrate for deposition of material mat Clean() ...................................clean the part Frame()...................................deposit frame material for powdered support structures Powder(h)...............................deposit powdered support structure up to a top level of h bond_mat(mat).......................returns the required bond material for material mat bond_thick(mat).....................returns the required thickness of the bond material for material mat is_blast(mat)...........................returns true if grit blasting is required for the deposition of material mat is_peen(mat)...........................returns true if shot-peening is required for material mat is_preheat(mat) ......................returns true if preheating is required for material mat For the implementation in the process planning and scheduling module the extraction of the geometric models used for the shaping and deposition operations has to be added to the process definitions. It has been omitted from the definitions in this chapter to emphasize on the processing aspects. The principle of generating the geometries has been shown in Chapter 2.4, 6.2 and 6.3.
6.5.1. Arbitrary 3D Structures The strategy for manufacturing three dimensionally shaped structures with no limitations in geometry uses the stepped compensation method shown in Section 6.2.2, and can be used for various deposition technologies. For materials requiring a preheated substrate, the part is preheated before the deposition process. For spraying operations, a thin layer of bond material, to improve the bonding conditions between the individual layers, is sprayed if required by a particular material. Grit-blasting is used to achieve the required surface roughness for the spray operations. The application of bond-materials and grit-blasting may not be required by micro-casting operations, and will be omitted during the evaluation of the process description by specifying the properties of micro-cast materials accordingly. The amount of materials penetration, the amount of material removed from a flat surface due to the grit-blasting operation, or essentially their sum, is assumed to the same for all materials (i.e., the value for the material with the biggest value is used for ∆h). The thickness of each layer must be big enough to allow all of the step-wise shifted compacts to
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be manufactured (hi > ∆h · (Ci+1 - Ci)). A base of support material is created underneath the part, to allow its removal from the substrate. The last compact of each layer completes the layer to provide a flat surface for the next layer. Therefore, the 2D and 3D contouring steps can be omitted from the last compact in each layer. Shot-peening is included in the process description, to compensate for thermal stresses created during the deposition process. If peening is required by a certain material, additional material with a thickness ∆hpeen is deposited. The layer is planed to provide a uniform surface for the peening process. The top portion of the layer (∆hpeen), which is severely deformed by the peening process is removed afterwards. For multi-compact parts, using the peening process for stress compensation creates a potential problem by peening earlier compacts of a layer several times. It is assumed, that this can be resolved, by depositing enough material to cover the whole area of the part, which will require additional peening and also protect previous compacts from the direct exposure to the peening, or by employing a sophisticated peening process, which is capable of affecting only selected areas of the structure. This process definition has been used to build the part shown in Chapter 8.1.1.2. Process Definition: Blast(substrate) Deposit(height_of_base, support) Plane(height_of_base, support) for layer i = 1 to layer i = L do1 for compact j = 1 to compact j = Ci - 1 if ((j == 1) and is_blast(mati,j)) Blast(mati,j) if is_peen(mati,j) d = di - ∆h · (Ci - j - 1) + ∆h · Ci+1 + ∆hpeen else d = di - ∆h · (Ci - j - 1) + ∆h · Ci+1 if bond_thick(mati,j) if is_preheat(bond_mat(mati,j)) preheat(bond_mat(mati,j)) Deposit(bond_thick(mati,j), bond_mat(mati,j)) Deposit(d - bond_thick(mati,j), mati,j) else if is_preheat(mati,j) preheat(mati,j) Deposit(d, mati,j) if is_peen(mati,j) Plane(hTi + ∆h · Ci+1 + ∆hpeen, mati,j) Peen(di - ∆h · (Ci - j - 1) + ∆h · Ci+1, mati,j) 1. See text for exceptions for the first and last layer.
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Plane(hTi + ∆h · Ci+1, mati,j) 2D-cut(hBi + ∆h · (Ci - j - 1), mati,j, 1) Mow(hBi + ∆h · (Ci - j - 1), mati,j, 1) for tool k = 2 to tool k = T 2D-cut(hBi + ∆h · (Ci - j - 1), mati,j, k) if is_blast(mati,j+1) Blast(mati,j+1) 3D-cut(hBi + ∆h · (Ci - j - 1), hTi + ∆h · Ci+1, mati,j, T) for compact Ci if not ((i == L) and (matL,CL == support)) if is_peen(mati,j) d = di + ∆h · Ci+1 + ∆hpeen else d = di + ∆h · Ci+1 if bond_thick(mati,Ci) if is_preheat(bond_mat(mati,Ci)) preheat(bond_mat(mati,Ci)) Deposit(bond_thick(mati,Ci), bond_mat(mati,Ci)) Deposit(d - bond_thick(mati,Ci), mati,Ci) else if is_preheat(mati,Ci) preheat(mati,Ci) Deposit(d, mati,Ci) if is_peen(mati,j) Plane(hTi + ∆h · Ci+1 + ∆hpeen, mati,j) Peen(di + ∆h · Ci+1, mati,Ci) Plane(hTi + ∆h · (Ci+1 - 1), mati,Ci) The base underneath the first layer does not have any additional, protective material, since stepped compensation is not possible because there are no compacts below the first layer of the part. To manufacture the first layer, the height informations in the described strategy have to be modified to use Ci - j = 0 for the deposition and Ci - j = 1 for the cutting routines. For manufacturing the last layer, the number of compacts in the following layer Ci+1 is set to one (CL+1 = 1). For manufacturing the last compact in a layer and the first compact of the following layer, the strategy can be optimized by combining the deposition steps, if both compacts are of the same material. The planing step for the last compact is then executed for the compact before the last compact in the layer, all other operations for the last compact are skipped. In the deposition step for the first compact of the following layer, the material necessary for both compacts will be deposited.
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6.5.2. 21/2D One-Material Structures Building sprayed, one-material structures in thin layers without three-dimensional shaping of the side-band, similar to the approach taken by SFF technologies, reduces the complexity of the strategy in the previous chapter. Only two compacts are necessary for each layer, the first compact being the material for the part, the second being the support material. A bondcoat layer is normally not used for the support material. In the following process description obj_mat is the material used for the part, and sup_mat is the support material. This process definition has been used to manufacture the examples shown in Chapter 8.1.1.1 and 8.1.2. Process Definition: Blast(substrate) Deposit(height_of_base, sup_mat) Plane(height_of_base, sup_mat) for layer i = 1 to layer i = L do if is_blast(obj_mat) Blast(obj_mat) if is_peen(obj_mat) d = di + ∆h + ∆hpeen else d = di + ∆h if bond_thick(obj_mat) if is_preheat(bond_mat(obj_mati)) preheat(bond_mat(obj_mat)) Deposit(bond_thick(obj_mat), bond_mat(obj_mat)) Deposit(d - bond_thick(obj_mat), obj_mat) else if is_preheat(obj_mat) preheat(obj_mat) Deposit(d, obj_mat) if is_peen(obj_mat) Plane(hTi + ∆h + ∆hpeen, obj_mat) Peen(di + ∆h, obj_mat) if (i < L) Plane(hTi + ∆h, obj_mat) else Plane(hTi, obj_mat) 2D-cut(hBi - ∆h, obj_mat, 1) Mow(hBi - ∆h, obj_mat, 1) for tool k = 2 to tool k = T 2D-cut(hBi - ∆h, obj_mat, k) if (i < L)
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if is_blast(sup_mat) Blast(sup_mat) if is_preheat(sup_mat) preheat(sup_mat) if is_peen(sup_mat) Deposit(di + ∆h + ∆hpeen, sup_mat) else Deposit(di + ∆h, sup_mat) if is_peen(sup_mat) Plane(hTi + ∆h + ∆hpeen, sup_mat) Peen(di + ∆h, sup_mat) Plane(hTi, sup_mat)
6.5.3. 3D Structures without Undercuts and Powder Support This strategy focuses on typical structures without undercuts, such as injection molding tools, which are built with the microcasting process, using a powdered support structure. There is no penetration resulting from the deposition of the support structure, while the penetration of the microcast material into the support structure to a depth of ∆hpen has to be removed. A solid base made from the material used for the part is first deposited, to allow for the finished structure to be cut off the substrate. The substrate is grit-blasted to remove surface oxides. The step for depositing the powder for the support structure includes all necessary processing steps, including the preparation of the top surface of the support material (see Chapter 5.4.3). This process definition has been used to manufacture a mild steel injection molding die shown in Chapter 8.2.3. Process Definition: Blast(substrate) Deposit(height_of_base, obj_mat) Plane(height_of_base, obj_mat) for layer i = 1 to layer i = L do if is_preheat(obj_mat) preheat(obj_mat) Deposit(di, obj_mat) Frame() Plane(hTi, obj_mat) if (hTi - ∆hpen > 0) if (hBi - ∆hpen > 0) hB = hBi - ∆hpen 2D-cut(hB, obj_mat, 1) Mow(hB, obj_mat, 1)
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for tool k = 2 to tool k = T 2D-cut(hB, obj_mat, k) 3D-cut(hB, hTi - ∆hpen, k) if (i < L) Powder(hTi)
6.5.4. Microcast, Non-Overhang Structures with Shrinkage Compensation Forces exerted on the underlying material by the shrinkage of microcast deposits during solidification and cooling causes the distortion of the previous layers. If the layer has been shaped, this results in distortion of the geometry of the layer (“Christmas Tree” effect). For structures with no undercut features, most of the distortion of the geometry of the part can be avoided, by depositing one layer, and then shaping the layer underneath (see Chapter 7.2.2.2). The distortion depth ∆hdis is the depth to which the underlying material is effected by the deposit is. The depth of penetration of the part material into the support material is ∆hpen. After a new layer has been deposited, the material underneath is shaped from the level of penetration into the support material up to the level, which is the distortion depth deeper than the top of the layer. The support material is then deposited and the top level of the support material is created at this same level. To create the trajectories for shaping the top level of the support material, the geometry of the part can be used. 2D- and mow-cutting of the part geometry will plane the top of the support material at a level lower than the top surface of the object, without removing any portions of the part. A shot-peening operation is used to relief the internal stresses of the part material only after the deposit. The irregular surface of the microcast material is planed before the peening process. An additional height is allowed for the peening process, which deforms the top portion of the layer to a depth ∆hpeen. The affected material is removed after the peening process. A cleaning cycle is included in this process definition, to clean the part after the cutting procedures. Process Definition: Blast(substrate) Deposit(height_of_base, base_mat) Plane(height_of_base, base_mat) for layer i = 1 to layer i = L do if is_preheat(obj_mat) preheat(obj_mat) Deposit(di + ∆hpeen, obj_mat) Plane(hTi + ∆hpeen, obj_mat) Peen(di + ∆hpeen, obj_mat) Plane(hTi, obj_mat)
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if (hTi - ∆hdis - ∆hpen > 0) if (hBi - ∆hdis - ∆hpen > 0) hB = hBi - ∆hdis - ∆hpen else hB = height_of_base 2D-cut(hB, obj_mat, 1) Mow(hB, obj_mat, 1) for tool k = 2 to tool k = T 2D-cut(hB, obj_mat, k) if (i < L) 3D-cut(hB, hTi - ∆hdis, k) else 3D-cut(hB, hTi, k) Clean() if (i < L) Deposit(hTi - ∆hdis - hB, sup_mat) 2D-cut(hTi - ∆hdis, sup_mat, 1) Mow(hTi - ∆hdis, sup_mat, 1) for tool k = 2 to tool k = T 2D-cut(hTi - ∆hdis, sup_mat, k) Clean()
6.5.5. Simplified Strategy for 3D Microcast Structures with Solid Support Microcast structures built from stainless steel with a copper support structure, do not require gritblasting to prepare the surface for the deposition process. With controlled droplet and substrate temperatures, penetration into dissimilar materials is very low. To simplify the manufacturing strategy and the planning process, compensation has been reduced to applying additional material only to the top surface during the buildup of a layer, and removing the material with a thickness of ∆h in the last planing step of the layer. Shot peening is included in the strategy. Additional material with a thickness of ∆hpeen is required and removed after the peening process. Peening operations in multi-compact layers present a potential problem, because of different compacts being subjected to a different number of peening operations and varying peening parameters for different materials. It is assumed, that this problem can be overcome by either covering the whole area of the part with deposit, so no previous compacts will be exposed to the peening process, and peening will be necessary for the whole area of the part, or having a sophisticated peening process, which is capable of selectively peening partial areas of the structure. This process definition has been used to manufacture the example shown in Chapter 8.2.4.
