THESIS. Presented in Partial Fulfillment of the Requirements for the Degree Master of Science in the Graduate School of The Ohio State University

Transcription

1 Design and LENS Fabrication of Bi-metallic Cu-H13 Tooling for Die Casting THESIS Presented in Partial Fulfillment of the Requirements for the Degree Master of Science in the Graduate School of The Ohio State University By Akshay Jain Graduate Program in Industrial and Systems Engineering The Ohio State University 2013 Master's Examination Committee: Dr. Jerald Brevick, Advisor Dr. Blaine Lilly

2 Copyright by Akshay Jain 2013

3 Abstract Thermal fatigue is one of the most common causes leading to die failure in die casting. This thesis investigates and presents the results of thermal fatigue life in a bimetallic H13-Copper die which can be manufactured using laser additive manufacturing technologies now commercially available. Using finite element method, computational models are developed to simulate the thermal fatigue tests of Wallace (Benedyk, Moracz, and Wallace 1970). Numerical solutions to the thermal-mechanical problem are obtained. The solutions include temperature, strain and stress distributions within the test sample. Solutions were obtained for varying amounts of copper in the test sample geometry. Results from the pure H13 sample computational model compared very well with experimental values obtained by Wallace (Benedyk, Moracz, and Wallace 1970). The maximum temperature reached by the test sample is shown to decrease with increasing amounts of copper. The fatigue life is calculated using the method of universal slopes which relates the calculated cyclic strain ranges to the number of cycles necessary for fatigue crack initiation. The specimen geometry consisting of a half thickness of Cu and the other half thickness of H13 at the thinnest point in the full crosssection of the wall thickness was shown to provide the best balance between thermal and fatigue life performance. ii

4 This document is dedicated to my Family and Friends iii

5 Acknowledgments I would like to thank each and every person who was instrumental in getting this research work done successfully and supported me both technically and emotionally. I would like to sincerely thank my advisor Dr. Jerald Brevick, Associate Professor, Industrial and Systems Engineering for his constant encouragement and guidance throughout this research work. I am grateful to him for his valuable time and shall always cherish my association with him. I would like to thank Dr. Blaine Lilly for standing on my Master s examination committee. I would also like to thank my friend Niranjan Rajagopal for sharing his knowledge and helping me. Lastly, but by no means least, I would like to thank my father Ashok Jain for always encouraging and supporting me in my education. iv

6 Vita March 4, Born Navi Mumbai, India B.S. Mechanical Engineering, Mumbai University Spring 2012 to present...m.s. Industrial and Systems Engineering, The Ohio State University August 2012 to present...graduate Research Associate, Department of Industrial and Systems Engineering, The Ohio State University Fields of Study Major Field: Industrial and Systems Engineering v

7 Table of Contents Abstract... ii Acknowledgments... iv Vita... v Table of Contents... vi List of Tables... ix List of Figures... x Chapter 1: Introduction... 1 The Casting Process... 1 Common Problems faced in Die Casting... 2 Thermal Fatigue... 2 Spray Lubrication in Die Casting... 4 Bi-metallic Die Components... 4 Laser Additive Manufacturing to Manufacture Bi-metallic Components... 5 Research Summary Chapter 2: Literature Review Physical Case Study: Thermal Fatigue tests by Wallace vi

8 Previous work with Bi-metallic Die Components Chapter 3: Methodology General Approach to Solution Temporal Division of the Casting Cycle Modeling of the Wallace Thermal Fatigue Test Temporal Division of the Wallace Test Cycle Modeling the Spray Time Evaluation Test Boundary Conditions for Computational Model Chapter 4: Results Numerical Solutions to the Thermal Problem of Wallace s Thermal fatigue Tests Determination of the right Boundary Conditions Multi-Cycle Run Thermal performance of Bi-metallic H13-Cu Models Numerical Solutions to the Displacement Problem of Wallace s Thermal fatigue Tests Stress-Strain Behavior of the Test Sample Discussion of the Stress-Strain Behavior: High Cycle vs. Low Cycle Fatigue Stress-Strain Performance of Bi-metallic Cu-H13 Models Calculating the fatigue life using Method of Universal Slopes vii

9 Spray Time using Bi-metallic Components Mesh Size and Solution Accuracy Physical Testing Chapter 5: Conclusion and Future Work Future Work References Appendix A: Method of Universal Slopes Appendix B: Properties of H13 Tool Steel and Copper used in Computer Model viii

10 List of Tables Table 1. Minimum and maximum temperature attained by all four test samples in steady state cycle Table 2. Maximum von Mises stress observed for all four test samples Table 3. Maximum strain range for pure H13, 25% Cu, 50% Cu and 75% Cu test samples Table 4. Cycles to failure for pure H13, 25% Cu, 50% Cu and 75% Cu test samples Table 5. Spray cooling required for pure H13, 25% Cu and 50% Cu test samples ix