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Process Definition: Blast(substrate) Deposit(height_of_base, support) Plane(height_of_base, support) for layer i = 1 to layer i = L do for compact j = 1 to compact j = Ci - 1 if is_preheat(mati,j) preheat(mati,j) Deposit(di + ∆h + ∆hpeen, mati,j) Plane(hTi + ∆h + ∆hpeen, mati,j) Peen(di + ∆h + ∆hpeen, mati,j) Plane(hTi + ∆h, mati,j) 2D-cut(hBi, mati,j, 1) Mow(hBi, mati,j, 1) for tool k = 2 to tool k = T 2D-cut(hBi, mati,j, k) 3D-cut(hBi, hTi, mati,j, T) Clean() for compact Ci if not ((i == L) and (matL,CL == support)) if is_preheat(mati,Ci) preheat(mati,Ci) Deposit(di + ∆hpeen, mati,Ci) Plane(hTi + ∆hpeen, mati,Ci) Peen(di + ∆hpeen, mati,j) Plane(hTi + ∆hpeen, mati,Ci) Clean()
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7. Mechanical Properties of Shape Deposited Materials and Stress-Related Problems in Layered Forming Fully functional parts are characterized bygood surface appearance, accurate dimensions and superior material quality. For directly fabricating functional, metal shapes with the SDM process, the three dimensional shaping process has a big impact on the appearance and surface quality of the parts. It is the deposition process, however, which determines the material properties of the fabricated structures and has an important influence on the dimensional accuracy and surface quality. In the following sections, first the properties of materials typically used in SDM are examined and compared with the properties of conventional materials. Then, problems associated with layered, thermal deposition and subsequent shaping are discussed.
7.1. Material Properties of Sprayed Structures The properties of sprayed materials are typically between one and two orders of magnitude below the properties of conventionally manufactured materials. Only at great expense (shrouding, inert chambers, post-treatment) higher material properties can be achieved [55]. However, sprayed materials can be deposited onto a wide range of substrates, and have great importance as coatings for many applications. While entirely sprayed structures will never achieve the material properties required for fully functional parts, thermal spraying is a part of the SDM process and can be used together with other deposition processes. Several parts have been built with thermally sprayed deposits (see Chapter 8.1) of a laminated material consisting of zinc and nickel/aluminum layers.
7.1.1. Properties of a Sprayed Zinc - Nickel/Aluminum Laminate While the main material of the laminate is zinc, the 95% nickel, 5% aluminum alloy serves as a bondcoat to increase the adhesion between the individual zinc layers. The process parameters, which were used to deposit the materials and the exact material specifications are given in Chapter 8.1. The test samples were prepared according to a sequence to manufacture a layer, encountered in a regular strategy. A 76 µm thick layer of Ni/Al bondcoat is sprayed onto the machined and grit-blasted surface of the previous zinc layer. The zinc is then sprayed onto the unchanged surface of the Ni/Al, and the top surface of the zinc was planed. Nitrogen shrouding was used to lower the oxygen levels to 1000 ppm. The samples were created on a 91% tin, 9% zinc alloy (melting point 473 K) and could be melted off after fabrication. Figure 7.1 shows a cross-section through one of the samples used for tensile testing. The darker layers are Ni/Al, the bright zones are zinc. The nature of the dark lines along the interface zones between the zinc and Ni/Al is not quite clear, but could be an indication of insufficient bonding.
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500 µm
Figure 7.1: Sprayed Zinc - Nickel/Aluminum Laminate (18.5x)
7.1.1.1. Tensile Testing Ten layers of the Zn - Ni/Al laminate, each with a thickness of 0.25 mm, were deposited for the tensile test specimens. Three specimens were cut with a 50 mm long reduced section, which was 12.5 mm wide and 2.5 mm thick. The tensile testing was carried out at a speed of 2 mm/min. Figure 7.2 shows the stress - strain curves for the three specimen. 2 00
1 90 80
3
σxx [MPa]
70 60 50 40 30 20 10 0
εxx [%]
Figure 7.2: Stress - Strain Curves for Sprayed Zinc Nickel/Aluminum Laminate The average tensile strength of 92 MPa and the average elongation at break at 0.49% is low, which is typical for untreated, sprayed materials. For comparison, a tensile strength between 150 and 200 MPa was found for commercial, rolled zinc. For cast zinc the typical tensile strength is 230 - 330 MPa at an elongation of 5 - 14% [56].
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test #
tensile strength [MPa]
elongation at break [%]
1
92.5
0.65
2
102.0
0.51
3
82.4
0.32
Table 7.1: Tensile Strength and Elongation for Sprayed Zn - Ni/Al Laminate 7.1.1.2. Adhesion Testing To test the adhesion between the individual layers, round specimen with a diameter of 25 mm were manufactured. Each specimen consisted of one layer of zinc and one layer of the Ni/Al bondcoat. Two sets of test were conducted, one with the zinc deposited onto the as sprayed surface of the bondcoat, one with the bondcoat sprayed onto the planed and grit-blasted surface of a zinc layer. The test speed was 1.27 mm/min. test #
adhesion strength [MPa] Zn on Ni/Al
Ni/Al on Zn
1
12.1
12.2
2
11.1
10.1
3
11.1
12.9
4
10.2
10.5
5
6.9
11.9
6
9.1
11.1
Table 7.2: Tensile Strength and Elongation for Sprayed Zn - Ni/Al Laminate The adhesion tests have shown, that for both types the specimen delaminated inside the zinc layer. The breakage surface was extremely smooth, and suggests, that the delamination occurred at the intralayer interfaces, which are created by the multiple pathes of the spray torch to achieve the required thickness of the deposit. The average value of the adhesion strength was 10.1 MPa for the zinc sprayed onto the Ni/Al bondcoat, and 11.5 MPa for the Ni/Al bondcoat sprayed onto the grit-blasted zinc layer. The values are two orders of magnitude lower than the tensile strength of conventional, rolled or cast zinc, and one order of magnitude lower than the tensile strength of the laminate parallel to the layers. The low values of intralayer adhesion, which will be further decreased during the melting of the support material are the cause of delamination, which occurred on some of the test parts (see Chapter 8.1).
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7.1.2. Properties of Microcast Steels For testing the properties of microcast materials 127 x 152 x 13 mm test plates were created on top of a 19 mm thick steel substrate (1020 steel). 8 layers were necessary for the deposit, the parameters used for deposition can be found in Table 8.5. The wavy surface of each deposited layer was planed to a layer thickness of 1.6 mm. Subsize tensile test specimens were cut from the center of the microcast plate, to avoid deposit, which was influenced by the steel substrate, or which did not have the full thermal history due to missing layers on the top. The specimens had a reduced section with a length of 25 mm and a cross-section of 6.35 x 6.35 mm. Microcast layers are typically created by depositing droplets along parallel lines. Along the direction of the lines the droplets get deposited in quick succession and the previous droplet is still at a fairly high temperature when the next droplet impinges and adheres to the substrate and the previous droplet. Perpendicular to the direction of the deposition path, i.e when two lines of droplets are joined, some time has passed since the deposition of the previous line. The droplets of the previous lines have already cooled down, resulting in different remelting conditions for the newly deposited droplets. Anisotropic material properties can be expected. Tensile tests of microcast specimens were carried out parallel and perpendicular to the direction of the deposit (Figure 7.3).
longitudinal
perpendicular
Figure 7.3: Orientation of Microcast Deposit for Tensile Test Specimens
7.1.2.1. Microcast ER70S-6 Mild Steel Tensile tests for mircocast ER70S-6 mild steel were conducted on 9 longitudinal and 8 perpendicular specimens. The minimum, maximum and average values for the tensile strength, 0.2% offset yield and the elongation at break are shown in Table 7.3. longitudinal
perpendicular
tensile strength [MPa]
0.2% offset yield [MPa]
elongation [%]
tensile strength [MPa]
0.2% offset yield [MPa]
elongation [%]
min.