11 List of Figures Figure 1. LENS process ( Laser Engineered Net Shaping, n.d.)... 7 Figure 2. H13 tooling made using LENS process ( Laser Engineered Net Shaping, n.d.)... 7 Figure 3. DMD rebuilding of turbine blade squealer tip (Dutta, 2013)... 9 Figure 4. Wallace thermal fatigue test sample (Benedyk, Moracz, and Wallace 1970, 189) Figure 5. Temperature vs. time traces for Wallace thermal fatigue test samples (Benedyk, Moracz, and Wallace 1970, 191) Figure 6. Plot of nd 2 vs. number of cycles for various hot work tool steels from Wallace thermal fatigue test (Benedyk, Moracz, and Wallace 1970, 201) Figure 7. Plot of average maximum crack length vs. number of cycles for various hot work tools steels from Wallace thermal fatigue test (Benedyk, Moracz, Wallace 1970, 201) Figure 8. Schematic drawing of core pins prepared for the HDPC trials (a) long bimetallic (b) long tool steel (c) short bi-metallic and (d) short tool steel core pin (Imran et al., 2012) Figure 9. Optical photograph of core pins after HDPC trials from left to right: bi-metallic short, bi-metallic long, tool steel long and tool steel short (Imran et al., 2012) x

12 Figure 10. Bi-metallic Cu-H13 core (Midson, 2013) Figure 11. Temporal division of a typical die-casting cycle Figure 12. Quarter symmetry used in computer model Figure 13. Representation of Wallace thermal fatigue test cycle. 1, 2 and 3 as delineated in figure Figure 14. Steps during spray time evaluation test cycle Figure 15. Temporal division of spray time evaluation test cycle Figure 16. Boundaries of the two dimensional Wallace quarter symmetry used in computer model Figure 17. Geometry and mesh used for computational model Figure 18. Temperature vs. time for exterior node of pure H13 geometry obtained from computational model Figure 19. Temperature vs. time graph for exterior thermocouple used in Wallace s calibration test (Benedyk, Moracz, and Wallace 1970, 191) Figure 20. Temperature vs. time traces for the first 3 cycles of the multi-cycle-run Figure 21. Cross section view of the four geometries tested. Shaded region represents copper. (a) Pure H13 (b) 25% Copper (c) 50% Copper (d) 75% Copper Figure 22. Temperature vs. time traces for the exterior corner element for pure H13, 25% Cu, 50% Cu and 75% Cu for first two cycles Figure 23. X component of total mechanical and thermal strain vs. time for the exterior corner element during first two cycles xi

13 Figure 24. Von Mises equivalent stress (exterior corner) vs. time for the first two cycles of pure H13 geometry; yield strength at temperature = 811 K Figure 25. Von Mises stress vs. time for exterior corner element for pure H13, 25% Cu, 50% Cu and 75% Cu samples; yield strength at temperature = 811 K; yield strength at 700 K Figure 26. Elastic and plastic strain range / fatigue life approximations for H-13 based on the material data in Appendix A (temperature = 294 K) and the method of universal slopes (Conway and Sjodahl 1991, 43) Figure 27. Plot of nd 2 vs. number of cycles for various hot work tool steels from Wallace thermal fatigue test (Benedyk, Moracz, and Wallace, 1970, 201) Figure 28. Plot of average maximum crack length vs. number of cycles for various hot work tool steels from Wallace thermal fatigue test (Benedyk, Moracz, and Wallace 1970, 201) Figure 29. Spray time required to cool test sample below operating temperature for pure H13, 25% Cu and 50% Cu test samples Figure 30. A 1013 element mesh used for 25% Cu model Figure 31. The Optomec LENS machine used for H13 deposition at Northern Illinois University Figure 32. Manufacturing of the test sample. (a) Copper Core. (b) Copper Core with H13 deposited. (c). Final machined test piece Figure 33. Dunk test apparatus used at Case Western University (Case Western University Website) xii

14 Figure 34. Schematic diagram of dunk test apparatus (Case Western University Website) Figure 35. Dunk test specimen post testing. (a) Side 1. (b) Side 2. (c) Side 3. (d) Side 4 59 Figure 36. Elastic and plastic strain range / fatigue life approximations for H-13 based on the material data in Appendix A (temperature = 294 K) and the method of universal slopes (Conway and Sjodahl 1991, 43) Figure 37. Stress vs. strain for H13 at different temperatures Figure 38. Stress vs. strain for Copper xiii

15 Chapter 1: Introduction The Casting Process Die Casting is a process in which molten metal flows into a mold of desired shape by gravity or other forces and solidifies to form the shape of the mold cavity. The part that is made using this process is also known as casting. Some of the advantages of using Die Casting are as follows: 1. It can be used to create complex part geometries, including both external and internal shapes. 2. Some die casting processes can create parts to net shape requiring no further operations. Other die casting processes are usually near net shape requiring a few additional operations to achieve the required dimensions and details. 3. Die casting is used for making a large number of complex parts which would be very uneconomical to make using conventional machining methods. 4. The casting process can be performed to any metal that can be heated to the liquid state. There are also a few disadvantages associated with the casting process. These include limitations on mechanical properties, porosity, poor surface finish and dimensional accuracy, safety hazards to humans when processing hot metals and environmental problems (Groover, 2010). 1

16 Die casting is primarily divided into two sub-groups Expendable-mold casting and Permanent-mold casting. In expendable mold die casting, the mold in which the metal solidifies must be destroyed in order to remove the casting. These molds are usually made from sand, plaster, or similar materials whose form is maintained by using various binders. A permanent mold is made out of metal that can withstand the high temperatures of the casting operation and is used over and over to produce many castings before it fails. Common Problems faced in Die Casting The basic principle of a casting process seems very simple: melt the metal, pour it into a mold and let it cool and solidify. Yet, there are many factors which must be considered in order to get a defect free casting. A common cause of poor quality castings and die failure in die casting is heat checking i.e. surface cracks occurring on the die due to a large temperature change on every cycle. According to Young (1979, 223), Heat Checking is probably the most frequent cause of failure of an aluminum die-casting die. Initial phases of heat checking leads to poor surface finish of the castings and further deterioration of the die due to heat checking may lead to stuck or defective castings. Thermal Fatigue If an unconstrained body is uniformly heated or cooled, it expands or contracts respectively by an amount based on its coefficient of thermal expansion. But if the body is constrained during the slow and uniform heating or cooling, it is unable to freely expand or contract leading to thermal stresses and strains being developed in the body. A body does not have to be constrained to develop thermal stresses and strains. If a body is 2

17 heated or cooled non-uniformly and or/rapidly, it will develop large thermal stresses. This is because when the body is heated or cooled non-uniformly the temperature distribution within the body will vary with the applied temperature field. For example, if the body is heated rapidly, the temperature on the surface layers will rise quickly whereas the average temperature of the body increases slowly. This leads to differential expansion within the body providing internal constraints as different parts of the body expand by different amounts, each part being constrained to some extent by its neighbors. Hence, thermal stresses and strains are developed in the body (Rosbrook, 1992). In die casting, the die is subjected to rapid and non-uniform heating and cooling as molten metal is poured in the die cavity heating the die and the die is then cooled rapidly using sprays and cooling lines. In high pressure aluminum die casting with H-13 dies, Papai and Mobley (1991) observed that the temperature recorded at a depth of 0.01 inches (0.254 mm) below the die cavity surface may rise by as much as 540 F (300 C) during the first second after the injection of molten aluminum (380 aluminum alloy poured at 660C, 1220 F). Similarly, a thermocouple placed at 0.06 inches (1.524 mm) beneath the surface recorded a maximum temperature rise of 360 F (200C) approximately 3 seconds after injection. At a depth of 0.5 inches the thermocouple recorded a maximum temperature rise of only 144 F (80C) and this maximum did not occur until 6 seconds after injection. From the above data it is clear that thermal gradients as high as 4500 F/inch (2500C/in) occur near the die surface during injection. The data for spray cooling of dies also shows similar thermal gradients as large as 2000 F/inch (1100C/inch) are present near the die surface during lubrication spraying (Chhabra, Chu, 3

18 & Altan, 1991). The temperature gradients during spray cooling are, of course, opposite in direction to the temperature gradients during injection. These large alternating temperature gradients lead to strains and stresses that often result in thermal fatigue as the die cyclically heated and cooled. Spray Lubrication in Die Casting In die casting, the surface of the die is sprayed with a lubricant between cycles to remove heat from the die after ejection of casting and deposit a parting agent to prevent soldering between the molten alloy and the casting and allow the part to be ejected easily. Heat has to be removed from the die to maintain the required operating temperature such that the desired casting rate can be achieved. The removal of heat from the die can take a long time depending upon the die geometry thus increasing the cycle time and losing costly production time. If the die can be cooled faster, the cycle time can be reduced considerably, thus increasing production capacity and also reducing the amount of costly die lubricant required. Bi-metallic Die Components The cycle time in a die casting process depends mainly on the cooling or solidification time required. It is a critical factor in the profitability of the die casting industries. During solidification, heat is transferred from the molten metal to the cooling channel through the die material. Hence, thermal conductivity of the die material is the primary governing factor in the cooling process. Copper and Copper alloys have high thermal conductivities and can extract heat from the casting faster, resulting in shorter cycle times, higher productivity and lower overall operating temperature thereby 4

19 increasing thermal fatigue life. Although copper alloys have been used extensively in manufacturing of molds for plastic injection molding, they cannot be used directly as cavity surfaces in high pressure aluminum die casting due to strong chemical affinity of molten aluminum for copper and the abrasive environment of the liquid melt (Zhu, Schwam, Wallace, & Birceanu, 2004). This problem can possibly be eliminated by coating H13 tool steel on the copper molds to act as the cavity surface to ensure sufficient strength and resistance to the harsh environment encountered in the high pressure die casting. Laser Additive Manufacturing to Manufacture Bi-metallic Components Additive manufacturing is a form of direct manufacturing which evolved from rapid prototyping technology in the 1990 s. While rapid prototyping is used to build nonfunctional or semi-functional prototypes, direct manufacturing aims to build fully functional components directly from 3-D computer models. In additive manufacturing, parts are produced by depositing material layer by layer. Laser Additive Manufacturing (LAM) is a fusion-based additive manufacturing process which was developed by Aeromet, a wholly owned subsidiary of MTS Corporation. The Aeromet LAM process evolved from a laboratory based process developed by a team of researchers from Johns Hopkins University and Pennsylvania State University under funding from the U.S. Office of Naval Research and Defense Advanced Research Projects Agency. The project aimed to reduce raw material usage and manufacturing lead time for large titanium aircraft structures by developing a solid free form manufacturing method for aircraft grade titanium (Kobryn, Ontko, Perkins, & Tiley, 2006). 5