548.6
502.7
2.9
445.7
437.3
1.5
avg.
586.0
526.1
7.2
475.8
460.8
3.1
max.
614.8
560.6
11.7
506.9
473.6
5.1
Table 7.3: Tensile Test Result for ER70S-6 Mild Steel Specimens All specimens showed a ragged breakage surface (Figure 7.4), with indications, that the failure occurred partially at the interdroplet boundaries. Also, some boundaries between the individual
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layers and small voids in the deposit were visible. This seems to indicate, that interdroplet and interlayer bonding is not established to the full strength of the material. As expected, the interdroplet bonding is weaker in the direction perpendicular to the orientation of the deposit than in the longitudinal direction. Figure 7.5 shows the average of the stress-strain curves of the specimens in the longitudinal and perpendicular direction.
Figure 7.4: Typical Breakage Surface of Specimens
Overall, the tensile strength and 0.2% offset yield of the microcast material are close to the values specified for ER70S-6 weldments. The tensile strength for weldments is specified at 616.4 MPa, the yield point at 481.3 MPa. The average, measured tensile strength of the microcast material is 95% of the specified values in the longitudinal direction and 77% in the perpendicular direction. The elongation of welded material is specified at 27%. The rapid cooling conditions and the resulting martensite structure of the microcast deposit are responsible for the rather low elongation in the microcast specimens and also for the 0.2% offset yield, which is above the values specified for welded material. The hardness of the material was measured in several places and was consistent, i.e. Rc = 26. 00
longitudinal 00
perpendicular
σxx [MPa]
00
00
00
00
0
εxx [%]
Figure 7.5: Stress - Strain Curves for Microcast ER70S-6 Tensile Test Specimens
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7.1.2.2. Microcast 308 Stainless Steel For mircocast 308 stainless steel tensile tests were conducted on 8 longitudinal specimens. The minimum, maximum and average values for the tensile strength, 0.2% offset yield and the elongation at break are shown in Table 7.3. longitudinal tensile strength [MPa]
0.2% offset yield [MPa]
elongation [%]
min.
663.2
406.9
34.1
avg.
677.2
481.1
44.8
max.
685.7
499.5
58.4
Table 7.4: Tensile Test Result for 308 Stainless Steel Specimens The values measured for tensile strength, 0.2% offset yield and elongation of the microcast material are above the values specified for 308 stainless steel weldments. The tensile strength for weldments is specified at 579.2 MPa, the yield point at 399.9 MPa and the elongation at 35%. The average tensile strength measured in the longitudinal direction of the microcast material is 17% higher than the specified values, the yield point is 20% higher, and the elongation is 28% higher. The increased values for the material properties are a result of the rapid cooling conditions, to which the individual droplets are subjected. Rapid cooling of stainless steels typically results in an increased percentage of austenite and better ductility of the material.
7.2. Residual Stresses in Layered Deposits A problem inherent to thermal deposition technologies is the buildup of internal stresses in the deposits. The mechanical performance and the geometric tolerances of a part are affected by residual stresses. Thermally introduced residual stresses can be very high, and have the effect of preloading the structure, and cause failure at small loads during operation. Various mechanisms, contribute to the creation of residual stresses during the deposition process. Among them are mismatches of material properties for the creation of multi-material structures, or solid-state transforms, resulting in density changes (= volume changes) during the cooling process. The most important and severe effects, however, are experienced from temperature gradients, which are present during the deposition shortly after the solidification process.
7.2.1. Temperature Gradients After Solidification Microcast deposits are typically made onto a substrate significantly colder than the deposited droplets. Also, solidification and cooling of the droplets progresses from the substrate upwards, creating temperature gradients between the substrate and the solidifying droplet and along the
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solidification front. Typically, the substrate temperatures would range from room temperature to several hundred degrees above room temperature during the deposition process. During the cooling process, the liquid material shrinks due to thermal contraction. When one layer of material is deposited on top of another, colder layer of material, no stresses are induced until solidification. After the top material starts to solidify, it bonds to the material underneath, and the thermal contraction of the top material tries to shrink the top layer, while it is held back by the bottom layer. An equilibrium will be established, where the top layer will be in tension, and the bottom layer will be in compression. This effect will not only take place at the interface between the substrate and the newly applied layer, but is inherent throughout the whole new layer, since nonuniform cooling progresses from the bottom upwards. Depending on the nearby geometry, the structure will be affected in different ways.
7.2.2. Effects of Residual Stresses on the Artifact Three effects have been identified, which are caused by the buildup of residual stresses. The first two, warpage and an effect typical to the layered manufacturing approach, the “Christmas Tree” effect, are related to the geometry of the artifact and result in a loss of tolerance. In addition to the loss of accuracy, the third effect, debonding, results also in the immediate failure of the material during manufacturing. 7.2.2.1. Artifact Warping If the underlying layers of the growing artifact are not rigidly attached to a substrate, cooling and shrinkage of the newly applied layer will result in a curvature of the layers. Once the layers are in contact, and the top layer starts to solidify, shrinkage of the material on the top exerts compressive forces onto the bottom layer (Figure 7.6). Equivalently, the bottom layer exerts tension forces onto the top layer. An estimate for the resulting curvature of two unsupported, beam-shaped layers with the assumption of temperature independent material properties, uniform contraction of the solidifying material and by neglecting any plastic deformation, has been given in [57]. shrinkage hot (liquid) tension compression
cold (solid) solidification of top layer
warpage
Σ σstraight ≠ 0 Σ σwarp = 0
Figure 7.6: Artifact Warping due to Non-uniform Cooling
However, a closer analysis of this effect has to include temperature dependent material properties, to account for plastic deformation inside the solidifying layer and also in the upper areas of the underlying layer, which are significantly heated by the deposit, and to incorporate non-uniform cooling conditions. Also, the geometry of the underlying material and the layers itself have to be taken into account. For structures built from multiple layers, the interaction between the newly
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applied layer and the underlying material are repeated for each additional layer. The curvature of the artifact, and the residual stresses are increased with the growing height of the part. 7.2.2.2. “Christmas Tree” Effect1 This effect is caused by the shrinkage of previously shaped layers, when a new layer is deposited and results in geometric inaccuracies and a saw-tooth like appearance of the part surface. The formation of the “Christmas Tree” effect is illustrated in Figure 7.7.
support
previous layer
new deposit
shrunk deposit
compressed bottom layer Christmas Tree Structure
a)
b)
c)
Figure 7.7: Formation of the “Christmas Tree” Effect
The previous layer has been shaped and embedded in the support material (a). The material for the new layer is then deposited on top, and begins to cool and shrink. After solidification of the new layer, forces from the shrinkage are exerted onto the previous layer. The top portions of the previous layer have been heated by the deposit, causing a significant decrease in the elastic properties of the material. With the continued cooling process and the shrinkage of the top layer, the bottom layer is also subjected to shrinkage, mostly causing plastic deformation. The shrinkage of the bottom layer, which has been previously shaped, now causes the sides of the layer to lean inwards (b). The continuous shaping and application of layers results in a structure similar to the one shown in Figure 7.7 c) after the support material has been removed. Figure 7.8 shows an example of the “Christmas Tree” effect on a 13 x 13 mm, microcast, 308 stainless steel cube, with four 1.50 mm thick layers. All layers, including the top layer, were embedded in microcast copper as a support material. A distinct “Christmas Tree” structure can be seen on the three bottom layers, while the top layer only shows some melting from the support material deposit on the very top corner. It has been found, that the “Christmas Tree” effect is dependent on the local geometry and the direction of the dripped deposit. Also, increased substrate temperatures seem to favor the effect, due to decreased elastic properties of the underlying material. Several measures can be taken to avoid or improve the “Christmas Tree” surface structure. Shotpeening of the newly applied layer can be used, to introduce stresses and plastic deformation counter-acting the effects caused by the shrinkage of the deposit. Since both the “Christmas Tree” effect and shot-peening cause plastic deformation, it is difficult, if not impossible, to bring the surface back to the specified geometry. An improvement of the surface appearance, however, will be possible. Another method, which requires a detailed thermal and mechanical analysis of the structure, is to compensate for the “Christmas Tree” structure by modifying the side band of each 1. The name “Christmas Tree” effect was given to the appearance of affected structures, which is similar to a popular way of drawing christmas trees.
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1 mm
Figure 7.8: “Christmas Tree” Effect layer, to allow for the shrinkage caused by newly applied layers. The most promising solution seems to a strategy, were the support material is kept at a lower level than the upper level of the growing part. This strategy, which can be used for parts without undercuts (limited use for undercut features might be possible) is illustrated in Figure 7.9. The previous layer has been shaped, and the support material is kept at a lower level (a). When the new material is deposited, it also covers the sides of the previous layers and adds material to compensate for the shrinkage of the layer (b). Both layers can now be shaped (c), and the process continues after embedding the previous layer into support material (d). previous layer
new deposit
shrunk deposit
shrinkage added material
support
a)
b)
c)
d)
Figure 7.9: Compensation Strategy to Avoid “Christmas Tree Effect
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7.2.2.3. Debonding Incremental stress buildup through the continued deposition of new layers can lead to stress concentrations, which are high enough to cause interfacial cracks. The cracks can start from the edges of the artifact and will propagate into the center. Eventually the cracks may propagate throughout the entire artifact, separating it into two pieces. To avoid delamination it is necessary, to counteract the incremental stress buildup, e.g. with shot-peening. Structures without underlying rigid support, which have to ability to warp, are less likely to show delamination. Due to the warpage, some of the internal stresses are relieved. In rigidly supported structures, such as layers, which are deposited onto thick, solid substrates, warpage is not possible, or only possible to a very small extent. Therefore, the internal stresses are not reduced and continuous stress buildup will lead to debonding at weak interfacial boundaries. Figure 7.10 shows two examples of interlayer delamination and layer warpage caused by incremental stress buildup. In both cases, the layer on the bottom was attached to a rigid 19 mm thick steel substrate. The residual stresses of the next two layers caused delamination between the first and the second layer. The second and the third layer show significant warpage. Further, it can be concluded, that the delamination occurred after the shaping operation, since the top surface of the third layer is also bent upwards. For the picture on the left, the delamination started at the corner and propagated into the material along the layer boundary. In the right picture, breakup of the two layers occurred along the edge, but away from the corners. Additional residual stresses from the structure above the third layer has contributed to the delamination and the warpage of the second and third layer.
1 mm
1 mm
Figure 7.10: Delamination and Warpage due to Buildup of Residual Stresses
7.2.3. Counteracting Internal Stress Buildup Incremental stress buildup during thermal deposition leads to warpage and the “Christmas Tree” effect, which result in loss of the geometric tolerances of the structure. Continued stress buildup
136
ultimately leads to delamination and failure of the material. It is inevitable to eliminate or compensate for these stresses. However, it is difficult to accurately counteract any geometric effects, such as warpage or the “Christmas Tree” structure, after they occurred, and some of the stresses were relieved due to the geometric deformation, i.e. the geometry of the structure cannot be accurately be brought back to its specifications. Methods for correcting the “Christmas Tree” effect have been shortly described in Section 7.2.2.2. To eliminate warpage, first it has to be ensured, that by the deposition of only one single layer no significant warpage will occur. That means, that each deposited layer has to be thin enough (i.e., small amount of residual stress) and the underlying substrate has to be rigid enough to prevent deformation. Warpage caused by continued incremental buildup of stresses can then be avoided by compensating for the stresses of each layer. One method for stress compensation is shot-peening [58]. Spherical particles are propelled at high speed onto the surface of the layer, and cause a thin, superficial layer of compressive stresses (through indentations), which try to expand the material on the top surface. Underneath the thin layer of compression, the expansion induces tensile stresses in the region underneath. Initial experiments have shown, that the tensile stresses are of significant magnitude, and result in plastic deformation of the region underneath the superficially deformed layer. The impacted zone can potentially be removed1, to leave a layer consisting of residual tensile stresses in the top portion, and the original residual compressive stresses from the thermal deposition in the bottom region. With the correct amount of peening, the tensile and compressive stresses could cancel each other out, and eliminate the incremental stress buildup leading to warpage and ultimately to delamination. Figure 7.11 tries to illustrate the effect. The thermal deposition leaves a layer with compressive stresses, and warpage is prevented by a rigid substrate (a). After the peening operation, the deposited layer shows tensile stresses in the top region, which were induced by the peening operation, and compressive stresses in the bottom portion, which remain from the non-uniform cooling process during deposition (b). Incremental buildup of layers and intermediate shot-peening will lead to structures with alternating layers of tensile and compressive stresses (c). While those stresses cancel each other to avoid deformation, they still present a substantial preload of the structure and can lead to early part failure under small loads. A possible method to relief the alternating stress layers is to remove the part from its original support structure, which was used for the building process and embed it in an extremely rigid (ceramic) support structure. The part can then be stress-relief heat treated [59] and the rigid support structure will prevent any deformation. The amount of residual stresses, which are relieved, are correlated to the duration and temperature of the treatments, and are based on typical stressrelaxation behavior, where the material undergoes microscopic (and even macroscopic) creep at the stress-relief temperature. Typical stress-relief temperatures for low-alloy ferritic steels are between 868 and 948 K and between 1173 and 1338 K for high-alloy steels. Slow and uniform reduction of the temperature of the part can then prevent the renewed buildup of residual stresses and warpage during the cooling process. For certain materials, such as austenitic stainless steel however, stress relief heat treating at higher temperatures (753 to 1198 K) and slow cooling can leave the material in an undesired metallurgical state. 1. Workhardening due to the peening process can be a potential problem.
137
solidification of layer
rigid substrate
compression (layer tries to expand)
tension (layer tries to shrink)
shrinkage
counteract shrinkage forces
a)
before peening
b)
after peening
c)
Σσ=0
Figure 7.11: Compensation of Residual Stresses through Shot-Peening
Extensive test have currently been started at the Shape Deposition Laboratory at Carnegie Mellon University to explore the issues of shot-peening and incorporate the process into the SDM environment and manufacturing strategies. With respect to multi-compact structures, a more selective peening process, such as a quickly oscillating hammer with a small spherical head and servo positioning control, might have to be developed, to allow accurate control and prevent overpeening of previous compacts.
138
8. Example SDM Parts To demonstrate the abilities of the Shape Deposition Process and the proposed strategies several test parts have been built. The examples show, that SDM is a feasible alternative to current rapid manufacturing systems. It is demonstrated, that SDM is capable to directly manufacture three dimensional, metal prototypes with arbitrary shapes. For tooling components, such as injection molds, SDM has shown the ability (with somefuture improvements) to produce fully functional parts.
8.1. Sprayed Parts The parts built with thermal spray deposition where built from zinc, with a low-melting tin-zinc alloy as support material, and a nickel-aluminum bond coating. Several other materials have been tried with a tin-zinc support structure, but did not lead to satisfying results. The required spray parameters were either too hot and melted the support material, or the material properties produced with low power settings lead to tough, uncutable coatings, debonding layers and brittle material, which chipped away during cutting operations. For the test parts shown below, zinc powder (Eutectic + Castolin CPW1960) was plasma sprayed with a “Plasma Technik F4-MB” plasma spray torch. A 76 µm thick bondcoat layer of a 95% nickel - 5% aluminum powder (Metco 450P) was used to enhance the interlayer bonding strength. A 6 mm nozzle was used in the plasma torch with a 6 mm gauge point for the powder ports. The zinc was injected into the plasma at an angle of 105°, the bondcoat at 90°. A shroud providing an inert shield of nitrogen was used with the plasma torch. With a nitrogen flow rate of 1.42 m3/min the oxygen content measured at the substrate was 1000 ppm. The support material was a 91% tin, 9% zinc alloy (“392”) sprayed with a “Miller BP-400” wire-arc spraying torch with nitrogen as the atomizing gas (wire diameter: 2.38 mm, atomizing gas pressure: 552 kPa). The melting point of the support material is 473 K (392 °F). Cross-hatch pathes, alternating in the x and y direction were used for the deposition. The number of pathes necessary was calculated from the required thickness of the layer. Before each spraying operation, the surfaces of the zinc as well as the support material were grit-blasted with aluminum-oxide particles. An abrasion depth ∆h of 0.127 mm was used to compensate for the material removed by the grit-blasting process. One serpentine path was used in each direction (x and y). Before spraying the bondcoat, the part surface was preheated by traversing the plasma torch several times over the part without material flow. For the cutting operations, cobalt coated HSS endmills and air coolant were used. Due to oxidation and the sprayed nature of the material, cutting operations were extremely abrasive on the cutter edges. Frequent replacement of the cutters was necessary, to avoid chipping and delamination of the layers during cutting operations. The parameters used for creating the example zinc parts are shown in Table 8.1 - 8.4. To remove the support material, the part is put into a furnace and heated to 483 K with a heating rate of 10°/min. While most of the support material will run off the part, remains of the melted tin-
139
zinc alloy can be removed with a soft brush. Extreme care has to be taken during brushing, to avoid delamination of the weak zinc layers while the part is at elevated temperatures. Once the part has cooled down, the surface can be cleaned from oxides with glass-beating. While the intralayer bond-strength of the plasma-sprayed zinc layers is already extremely weak, it gets further reduced when the part is heated to remove the support material. Several of the zinc nickel/aluminum parts showed delamination, which occurred during the meltout process, in areas with increased stress concentration. The delamination occurred on the intralayer interfaces, which are created when several pathes of the spray torch are necessary to achieve the desired layer thickness. The bad bonding conditions are most likely caused by oxidation of the intralayer surfaces with the remaining oxygen in the shielding atmosphere. While the combination of zinc and a tinzinc alloy as the support material has shown the feasibility of the concepts of SDM, it is not usable for the reliable production of parts. It seems possible, that with necessary effort it is possible to identify other material combinations, which are suitable for the thermal spraying approach. However, the materials created through thermal spraying will never produce fully functional parts. Due to limited resources and time constraints the search for suitable materials for thermally spraying SDM parts was abandoned for this work and the focus was given to microcast deposits, which have shown the ability to produce high quality materials. Amps [A]
Volts [V]
Plasma Gas (Ar) [l/min]
Plasma Gas (H2) [l/min]
Carrier Gas (Ar) [l/min]
Disk [%]
Stirrer [%]
path spacing [mm]
standoff [mm]
speed [mm/s]
Deposit [µm/path]
preheat
200
~63.7
47.4
2.0
0
0
0
5.0
152
254
0
Zinc
275
~55.0
50.0
1.0
3.0
24.0
50.0
5.0
152
254
64
450P
400
~70.1
47.4
8.0
3.5
19.0
50.0
5.0
152
254
38
Table 8.1: Plasma Spray Parameters for Zinc and Metco 450P Powder
“392”
Amps [A]
Volts [V]
Wire Feed [mm/s]
Atomizing Gas (N2) [kPa]
path spacing [mm]
standoff [mm]
traj. speed [mm/s]
Deposit [µm/path]
~60
26.8
53.3
552
20.0
230
254
35
Table 8.2: Arc Spray Parameters for “392” Tin-Zinc Alloy
Zn/392
Nozzle ∅ [mm]
Grit Size [mm]
Air Pressure [kPa]
path spacing [mm]
standoff [mm]
speed [mm/s]
6.35
0.35 - 1.00
207
25.4
152
30
Table 8.3: Grit-Blasting Parameters for Zinc and “392” Tin-Zinc Alloy
8.1.1. Zinc Laminate Test Parts 8.1.1.1. Thin, L-Shaped Wall Figure 8.1 shows a small test part sprayed from zinc - nickel/aluminum laminated material. The
140
Cutters
Zinc
“392”
Diameter [mm]
Cutter Type / Flutes
Spindle Speed [rpm]
Feedrate [mm/min]
Spindle Speed [rpm]
Feedrate [mm/min]
19.05
HSS / 4
800
254
800
254
12.70
HSS / 4
1200
127
1200
254
6.35
HSS / 4
1600
190
1600
217
3.18
HSS / 4
3200
254
3200
381
Table 8.4: Cutting Parameters for Zinc and “392” Tin-Zinc Alloy artifact was built with the 21/2D strategy shown in Chapter 6.5.2 without 3D contouring. A layer thickness of 0.25 mm was used. The part shows a thin, L-shaped wall, which is sitting on top of a base. The base is 32 x 25 mm in size, and 1 mm high. The wall is 5 mm high and has a thickness of 1 mm. The slots in the wall have a height of 1 mm. The inclined ends of the wall show the stairstep texture typical for SFF parts. The surface texture on the straight sides of the wall clearly shows the individual layers. Small remains of the support material are visible on the base and on the edge from the longer wall to the base. Through extremely careful removal of the support material while the part was heated to 483 K delamination could be avoided.