20 In recent years, many Laser Additive Manufacturing (LAM) methods have been developed and commercialized. LAM has played a key role in repair of aerospace components for more than a decade and now has been further developed for a wide range of applications. Manufacturing of multi material components is one such application. Laser Engineered Net Shaping (LENS ) is a LAM technology developed by Sandia National laboratories which can be used to fabricate three dimensional metallic components directly from CAD solid models. The process fabricates metal parts using a metal powder injected into a molten pool created by a focused, high powered laser beam. Simultaneously, the substrate on which the deposition is occurring is scanned under the beam/powder inter-action zone to fabricate the desired cross-sectional geometry. Consecutive layers are sequentially deposited, thereby producing a three-dimensional metal component. LENS is claimed to produce parts with mechanical properties similar or better than traditional processing methods. In the LENS process, the build takes place in a closed chamber that is vacated, and then back filled with an inert gas. This limits the size of builds that can be accomplished using the LENS process. Figure 1 shows the LENS process in action and figure 2 shows a H13 tooling manufactured using LENS. LENS can possibly dramatically reduce the time and cost required to realize functional metal parts. LENS can also be used to modify or repair existing hardware ( Laser Engineered Net Shaping, n.d.). 6

21 Figure 1. LENS process ( Laser Engineered Net Shaping, n.d.) Figure 2. H13 tooling made using LENS process ( Laser Engineered Net Shaping, n.d.) 7

22 Direct Metal Deposition (DMD), formerly known as Powder on Metal (POM) is another Laser Additive Manufacturing (LAM) process which originated at the University of Michigan and is now further developed and commercialized by DM3D Technology. DMD provides a major advancement of the current LAM technology. In DMD, an industrial laser beam under CNC/robotic control is focused onto a work piece producing a melt pool into which a small amount of powder metal is injected, building up the part in a thin layer. Based on the CAD geometry, the beam is moved to the next location, tracing out the part layer by layer. Unlike the LENS process, DMD takes place in an open space in air with the build area being flooded with an inert gas making larger build possible. Figure 3 shows DMD being used to a turbine blade squealer tip. DMD has the following key features 1. Patented closed loop feedback control for process. 2. Coaxial nozzles with local shielding of melt pool axis DMDCAM software for additive manufacturing axis moving optics for heavy parts. 5. A user friendly DMD user interface. The DMD closed loop feedback system uses high speed sensors to collect melt pool information, which is directly fed into a dedicated controller that adjusts process inputs such as laser power to maintain part dimensions. 8

23 Figure 3. DMD rebuilding of turbine blade squealer tip (Dutta, 2013) DM3D claims that DMD deposited material is fully dense so the mechanical and physical properties of the DMD material are as good as or better than that of comparable wrought or cast materials. DMD has a wide range of applications covering industries from automotive and aerospace, to oil and gas to power generation and others (Dutta, 2013). Some examples of such application are: 1. Multi-material components 2. Repair and restoration 9

24 3. Hard facing for wear and corrosion protection 4. Free form fabrication Both LENS and POM offer the opportunity to create bi-metallic die components which can potentially lead to better thermal management and increased service life of the die components. Research Summary This thesis discusses the problem of thermal fatigue and lubrication spraying faced in die casting and then presents results of an experiment to study the effect of using a bi-metallic H13-Cu composite in Aluminum Die casting and evaluate the best relative composition for the same. A finite element analysis model is created to predict the fatigue cracking in a geometry studied experimentally by Wallace (Benedyk, Moracz and Wallace 1970). Using finite element analysis, the thermal fatigue tests done by Wallace are simulated. The results from this analysis are compared to results of the Wallace test for verification. The composition of the test sample is then varied to simulate and study the thermal and structural behavior of a bi-metallic H13-Cu composite. Objectives 1. Develop Finite Element Analysis models to study the thermal and structural behavior of a bi-metallic H13-Cu composite. 2. Determine the optimum relative composition of H13 and Cu in the H13-Cu composite for the best thermal and structural performance. 10

25 Approach: 1. Choose sample geometry to be studied. Geometry used by Wallace for thermal fatigue experiments has been used. (Benedyk, Moracz, and Wallace 1970) 2. Choose the relative compositions of H13 and Copper in the H13-Cu composite to be studied. The full cross-section of the wall thickness at its thinnest point in the specimen geometry was divided into the following ratios: a. 75 (H13) : 25 (Cu) b. 50 (H13) : 50 (Cu) c. 25 (H13) : 75 (Cu) 3. Determine the temperature field within the sample using appropriate boundary conditions by finite element method. 4. Determine the stress and strain distributions within the sample using appropriate boundary conditions by finite element method. 5. Predict the fatigue life of the sample by relating displacements, stresses and strains via empirical formulas. 6. Determine reduction in die casting cycle times of the H13-Cu composite as compared to standard H13. 11