Figure 8.1: Sprayed Zinc - Nickel/Aluminum Test Part 8.1.1.2. Interlocking Frames with 3D-Shaped Surfaces The part shown in Figure 8.2 is another zinc - nickel/aluminum laminate showing two interlocking frames, which are inclined under angels of 45°. The artifact was built with full 3D contouring and adaptive layer thickness. The size of the base is 50.8 x 38.1 x 6.4 mm, the beams of the frames are 6.4 x 6.4 mm. The overall height of the part is 24.6 mm. The layer thickness ranged from 1.0 to 2.5 mm. This part shows the ability of SDM to avoid the stair-step texture usually encountered on inclined surfaces in conventional SFF processes. On the edges, some chipping of the brittle zinc structure is visible. Underneath the inclined beams of the frames, marks are visible on the base of the part, and the surface of the base is slightly elevated. The support material, which was
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used to shape the undercut features of the beams, protected the portions of the base underneath the beams, while the remaining surface of the base was abraded during grit-blasting. Due to the extremely weak material strength of the zinc layers at elevated temperatures, one of the frames delaminated from the base during the meltout process. The frame has been glued to the part, to demonstrate the as-built geometry.
Figure 8.2: Sprayed Zinc - Nickel/Aluminum Test Part with 3D Shaped, Inclined Surfaces
8.1.2. IMS-T1 Test Part The IMS-T1 part was one of the test parts used in a world wide assessment of rapid prototyping technologies [60]. A half-scale version has been built from the sprayed zinc - nickel/aluminum laminate with a total of 149 layers. The 21/2D strategy of Chapter 6.5.2 was used with a layer thickness of 0.25 mm. A cross-section through typical layers of the laminate, and tensile and adhesion test data are presented in Chapter 7.1.1. The overall dimensions of the part are 75 x 50 x 37 mm. Figure 8.3 shows the part during the building process. The current layer of the part, which has been shaped in the previous step, is visible slightly raised from the surrounding support material. the surface has been grit-blasted and the support material is deposited in the next step, to embed the layer. Figure 8.4 shows the finished part still embedded in the support structure on the substrate and pallet, which is held by the transfer robot. Figure 8.5 shows the final part after the tin-zinc support material has been melted away. The stair-step texture caused by the 21/2D manufacturing strategy is clearly visible on the inclined surfaces of the part. With this part is has been demonstrated, that SDM can handle intricate geometries comparable to SFF processes. However, several small defects are visible on the part. As can be seen from the picture in Figure 8.5, the staired geometry at the back of the part delaminated at the end of the meltout process. The meltout process lasted for about 1.5 hours, and the part was frequently moved out of the furnace to manually brush away adhering support material. Stresses caused by the thermal cycling, and support
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material creeping into small defects between the layers finally lead to the delamination. Also, the top portion of the cylindrical/conical shaped feature in the back left corner delaminated during the removal of the support material. During the meltout process, some zinc-oxide formed at the surface of the part. Glass-beating was used to remove most of the oxide. The small, shiny dots, which can be seen on the surface of the part, are residual support material.
Figure 8.3: Intermediate Layer of the IMS-T1Test Part
Figure 8.4: IMS-T1 Test Part Embedded in Support Material
143
144
Figure 8.5: IMS-T1 Test Part Sprayed from a Zinc - Nickel/Aluminum Laminate
8.2. Microcast Parts The metallurgical bonding of microcast droplets results in parts with strong interlayer bonds and superior material properties than experienced in the sprayed examples. Several different materials and strategies have been developed. The following examples show parts created with low melting support materials, protected by thermally sprayed layers, parts created with a bound metal powder support structure, and stainless steel parts created with a microcast copper support material.
8.2.1. Steel Parts with Sprayed Support 8.2.1.1. Injection Molding Die with Internal, Serpentine Cooling Channel The part shown in Figure 8.6 is a mild steel model of an injection molding die with internal cooling channels, to create plates with the raised letters “CMU”. While the outer shape of the tool (76.2 x 63.5 x 25.4) could have been easily built with conventional manufacturing techniques, the internal features of the tool would require specialized casting techniques, and show the potential of the SDM process with microcast deposits. The two rectangular shaped holes (6.4 x 5.1 mm) on the front of the die are connected by a serpentine cooling channel.
Figure 8.6: Injection Molding Die with Internal, Serpentine Cooling Channel
The part was built by first depositing the mild steel (ER70S-6) for the bottom half of the die through deposit MIG-welding. The serpentine cooling channel and a recess for the sprayed material were cut into the material. The cooling channel was filled with a 91% tin, 9% zinc alloy (“392”) (Figure 8.7 a). A 1 mm thick layer of 410 stainless steel was wire-arc sprayed (with nitrogen atomizing gas), and planed to level with the top surface of the bottom half of the die (b). Two layers of microcast ER70S-6 with an approximate height of 5 mm where put on top, the remaining material was deposited by welding (c). During the first layer of microcast deposit, some of the “392” support material melted and flowed out of the cooling channels, without disturbing the deposit. Finally, the side and top of the die was cut, and the tin-zinc support was melted out.
145
low melting support
a)
fill cooling channel with low melting support material
sprayed layer
b)
spray protective layer
c)
microcast and weld top half of die
Figure 8.7: Creation of Internal, Serpentine Cooling Channel
8.2.1.2. Steel Cube with Internal Cavity and Embedded Sphere A similar artifact, a hollow, microcast ER70S-6 mild steel cube (38.1 mm) with an inserted steel sphere is shown in Figure 8.8. This part shows the ability of SDM and microcasting, to produce intricate, hollow objects, with embedded components. To create the cube, material was deposited through microcasting up to the ceiling of the cube. The internal cavity of the cube was machined and the steel sphere was inserted (Figure 8.9 a). The cavity was filled with the tin-zinc alloy (“392”) and a 1mm thick layer of 410 stainless steel was wire-arc sprayed (with nitrogen atomizing gas) to protect the low melting alloy (b). The material for the ceiling was then microcast on top of the sprayed layer (c), and the outside of the cube was shaped. To remove the support material, holes were drilled into the top and bottom of the cube, and the tin-zinc alloy was melted out.
Figure 8.8: Microcast Steel Cube with Internal Cavity and Embedded Sphere The injection molding die and the cube with the embedded sphere where built with the early version of the microcasting process, using a MIG-welding torch and a rotating, water-cooled copper electrode (see Chapter 4.3.1) to create the droplets. Nitrogen shrouding was not used with this particular setup. The deposited material shows a mainly martensitic structure and oxidation, resulting in unfavorable conditions for the cutting process. The instability of this microcasting
146
sprayed layer
low melting support
a)
cube with inserted sphere
b)
cube filled with support material and sprayed layer
c)
microcast top of the cube
Figure 8.9: Creation of Internal Cavities with Embedded Objects setup also lead to the creation of occasional voids in the deposit. Droplet temperatures and sizes have not been measured.
8.2.2. Copper-Steel Multimaterial Parts The multi-material capabilities of the SDM process have been demonstrated with the manufacture of parts consisting of steel and copper. 8.2.2.1. Laminated Tube with Copper and Steel Layers The laminated tube shown in Figure 8.10 is built from alternating, microcast layers of mild steel and copper with a thickness of 1.9 mm. The tube has a height of 19 mm, a diameter of 28 mm and a wall thickness of 2.5 mm. The material was deposited with the MIG-microcaster, which was not capable of generating droplets close to the evaporation point of the material. While the copper layers adhere well to the steel, the bond at the interfaces, where steel was deposited onto the copper are weak, and tiny voids are visible in a few spots. This was mainly caused by insufficient temperatures of the steel droplets. The tube was built without any support material. Each of the copper and steel layers was first deposited on top of the previous ones and planed, and the tube was shaped after all layers have been deposited. 8.2.2.2. Copper Wheel on a Steel Axle The artifact displayed in Figure 8.11 not only shows the possible creation of multi material structures, but also the feasibility of manufacturing multi-part assemblies. Support material layers can be used, to keep the individual parts of the assembly in position. The copper wheel, which is mounted on a mild steel (ER70S-6) axle was built from microcast material with a tin-zinc support structure and layers of sprayed material. The copper wheel has a diameter of 38.0 mm and a thickness of 7.6 mm. The diameter of the axle is 12.6 mm, the hub diameter 19 mm, and the clearance between the steel axle and the copper wheel in a concentric position is 1.0 mm. To manufacture the assembly (see Figure 8.12), first the bottom hub and the axle were deposited, shaped (a) and embedded in support material. The support material was shaped, to provide the surfaces for creating the bottom and inside surface of the copper wheel (b). A 1 mm thick layer of copper was ther-
147
Figure 8.10: Copper - Steel Laminated Tube mally sprayed for the bottom and inside wall of the wheel (c). For the remaining portions of the wheel, copper was deposited with the MIG-microcaster (d). After the outer and top surface of the wheel were created, the wheel was embedded in support material (e). The upper hub of the axle was then created by wire-arc spraying a 1 mm thick layer of steel onto the planed support material (f), microcasting the remaining material and shaping the top of the axle (g). After removal of the low melting support material, the wheel was able to run on the axle (h).