26 Chapter 2: Literature Review Physical Case Study: Thermal Fatigue tests by Wallace To verify the results of a computational model, physical test samples with known process conditions are required. A test sample such that its process parameters, material properties and the relationship between extent of heat checking and the number of thermal cycles is known would be preferred. Also, to reduce computational intensity, simplify modeling and ease the interpretation of the results, test pieces with simple geometries are desirable. Based on the above mentioned criteria, test pieces used in the laboratory thermal fatigue studies of Wallace (Benedyk, Moracz, and Wallace 1970) have been chosen for development and evaluation of the computer model. Wallace conducted several tests to study thermal fatigue in Die Casting (Benedyk, Moracz, and Wallace 1970). The effects of variable material composition, heat treatment, surface treatment and coatings on the thermal-cyclic behavior of various tool steels, including H13 were evaluated. To simulate the temperature cycles encountered in die casting dies in a laboratory, Wallace developed a dunk-test. The test was calibrated such that the temperature fluctuations in test pieces measured during testing were similar to those measured in production H-13 die casting dies. The test pieces used were rectangular parallelepipeds, 2x2x7 in. (50.8x50.8x117.8 mm.) as shown in figure 4. The edges of the piece were 12

27 shaped to (0.254 mm.) inch radii. A 1.5 in. (38.1 mm.) diameter hole was drilled in the center of the test piece for internal water cooling. The water used was tap water circulating at 0.85±0.10 gallons per minute (3.22±0.38 liters per minute). Figure 4. Wallace thermal fatigue test sample (Benedyk, Moracz, and Wallace 1970, 189) 13

28 The test cycle consisted of 12 seconds of immersion in an aluminum 380 bath, followed by 22 seconds of air surface cooling and lastly spraying of the exterior surface with a water base lubricant for 1 second immediately prior to subsequent immersion for the next cycle. The outer surface temperature of the test piece, measured at an outer edge, cycled between 200 F (366.5 K) immediately following the spray and 1075 F (852.5 K) at the end of immersion. This is shown in figure 5. Figure 5. Temperature vs. time traces for Wallace thermal fatigue test samples (Benedyk, Moracz, and Wallace 1970, 191) 14

29 The samples were evaluated after a specific number of cycles. A three inch edge section, equidistant from the ends was examined for the cracks. Two evaluation methods were used to study the cracks. In the first method, a parameter that considers total crack pattern area and crack lengths, nd 2, where n signifies number of cracks of a particular size, and d signifies the crack length for that particular size after a specified number of cycles. In the second method, the average maximum crack length, the average length of the longest crack on each of the four edges, was also recorded. Figure 6 shows a plot of nd 2 versus the number of cycles for relative comparison of different die material. For a particular number of cycles, a larger nd 2 value resulting from either a large cracking pattern or large cracks in indicative of inferior resistance to thermal fatigue cracking. Figure 7 shows a plot of average maximum crack length versus number of thermal cycles. These curves were extrapolated to zero crack length to determine the average number of cycles for crack initiation. It should be noted that Wallace s dunk test was created solely to simulate thermal cycles encountered in die casting, it does not take into account the mechanical loading cycles present largely due to cavity pressures experienced in high pressure die casting. 15

30 Figure 6. Plot of nd 2 vs. number of cycles for various hot work tool steels from Wallace thermal fatigue test (Benedyk, Moracz, and Wallace 1970, 201) Figure 7. Plot of average maximum crack length vs. number of cycles for various hot work tools steels from Wallace thermal fatigue test (Benedyk, Moracz, Wallace 1970, 201) 16

31 Previous work with Bi-metallic Die Components Copper based bi-metallic core pins manufactured using DMD were evaluated by Imran, Masood, Brandt, Gulizia and Zahedi (Imran, Masood, Brandt, Gulizia, & Zahedi, 2012). In their research, accelerated die casting trials were performed by using a specially designed die, fabricated from H13 tool steel that had the provision to investigate core pins under the same conditions. The objective of the trial was to evaluate the performance of the bi-metallic core pin in the specially designed die under semi industrial high pressure die casting (HDPC) conditions. Two sets of core pins with different lengths were prepared for HDPC trials. Each set consisted of one pin made from tool steel and another bi-metallic made from copper alloy and tool steel. Figure 8 shows the geometry of the pins used. 17

32 Figure 8. Schematic drawing of core pins prepared for the HDPC trials (a) long bimetallic (b) long tool steel (c) short bi-metallic and (d) short tool steel core pin (Imran et al., 2012) 18

33 The long core pins were designed such that they could be positioned in front of gate entry in the die cavity to evaluate performance under direct liquid aluminum impingement condition. The short pins were prepared to avoid the direct impingement of the injection of molten aluminum. According to Gulizia (Gulizia, 2002), fifty HDPC cycles of the die used in this experiment were equivalent to several thousand in industrial HDPC machine. Hence, each of the four pins was designed for fifty cycles. The experiment found that tool steel core pins could not survive intended number of HDPC shots during trials and resulted catastrophic failure due to severe soldering. This catastrophic soldering particularly substantiated that the die holding time was not long enough for the casting part around the core pin to be solidified completely. Unlike tool steel, the bi-metallic core pins facilitated quick solidification of the casting material around the pin that in turn reduced the soldering and survived the designed number of HDPC shots without any severe failure. Figure 9 shows the core pins after finishing HDPC cycles. Figure 9. Optical photograph of core pins after HDPC trials from left to right: bi-metallic short, bi-metallic long, tool steel long and tool steel short (Imran et al., 2012) 19

34 North American Die Casting Association (NADCA) investigated the performance of a bi metallic Cu-H13 core comparing it to conventional H13 core. A bi-metallic Cu- H13 core was used in production at Albany Chicago Die Casting Company. The bimetallic core reduced operating temperature by ~100 F (~56K) as compared to a H13 core. A conventional core would solder after 7000 shots whereas the bi-metallic core showed no signs of soldering after 7000 shots. The bi metallic core developed a crack in the H13 and had to be removed after ~10,000 shots. The bi-metallic core used is shown in the figure 10 below (Midson, 2013). Figure 10. Bi-metallic Cu-H13 core (Midson, 2013) 20

35 Chapter 3: Methodology To study thermal fatigue and predict heat checking in die-casting dies, a computer based model was developed. The model was developed based on finite-element-method solutions to the thermal-mechanical problem encountered in the experimental thermal fatigue tests conducted by Wallace (Benedyk, Moracz and Wallace 1970). To ensure that the model predicts thermal fatigue and heat checking accurately, it is important that the right initial and boundary conditions are used. The software package used for obtaining the numerical solutions was Ansys Mechanical APDL (2010). The material model includes elastic, plastic and thermal behavior of H13 dies steel and pure Copper (C10100). The elastic behavior is modeled as linear and isotropic. The stress vs. strain graph used for plastic behavior of copper and H13 is presented in Appendix B. The thermal conductivity is assumed to be isotropic, density is assumed to be uniform and the specific heat is also assumed to be uniform. The material properties have been enumerated in Appendix B. The preprocessing, creation of the geometry and finite element meshes and the post processing, i.e. graphical display of results were all done in Ansys Mechanical APDL. 21

36 General Approach to Solution Fatigue failure in die casting is a result of cyclic forces to which the die is subjected during operation. As outlined earlier in chapter 1, these forces are a result of both cyclic mechanical and cyclic thermal loads. (Mechanical loads are not considered in Wallace s tests). Hence, localized displacements of points within the die are functions of both mechanical and thermal aspects of die casting. To determine the displacement of these internal points, it is required that the displacements due to thermal effects are superposed with the displacements due to mechanical effects. The total displacement field due to the superposition of these two sets of displacements can be fully described as u(x, t). (Underlining denotes vectors of dimension three i.e. u, a vector field of dimension three, is a function of three dimensional position vector x, as well as time t, a scalar quantity) The displacements due to the thermal effect can be completely determined from properties of the material and the current temperature field T(x, t). (Thermal effects produce displacements in accordance with thermal expansion which is discussed in Appendix A). In a general thermal-mechanical problem, the displacement field is dependent on the temperature field and the temperature field is dependent on the displacement field, i.e. they are coupled. u = u(t) T=T(u) This coupling is the result of two effects, namely, thermal expansion, and the second law of thermodynamics. The displacement varies with temperature in accordance with 22

37 thermal expansion. Irreversible strain is accompanied by an increase in temperature as explained by the second law of thermodynamics. In die casting, rapid and non-uniform heating and cooling is present leading to large temperature spatial temperature gradients within the body, and where the plastic strain rate within the die is small, the heat generated by irreversible deformation can be neglected (Kovalenko 1969, 33). Based on this assumption, the coupled thermal mechanical problem can be decoupled into two separate problems, one thermal and one mechanical. With this assumption, the temperature field is no longer dependent on the displacement field. u = u(t) T T(u) As the thermal field is independent of the displacement field, solution of the thermal problem requires knowledge of the thermal history and current thermal conditions only. The computational intensity is further reduced when the dynamic effects are neglected. These effects can be neglected in cases of normal heat transfer, where the rate of temperature change is small compared with the speed of sound in the material (Kovalenko, 1969, 33). Neglecting these effects, time becomes a parameter in the quasistatic displacement problem. The thermal and displacement problem can be attacked as follows. Solve the thermal problem in n time steps, t1, t2,, tn to determine the temperature field as function of position and time, T(x,t): 23

38 1. Use the initial thermal conditions, T(x, t 0 ), and the appropriate thermal boundary conditions to solve the heat equation within the bounded region of the die to determine the temperature distribution at time t1, T(x, t 1 ). 2. Use thermal conditions at time t1, T(x, t 1 ), along with thermal boundary conditions to solve for temperature distribution at time t 2, T(x, t 2 ). 3. Continue to increment t and solve the thermal problem until t n, the desired end of the period to be modeled is reached. Using the thermal problem solution, the displacement problem solution is obtained to determine displacements, stresses and strains as a function of position and time, u(x, t), σ(x, t) and ϵ(x, t) respectively. 4. Use the temperature distribution at time t 1, T(x,t 1 ) and initial displacement conditions u(x,t 0 ), along with the appropriate displacement boundary conditions to solve for displacements, stresses and strains at time t 1, u(x,t 1 ), σ(x,t 1 ) and ϵ(x,t 1 ) respectively. 5. Use the temperature distribution at time t 2, T(x,t 2 ), and displacement, stresses and strain conditions at time u(x,t 1 ), σ(x,t 1 ) and ϵ(x,t 1 ) respectively, along with appropriate displacement boundary conditions to solve for displacements, stresses and strains at time t 2, u(x,t 2 ), σ(x,t 2 ) and ϵ(x,t 2 ) respectively. 24