Figure 8.11: Copper Wheel and Steel Axle
Using the unshrouded MIG-microcaster, with insufficient control over the droplet temperatures led to the appearance of voids in the deposited material. The steel portions show voids created by
148
a)
bottom hub and steel axle
e)
deposit support material
b)
deposit support material
f)
spray steel layer
c)
g)
spray copper layer
microcast top steel hub
d)
microcast copper wheel
h)
melt support material
Figure 8.12: Manufacturing Steps for Copper Wheel - Steel Axle Assembly undercut shapes of the deposited droplets, which were caused by insufficient droplet temperatures and unfavorable wetting conditions due to droplet oxidation. The voids in the copper deposit were mainly caused by uncontrollable droplet trajectories.
8.2.3. Steel Mold Half A powder support structure and the strategy from Chapter 6.5.3 were used to create a mold half for an automotive part. The mold created with the SDM process is shown on the left hand side of Figure 8.15. The mold-half on the right, was conventionally manufactured and assembled from several different pieces. For the SDM mold-half ER70S-6 mild steel with a wire diameter of 0.9 mm was deposited with a plasma microcaster. The droplets had a diameter of 3.6 mm and a temperature of 2596 K. The microcaster settings and the parameters for the deposition trajectories are shown in Table 8.5. The support material was mild steel powder (Plasmalloy AI-P230) mixed with 10% (weight) sodium silicate, which was dehydrated and bound with CO2 (see Chapter 5.4.3). To hold the support material in place, a 120 x 120 mm frame was deposited during the microcasting process. Figure 8.13 shows a top view of the injection mold during the manufacturing process, after the support material has been deposited. The grey area between the frame and the cross-section of the part is the support material. In Figure 8.14 the material for the next layer has been deposited. The maximum penetration of the steel droplets into the cured powder was 1.25 mm. 29 layers with an adaptive thickness between 1.15 and 1.49 mm were necessary to build the structure. A 76.2 mm cutter with 6 carbide inserts was used for the planing operation, titanium-corbonitrate coated HSS endmills with 12.70, 6.35 and 3.18 mm diameter were used for cutting the contours of the mold. The channels in the mold cavity were cut with 1.59 mm endmills. The cutting parameters are listed in Table 8.6. The mold has a diameter of 88 mm (top) and an overall height of 354 mm. The channels in the mold cavity are between 2.75 and 5.24 mm deep and 2.54 mm wide. Inclined surfaces had to be shaped inside the channels with a 3D cutting strategy.
149
Figure 8.13: Injection Molding Tool with Powder Support Structure before the Deposit of the Next Layer
Figure 8.14: Injection Molding Tool and Powder Support Structure with Newly Deposited Layer
No strategies to reduce or compensate for the effects of residual thermal stresses were used to create this part. As a result, a “Christmas Tree” surface structure, with a maximum surface deviation of 0.1 mm was measured on the outside of the injection molding tool. The step, which can be seen at approximately half the height of the mold, was caused during a planing operation with a 76 mm carbide insert cutter. Due to insufficient horsepower the spindle of the milling machine stalled and forces exerted by the x-y feed caused a misalignment between
150
151
Microcast Mold Manufactured with SDM
Conventionally Manufactured Mold
Figure 8.15: Injection Mold-Half for Automotive Part
the substrate and the pallet receiver. Amps [A]
Volts [V]
Wire Feed Rate [cm/s]
Plasma/ Shield Gas Ar [l/min]
Shroud Gas N2 [l/s]
drop rate [drops/ s]
drop diam. [mm]
path spacing [mm]
standoff [mm]
speed [mm/ min]
max. height [mm]
ER70S-6
68
~29.2
5.33
1.15/11.5
7.87
1.20
3.61
4.57
76
127
1.55
308 Stainless
70
~30.3
5.84
1.15/11.5
7.87
1.68
3.37
4.57
76
152
1.65
Copper
58
~26.9
5.84
1.15/11.5
7.86
2.11
3.19
4.32
76
203
1.52
Table 8.5: Microcast Parameters for Steel and Copper Cutters
Mild Steel ER70S-6
308 Stainless Steel
Copper
Diameter [mm]
Cutter Type / Flutes
Spindle Speed [rpm]
Feedrate [mm/min]
Spindle Speed [rpm]
Feedrate [mm/min]
Spindle Speed [rpm]
Feedrate [mm/min]
76.20
Carbide / 6
450
254
350
220
500
260
50.80
Carbide / 4
800
355
800
244
1200
304
12.70
HSS / 2
800
127
350
64
611
91
6.35
HSS / 2
1600
228
700
114
1222
185
3.18
HSS / 2
3200
304
1500
203
2445
356
1.59
HSS / 2
7000
178
7000
178
7000
178
Table 8.6: Cutting Parameters for Microcast Steels and Copper
8.2.4. IMS-T2 To demonstrate the capabilities of the SDM process to create three dimensional structures with arbitrary geometry, a half-scale version of the second test part, IMS-T2, of a recent world-wide assessment of rapid prototyping technologies [60] has been built from microcast stainless steel. Microcast copper was used as a support material. The stainless steel droplets were deposited at a temperature of 2555 K, the copper droplets had a temperature of 2588 K. The big heat sink created by the big amount of copper support material required preheating the part to approximately 453 K before the microcast deposition. The part was created from 29 layers with a thickness between 1.0 and 1.7 mm and a total of 79 compacts were necessary to create the geometry. The simplified strategy for microcast 3D structures from Chapter 6.5.5 was used with an additional material height ∆h of 0.127 mm for each compact, which was removed upon completion of the layer. Shot-peening was omitted to build this part. The overall dimensions of the part are 75 x 50 x 42 mm. The IMS-T2 part shows some intricate features, such as the pocket on the right side of the part, which has a wall thickness of 1.59 mm, and the wall on the left side with a thickness of 1.14 mm at a height of 13.65 mm. Portions of the arc in the back of the part have cross-sections with an area as small as 5.04 mm2. Figure 8.16 through 8.18 show the part during the building process. In Figure 8.16 a new layer of stainless steel has been deposited. Figure 8.17 shows the part after shaping the stainless steel. In
152
Figure 8.18 the part is completed, but still enclosed in the copper support structure.
Figure 8.16: IMS-T2 Test Part during Buildup with Newly Deposited Stainless Steel Layer Due to the material properties of the microcast stainless steel, the cutting operations were rather
Figure 8.17: IMS-T2 Test Part after Shaping of Stainless Steel Compacts slow. In some portions of the part, the 3D cutting strategies could not be used, since filleting and chamfering of some edges produced many tiny facets. A 21/2D (stair-step) contouring strategy with a step height of 0.127 mm was used instead. The current programs for the generation of the 3D cutting strategies cannot interpolate between those facets, and resulted in highly inefficient and long machine code, which could not be run due to time limitations (e.g. one 3D contouring file had over 40000 lines!). This problem will be avoided in the future by using nonlinear CAD models, which will result in the generation of more efficient CNC code. The large amount of copper support material, which had to be deposited created a big heat sink, and created problems for
153
Figure 8.18: Completed IMS-T2 Test Part Fully Embedded in Copper Support the deposition of the copper support material. Due to reduced initial interface temperatures, resulting in insufficient flow characteristics of the copper droplets, voids were present in the copper support material. In some cases, the voids caused penetration of the stainless steel along the edges of the layers, resulting in little extensions on the bottom of the layer (similar to a tiny flash) (Figure 8.19). Also, stress relieving was not used to manufacture the IMS-T2 part, and in certain areas warpage, a few cases of delamination, and some surface areas which show a “Christmas Tree” structure are present. Figure 8.20 a) shows the left corner on the front of the part. Delamination occurred between the first and the second layer on the base, due to increased residual stresses from the layered build up. The corner edge of the wall shows a “Christmas Tree” structure and the layer interfaces warp upwards. Figure 8.20 b) shows the right corner on the front of the base. Warpage and delamination between the first and the second layer are visible.
1 mm
Figure 8.19: Penetration of Stainless Steel into the Copper Support
154
1 mm
b)
1 mm
a) Figure 8.20: Delamination, Warpage and “Christmas Tree” Structure on the IMS-T2 Test Part
The copper support structure of the part was removed in a bath of 70% nitric acid at a temperature between 333 and 348 K. Figure 8.21 shows the IMS-T2 part during the etching process.
Figure 8.21: IMS-T2 Test Part during Removal of the Copper Support Structure
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Overall, however, the IMS-T2 part has shown, that the SDM process and microcasting are able to produce three dimensional, metal structures with arbitrary shape. The quality of the surface finish still needs to be improved, and stress relieving techniques have to be employed, to eliminate the effects caused by residual stresses, to meet the specifications of fully functional parts. With respect to other rapid prototyping processes, such as stereolithography, the IMS-T2 part created with SDM reached comparable geometric tolerances, while greatly improving upon the material strength and quality. As of the time this thesis was written, SDM is the only rapid prototyping process which was able to directly produce the IMS-T2 part in metal. 8.2.4.1. Materials Usage for IMS-T2 In the current, experimental setup of the SDM process, not all operations are optimized with respect to material usage and time consumption. To deposit the part and the support material, 4.5 kg of 308 stainless steel and 3.6 kg of copper wire were used. Approximately 5.7 m3 of Argon gas was necessary for the microcasting process. 640 l of liquid nitrogen were used for shrouding. Preheating the substrate with the plasma spray torch consumed approximately 5.7 m3 of argon and 0.8 m3 of hydrogen. Material usage by the deposition operation can be optimized through more selective deposits and covering only the minimally required areas. Analyzing the cross-sectional areas and the heights of the layers of the test part results in a required deposit (including approximately 15% overdeposit on the edges) of 32 cm3 or 252 g of stainless steel and 154 cm3 or 1358 g of copper. Considering the wavy surface of microcast deposit, which amounts to roughly half the volume of the deposited material, and which has to be removed, the required amounts of material in an optimized deposition process are 500 g of stainless steel, 2700 g of copper, 35 m3 of argon and 252 l of liquid nitrogen. The amount of gases used for the preheat will not change. The amount of copper could be further optimized, by smart design of the support structure, keeping the support structure geometry as close to the part surface as possible. The final weight of the part is 183 g. The intricate geometry of the part, material properties produced by the microcasting process, which create hard to cut materials, and the delicate nature of stainless steel in terms of machinability, resulted in the use of a great number of cutting tools. For the planing operation, a face mill with 4 carbide inserts was used. 7 changes of inserts were necessary to complete the part. For the contour cutting operations, HSS endmills with a titanium-corbonitrate coating for cutting stainless steel and HSS endmills with titanium-nitrate coating for the copper were used (see Table 8.7). Optimizing the amount of material deposited can mainly improve upon the amount of tools required for planing the wavy surface of the microcast deposit. The influence on the amount of tools required for shaping is small, since the length of the contours remains the same. The copper support material for the IMS-T2 part was removed in a nitric acid bath. 7 l of 70% nitric acid at a temperature between 438 and 448 K were necessary, to remove 1.041 kg of microcast copper from the part. A processing time of 6 hours was required to remove the support structure. Reducing the amount of deposited support material, and removing accessible support material through cutting before the etching process can decrease the amount of acid and time necessary to remove the support material.
156
157
Figure 8.22: Stainless Steel IMS-T2 Test Part Manufactured with the SDM Process
Table 8.7 summarizes the amount of materials required to manufacture the IMS-T2 test part. An estimated amount of materials required is shown for an optimized process. The cost of the materials and tools are based on the actual amounts paid. In case of an industrial production environment, the cost can decrease, especially for the gasses. If liquid, instead of compressed argon is used, the cost for argon can decline by a factor 10 to 30. Actually Used Material
Est. Usage for Optimized Process
Amount
Cost [US $]
Amount
Cost [US $]
Stainless St.
4.5 kg
56
0.5 kg
6
Copper
3.6 kg
32
2.7 kg
24
Argon
5.7 m3
30
2.2 m3
12
Liquid Nitrogen
640 l
180
252 l
71
Argon
5.7m3
30
5.7 m3
30
Hydrogen
0.8 m3
3
0.8 m3
3
Planing
Carbide Inserts
7
63
2
18
Endmills St.St.
12.70 mm
6
173
4
116
6.35 mm
6
114
2
38
3.18 mm
7
133
7
133
12.70 mm
2
39
2
39
6.35 mm
2
26
2
26
3.18 mm
3
39
3
39
Nitric Acid
7l
66
4l
38
Deposition
Preheat
Endmills Cu
Cu Removal
984
593
Table 8.7: Material Usage for the IMS-T2 Test Part
8.3. Building Times To compare the performance of the SDM process to other rapid prototyping systems, the processing time is analyzed for two different cases. Typically, three phases are required to build a rapid prototype part: pre-processing, building and post-processing. During the pre-processing phase the computer model of the part is analyzed and decomposed into slices or in the case of the SDM system into compacts and manufacturing instructions, which are used during the building phase to create the part. The post-processing phase includes the removal of the support structures and cleaning or finishing of the surface. For the pre-processing phase in the SDM process, realistic values for the pre-processing times cannot be given to this date. For most of the cases, the planning and scheduling algorithms can
158
derive the building instructions with out external interaction. However, the planning system and the process itself, are still under development, not all available routines are fully stable, and in certain cases adjustments have to be made. For a fully developed version of the SDM process planner, pre-processing is expected to function without any manual interaction from a user, and the pre-processing times are therefore reduced to the machine-time required for computation.
8.3.1. IMS-T1 The IMS-T1 test part was built in 149 layers with a thickness of 0.25 mm from a sprayed zinc nickel/aluminum laminate without 3D contouring. The building time for the IMS-T1 test part is analyzed for the cross-section of layer 61, which is typical for the part. The overall area (projection) of the part is 38.7 cm2, the area for the support material 55.7 cm2. The cross-sectional area of layer 61 is 14.4 cm2 with a perimeter of 267 mm. The cutter path after offsetting the perimeter of the cross-section has a length of 378 mm for the rough contouring (12.70 mm cutter) and a length of 296 mm for the final contouring (3.18 mm cutter). The manufacturing sequence consisted of grit-blasting, deposition of the part material, planing and cutting the shape of the current cross-section, grit-blasting, deposition of the support material and a planing operation to finish the layer. Table 8.8 shows the actual time in seconds required for each of the operations, and estimated values for optimizing the system.
Part
Grit-Blasting Spray
Transfer
Preparation and Finish
Operation
Optimized Operation
Future Optimization
69
74
55
30
30
27
20
20
27
20
5
106
60
15
15
449
0 ÷ 100
0
15
163
120
120
15
342
10 ÷ 150
10
15
83
60
60
Preheat Bond-Coat
86
198
Zinc Cut
Plane Rough Contour Mow
93
Fine Contour Support
Grit-Blasting
69
74
55
30
30
Spray “392”
86
63
117
60
15
Plane
93
15
203
80
80
496
484
1627
430 ÷ 670
385
2607
Table 8.8: Processing Time in Seconds per Operation for IMS-T1 Layer #61 The actual time required to build layer 61 is broken into 3 categories. The transfer time shows the time required to deliver the part to a particular station, and to pick it up after the processing step
159
has been completed. For 6 transfers per layer a total time of 8 min 16 s was required accounting of 19% of the total layer build time. Due to the robotic transfer and pelletizing system, which requires the pallet to be taken off the receiver and being put onto a different one for every operation, this time is rather long. A newly developed manufacturing cell, that does not require the flexibility of a programmable transfer system, could use dedicated transfer stages, and an estimated total transfer time between 45 and 60 s seems possible. The ‘preparation and finish’ column shows the time which is spent to prepare and shutdown a station or process, after the pallet has been delivered or before it is picked up. For the deposition processes this means mainly getting the gun off the rack, starting it up, shutting it down and putting it back, for the grit-blasting process most of this time is spent by the transfer robot taking the deposition rod and moving it in place. Optimizing the process operation, by executing those task before the pallet is brought to a particular station, or after it has been picked up can potentially eliminate all of this time. While this change would be absolutely required for a production type setup, it is to some extent impractical for an experimental setup, which requires manual interference and inspection. The total time used for process preparation and finishing is 8 min, accounting for 19% of the total build time. The time consumption for the actual operations is 27 min 7 s or 62% of the total time of 43 min 27 s required to build the whole layer. 7% of the time is used for grit-blasting, 17% for spraying and 76% for the shaping operations. The distribution between the materials is 77% for the part and 23% for the support material. The processing steps for the support material require less time, since the deposition is slightly faster, and mainly because no contouring is required. Using the manufacturing time for layer 61 as a base, the time required to manufacture a 149 layer part is approximately 108 h. In the current experimental setup, the operation times can be improved by optimizing process planning, path generation and to a small extent by improving upon the process parameters. More efficient path planning and selective deposition can reduce the time required for the deposition steps. If the part material is deposited with strict control of the layer thickness, the resulting layers eliminate the need for the intermediate planing step. The whole layer can be planed after the deposition of the support material. Path planning for contouring can incorporate selective removal of material and therefore shorten cutting durations (especially for mowing) significantly. Using these changes to optimize the processes will result in an approximate building time between 7 and 11 minutes for a similar layer, resulting in a total building time for the part between 18 and 28 h (+ transfer time) with the currently installed system. The last column in Table 8.8 shows estimated times for the individual processing steps for a system with optimized equipment. The one nozzle spray torches can be replaced by multi-head torches, which can speed up the deposition process. Fully optimized planning and path generation strategies are assumed. With an assumed transfer time of 60 s the layer build time decreases to approximately 61/2 min, and the whole part can be built in 16 h. Melting the support structure and cleaning of the part surface required about 2 hours of post-processing time. A direct comparison of part building times between the SDM process using sprayed material dep-
160
osition and other rapid prototyping technologies is not possible, since data is not available for processing times of IMS-T1 on other systems. However, processing times for the IMS-T2 part are available from [60]. The building times for a full scale model of the IMS-T2 part varied between 21/4 and 9 h for lithography based processes1. While the current experimental operation exceeds the building times by far, a future, optimized system can compete with current rapid prototyping systems.