39 Temporal Division of the Casting Cycle The casting cycle is divided into three steps. Each casting cycle consists of a heating, convection cooling and spray cooling. The three steps and their characteristic boundary conditions are shown in figure 11. Figure 11. Temporal division of a typical die-casting cycle Modeling of the Wallace Thermal Fatigue Test For the thermal fatigue tests, Wallace (Benedyk, Moracz, and Wallace 1970, 189) used a simple geometry with a constant cross section as shown in figure 4. Modeling such a geometry is much easier than modeling details from actual dies and the constant cross section of the geometry also allows for a two dimensional version of the test sample to be 25

40 modeled. As the length of the specimen perpendicular to the cross section is much greater than the dimensions of the cross section, strains in the perpendicular direction are assumed to be negligible. Thus, the two dimensional problem is modeled as plane strain. To further reduce the computational intensity, a quarter symmetry model of the geometry is used as shown in figure Figure 12. Quarter symmetry used in computer model Temporal Division of the Wallace Test Cycle The thermal cycle used by Wallace for the dunk test, as shown in figure 13, was used in the modeling. This three region cycle approximation is easier to model than an actual casting cycle. Although the proportional durations of the three regions do not resemble the actual casting cycle, the Wallace test was calibrated such that the 26

41 temperature fluctuations in the test samples during testing were similar to those measured in H13 die-casting dies. Figure 13. Representation of Wallace thermal fatigue test cycle. 1, 2 and 3 as delineated in figure 11 Modeling the Spray Time Evaluation Test To illustrate the effect of bi-metallic die components on the amount of lubricant spray required, a 20 second cycle representative of actual die casting conditions was simulated. The cycle consisted of a 12 second immersion in molten aluminum, a 2 second dwell time where it is air cooled, followed by lubricant spraying to cool the die to operating temperature. Figure 14 shows the different steps during the test cycle and figure 15 shows the temporal division of the test cycle. 27

42 Figure 14. Steps during spray time evaluation test cycle Figure 15. Temporal division of spray time evaluation test cycle 28

43 Boundary Conditions for Computational Model Figure 12 shows the two dimensional geometry used in the computer model. Boundary 1 and Boundary 2 are defined by the symmetry lines 1 and 2 of figure 12 respectively. Symmetry boundary conditions are applied to these boundaries, that is, no heat flux across the boundary and no displacement perpendicular to the boundary. Boundary 3 is in constant contact with the circulating cooling water. The water pressure is neglected and hence there are no mechanical constraints along this boundary. The heat loss along this boundary is governed by the rate of heat transfer from the metal (H13 or Copper) to the water. The boundary conditions at boundaries 1,2 and 3 remain constant throughout the casting cycle, whereas the boundary conditions at boundary 4 and 5 change with each of the 3 steps in the cycle. Figure 16. Boundaries of the two dimensional Wallace quarter symmetry used in computer model 29

44 All the thermal and mechanical boundary conditions are summarized below Boundary 1 and 2 Thermal No heat flux across boundary Mechanical No displacement perpendicular to boundary Boundary 3 Thermal Convection to cooling water Mechanical No mechanical constraints Boundary 4 and 5 Step 1 Thermal Convection from aluminum to steel Mechanical No mechanical constraints Step 2 Thermal Convection from steel to air Mechanical No mechanical constraints Step 3 Thermal Convection from aluminum to cooling spray Mechanical No mechanical constraints 30

45 Chapter 4: Results Numerical Solutions to the Thermal Problem of Wallace s Thermal fatigue Tests The accuracy of the computer model was verified by comparing the results of the pure H13 sample with the experimental results from Wallace s (Benedyk, Moracz and Wallace, 1970) thermal fatigue test. As discusses earlier, the thermal displacement problem is solved as two separate problems, first thermal and then displacement. The experimental temperatures recorded by Wallace during his test calibration were compared to the temperatures obtained from the numerical solution to the thermal problem for the pure H13 geometry. Determination of the right Boundary Conditions The correct boundary conditions for the computer model were obtained from previous work done by Rosbrook (Rosbrook, 1992) on the analysis of thermal fatigue and heat checking in die casting. Figure 17 shows the geometry and mesh used for the computational model. The heat transfer co-efficient between at the sample-melt interface was chosen as 4016 W/m 2 K. The heat transfer coefficient governing the heat loss from the sample to surrounding air was chosen as 24 W/m 2 K. The heat transfer coefficient governing the heat loss from the sample to cooling water was chosen as 7112 W/m 2 K. The heat transfer 31

46 coefficient governing the heat loss from the sample to cooling spray was chosen as 4184 W/m 2 K. Exterior corner element Figure 17. Geometry and mesh used for computational model The heat transfer coefficients used are summarized below: h sample-melt = 4016 W/m 2 K h sample-air = 24 W/m 2 K h sample-cooling water = 7112 W/m 2 K h sample-cooling spray = 4184 W/m 2 K 32

47 The heat transfer coefficients are convection coefficients which relate heat flux per unit area, q, to the difference between the surface and fluid temperature (T s and T respectively), q = h(t s- T ) The surface temperature of H13 and fluid temperatures for air, coolant and spray were estimated as follows: T melt = 950 K T air = 366 K T cooling water = 283 K T cooling spray = 283 K The material properties for H13 and Copper used in the computer model are enumerated in Appendix B. Figure 18 shows the temperature vs. time graph obtained from the computer model for a pure H13 geometry at the exterior node (see figure 17). This exterior node corresponds to the exterior thermocouple used by Wallace in his experiments. Figure 19 shows the temperature vs. time graph measure by Wallace s thermocouple. 33