8.3.2. IMS-T2 A half-scale model of the IMS-T2 test part was built with stainless steel and copper as a support material. 29 layers with a total of 79 compacts (31 part, 48 support) were necessary. The average layer height is 1.45 mm. The overall area (projection) of the part is 38.7 cm2, the area of the support material 49.0 cm2. The number of compacts per layer changes throughout the part, the average layer build time is therefore derived from the total build time for the part. The manufacturing sequence for each compact consisted of preheating the substrate, microcasting the material, planing, contouring, 3D shaping and washing. Contouring and 3D shaping is not necessary for the last compact in each layer. The times calculated for each operation and material and values for an optimized process are shown in Table 8.9. The deposition process consists of a preheating step and the microcasting step. While the time required for preheating remains the same for each compact, the time required for material deposition depends on the area which has to be covered. For the calculation of the time required for the deposit it will be assumed, that the material is deposited selectively and will only cover the necessary areas. The CAD model of the part was used to calculate the area of the bottom cross-section of each compact, to evaluate the total deposition time. The total area for all compacts is 225 cm2 of stainless steel and 916 cm2 of copper. With a coverage of 6.91 cm2/min and 8.80 cm2/min, 33 min are required to deposit the stainless steel and 104 min for the copper. Transfer times and process preparation and shutdown times are shown in separate columns. For the cutting process, the contours of the cross-sections of the CAD model were offset to generate the cutter pathes for all compacts. Three different size endmills (12.70, 6.35 and 3.18 mm) were used. The length of the pathes are a total of 10.52, 9.75 and 9.12 m for the stainless steel compacts, and 3.68, 2.97 and 2.69 m for the copper compacts. Several pathes have to be taken with each tool to cut the required depth for one layer (2, 2, 3 for stainless, 1, 2, 3 for the copper). With the specified cutting parameters from Table 8.6 this results in total 2D contouring times (without transport and process preparation) of 10 h 38 min for the stainless steel compacts and 96 min for the copper compacts. The time needed for the 3D shaping is calculated by assuming about 25% of the cutting length of the smallest cutter contains surfaces requiring 3D cutting, half of 1. It should be noted, that the IMS-T2 part has a slightly different geometry, and is mainly constructed from thin walls. While some features of the outer shape are the same, the cross-sectional area at the same height is only 30% of the one from IMS-T1. Considering the fact, that processing times increase with the cross-sectional area for lithography processes with laser scanning, and that the IMS-T1 part built with SDM is a half scale model, allows a direct comparison of the building times (The cross-sectional area of the full-scale IMS-T2 is only 16% larger than the cross-sectional area of a half-scale IMS-T1 at the height of layer 61).
161
Stainless Steel 31 compacts
Deposit
Cut
Transfer
Preparation and Finish
Operation
Future Optimization
Preheat
44
68
78
20
Microcasting
44
15
33
33
Plane
8
145
110
Rough Contour 1
8
332
300
8
171
60
Fine Contour
8
135
100
3D Contour
8
338
300
37
0
93
60
Preheat
44
106
120
30
Microcasting
44
24
104
104
Plane
12
225
180
Rough Contour 1
12
40
35
12
33
15
Fine Contour
12
23
20
3D Contour
12
58
50
58
0
144
100
367
313
2072
1517
Rough Contour 2
Washing Copper
Deposit
48 compacts Cut
Rough Contour 2
Washing
48
48
2752
Table 8.9: Total Processing Times in Minutes for IMS-T2 which will be shaped with a 5-axis strategy, half with a zig-zag strategy. A maximum surface roughness of 2 µm requires a path spacing of a tenth of the cutter radius (see Figure 2.21). The length of the whole 3D cutting path results to 2.5 times the lengths of the 2D contouring path. A total of 6 h 36 min is required for the execution of the 3D contouring steps. The planing steps are assumed to be the same for each compact. For the overall dimensions of 84 x 59 mm a path length of 381 mm is assumed. For removing the wavy surface of the microcast material three consecutive planing operations are necessary for each compact. The total, estimated build time for the half-scale, microcast IMS-T2 part is approximately 46 h of continuous manufacturing on the current, experimental setup. 14% of the total time are used for transportation of the pallet, 11% for process preparation and shutdown, and 75% are the actual operations. Of the time required for the operations, 16% are used for deposition, 72% for shaping and 12% for washing. The distribution between the part material (stainless steel) and the support material (copper) is 64% and 36%. Several measures can be taken, to further optimize the building process. Most of the transfer time can be eliminated by using a different pallet shuttling system, which does not require frequent placement of a pallet with a robot. If the preheating step can be done at the same station as the
162
microcasting deposition, the transfer time can be reduced to approximately 60 s per compact, resulting in a total transfer time of 79 min. Through improved planning and control mechanisms, the process preparation and shutdown times can be eliminated, if processes are started before the pallet reaches the particular station. While the microcasting process does not leave room for improvement in deposition times, the use of a different preheating process, and selective cutting strategies can further reduce the processing time. The estimated build time for the half-scale IMST2 test part with an improved process is approximately 26 h. To compare the performance of the SDM process with the current rapid prototyping technology, the time for building a full-scale version of the test part has to be estimated. While transportation and washing times would remain constant for one layer, deposition and planing times would increase by a factor 4, cutting times would approximately increase by a factor 2, and the part would require twice as many layers. Using the optimized values from Table 8.9 and the optimized transfer time of 79 min, the total time required to build a full-scale version of the IMS-T2 part out of stainless steel is 130 h. If two parts are built simultaneously (i.e., when material is deposited on one part, the other part is shaped), the build times can be reduced by approximately 25%, which would result in a build time of 98 h. Compared with stereolithography, this is between a factor 10 and 40 slower [60]. It has to be considered, however, that the parts created with stereolithography are non-functional plastic structures, while the SDM part has the material properties of stainless steel.
8.4. Conclusions 8.4.1. Summary of Current Work To extend the capabilities of current rapid prototyping technology towards the fabrication of fully functional parts and assemblies, a novel process, Shape Deposition Manufacturing (SDM), has been developed. In its concept, SDM is a layered manufacturing process, employing separate three dimensional shaping and material deposition steps, to achieve high material quality and accuracy, which is required by fully functional structures. The requirements for a process planning environment, incorporating the three dimensional nature of the SDM process, have been identified and implemented to test the concept of the process. A novel, droplet based deposition process for metals, microcasting, has been developed, to create layered deposits with high material quality under conditions suitable for a layered manufacturing technology. Support materials, which are required by layered manufacturing processes to create undercut features, have been identified. Various strategies, to create parts with different complexity and material properties have been developed. To prove the concept, the SDM process was implemented in a testbed manufacturing cell. A variety of test parts have been successfully built, which show the potential of SDM as a future alternative, rapid manufacturing process. Three dimensional shaping of each layer has shown improvement on the surface quality of the parts, and eliminated the stair-step texture inherent to conventional rapid prototyping technologies. Materials deposited with the microcasting process
163
show strength properties, which are comparable to (for mild steel) or even above (for stainless steel) the values for conventionally manufactured materials and lead to structures suitable for full functionality. However, residual stress built up during thermal deposition, combined with a layered manufacturing approach lead to problems, which cause warpage and possible delamination of layers, and an effect, which causes distortion in underlying geometries (“Christmas Tree” effect). The causes for these problems have been identified, possible solutions have been discussed, and will be the focus of future research efforts. Overall, it was shown, that SDM is capable of producing metal parts with high strength and surface qualities comparable to conventional rapid prototyping processes (e.g., stereolithography). It is believed, that the problems related to the residual stress buildup can be resolved, and the rapid production of parts with superior material properties is possible. Future commercialization of the process is expected to have a major impact upon manufacturing, especially in tool fabrication, where short turn-around times are required, and one-of-a-kind manufacturing. To this date, SDM is the only solid freeform fabrication process known to the author, capable of directly producing fully strong metal parts and assemblies.
8.4.2. Future Research and Goals In addition to resolving the problems related to internal stress buildup mentioned above, several other areas are open for improvement of the SDM process in the future. The process planning environment has to be extended into a non-linear CAD planning system, to eliminate some of the limitations imposed by the current linear version, and to allow the implementation of improved planning strategies. Improvements on the microcasting process include the implementation of better process control, to allow monitoring and independent adjustment of droplet temperature, size and rate during the deposition process. Substrate temperature measurements have to be included in combination with active substrate cooling and heating. The concept of a laser-based microcaster has to be further investigated, and the range of materials used in the microcasting process has to be extended. An automated, image-based inspection system, to check the integrity of each layer, also has to be implemented in the SDM process. On a wider scale, the SDM process can be extended into several other areas. Currently, efforts are under way to manufacture conformable, embedded, electro-mechanical systems with the SDM process. Three dimensional epoxy structures and stacked, planar circuit boards using conventional surface mount technology are simultaneously manufactured in the layered approach (see Figure A.14 and A.15 in the Appendix). Future research has to investigate the possibility of true three dimensional component placement and routing, and embedding electronic devices in microcast material to produce smart, functional shapes. Other possibilities are the creation of sensor equipped mechanical structures. The layered manufacturing approach allows placement of the sensors at any given place inside the structures. Sensors suitable to be embedded in microcast material have to be developed (A type K thermocouple has been successfully embedded into microcast stainless steel; see Figure A.16 in the Appendix). Another goal of the author is to extend the concept of SDM and the creation of smart, functional shapes into a millimeter size scale, to produce “milli-mechanisms”, and “milli electro-mechanical
164
assemblies”. To accomplish this goal, the development of millimeter or submillimeter sized sensors and actuators, suitable to be embedded in assemblies and structures, is necessary.
165
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Appendix A: Pictures
Figure A.1: Transfer Robot Loads Pallet onto Milling Machine
A-1
Figure A.2: Planing of Sprayed Deposit
Figure A.3: 3D Contouring of Conically Shaped Layer
A-2
Figure A.4: Thermal Spray Torches and Deposition Robot
Figure A.5: Plasma-Arc Spraying of Zinc
A-3
Figure A.6: Microcasting Deposition Station
Figure A.7: Microcasting of Copper
A-4
Figure A.8: MIG-Microcaster with Water-Cooled Copper Electrode
A-5
Figure A.9: Plasma Microcaster without Shroud
A-6
Figure A.10: Shrouded Plasma Microcaster
A-7
Figure A.11: Microcasting Droplet Formation and Deposition
A-8
Figure A.12: Grit-Blaster/Shot-Peener - Pallet Loading Operation
Figure A.13: Grit-Blasting Operation
A-9
Figure A.14: Embedded Electronic Circuit during Buildup
Figure A.15: Embedded, Electro-Mechanical Structure
A-10
Figure A.16: Type K Thermocouple Embedded in Microcast Stainless Steel
A-11