48 Temperature (K) Temperature (K) Time (Seconds) Figure 18. Temperature vs. time for exterior node of pure H13 geometry obtained from computational model Time (Seconds) Figure 19. Temperature vs. time graph for exterior thermocouple used in Wallace s calibration test (Benedyk, Moracz, and Wallace 1970, 191) 34

49 From figure 18 and figure 19 it is clear that the temperature trend of the computer model compares very well with the temperature profile measured by Wallace. The maximum temperature reached by the exterior thermocouple in Wallace s test (Benedyk, Moracz, and Wallace 1970, 190) is K (1075 o F); the maximum temperature calculated by the computer model at the exterior node during the first cycle is 849 K ( o F) which is very close to the experimental value measured by Wallace. Comparison of temperatures from the first cycle of the simulation to temperatures measured during the steady state experimental cycles is not very accurate and hence a multi cycle simulation is conducted for a better comparison. Multi-Cycle Run Figure 20 shows the temperature vs. time traces for the exterior node for the first 2 cycles. The test piece attains steady state conditions after the first two cycles. Inspection of the numerical temperature data for the exterior element reveals that the temperature values for cycle 2 are very close to the temperature values of cycle 3. Hence, steady state operating conditions are achieved after the first two cycles. In actual die casting, these conditions would not be achieved as rapidly. This is expected as the mass of the steel specimen used by Wallace is very small compared to the mass of the molten aluminum. In actual die casting the mass of steel is relatively large compared to the mass of the casting alloy. 35

50 Temperature (K) Time (Seconds) Figure 20. Temperature vs. time traces for the first 3 cycles of the multi-cycle-run The maximum temperature attained by the exterior element in the second cycle is 851 K (1072 o F); the maximum temperature recorded by the exterior thermocouple in Wallace s experiment was K (1075 o F) which is very close to the value obtained from the computer model. Hence, we can conclude that the computer model is accurate enough. Thermal performance of Bi-metallic H13-Cu Models The thermal performance for varying amounts of Copper in bi-metallic H13-Cu composite samples was studied. The minimum wall thickness of the Wallace test sample was divided into the following ratios: a. 75 H13 : 25 Cu b. 50 H13 : 50 Cu c. 25 H13 : 75 Cu 36

51 Figure 21 shows cross sections of the different geometries tested. (a) (b) (c) (d) Figure 21. Cross section view of the four geometries tested. Shaded region represents copper. (a) Pure H13 (b) 25% Copper (c) 50% Copper (d) 75% Copper 37

52 Temperature (K) % H13 25% Copper 50% Copper 75%Copper Time (Seconds) Figure 22. Temperature vs. time traces for the exterior corner element for pure H13, 25% Cu, 50% Cu and 75% Cu for first two cycles Figure 22 shows the temperature vs. time traces for all the four geometries, i.e. pure H13, 25% Cu, 50% Cu and 75% Cu. From the figure, we see that as the amount of copper is increased in the geometry, the maximum attained by the exterior element falls. This is due to the fact that copper with a higher thermal conductivity than H13 steel leads to faster heat transfer. Table 1 shows the maximum and minimum temperature attained 38

53 by the exterior corner element for all four geometries during the second steady state cycle. Geometry Minimum Temperature (K) Maximum temperature (K) Pure H % Cu % Cu % Cu Table 1. Minimum and maximum temperature attained by all four test samples in steady state cycle Numerical Solutions to the Displacement Problem of Wallace s Thermal fatigue Tests The solutions of the thermal problem were shown to compare very well to the experimental values measured during Wallace s tests. The thermal field is now used an input to the displacement problem. Stress-Strain Behavior of the Test Sample The strain behavior for the pure H13 geometry is shown in figure 23: X- component of total mechanical and thermal strain. As the specimen is heated from an initial stress-less, constant temperature (293 K) condition (point A), compressive stresses are developed in the exterior surface elements due to the cooler interior elements which constrain thermal expansion of the warmer surface elements. The strains are positive as the elements expand while been heated. As the temperature of the neighboring elements 39

54 rise and thermal gradients decrease, the compressive strains decrease. The strains continue to increase and the compressive stress gradually decreases. The strain reaches a maximum value as the sample is removed from the molten bath (point C) after which the surface elements begin to cool and contract and the strain decreases. As the surface elements are now cooler than the interior elements, tensile stresses are developed because the surface elements are constrained from thermal contraction. The tensile stresses keep decreasing as the temperature gradient decreases and the strains also gradually decrease. When the specimen is sprayed (point D), the sudden removal of heat cause large tensile stresses and the minimum strain is reached at the end of the spray period (point E). 40

55 Strain C A B D E F Time (Seconds) Figure 23. X component of total mechanical and thermal strain vs. time for the exterior corner element during first two cycles Immediately after the specimen is sprayed, it s again immersed in the molten aluminum bath (point F) producing compressive stresses and increasing strains. The maximum compressive stress is less than that during the first cycle as the overall body temperature of the specimen is now higher and temperature gradients due to surface heating are smaller. The stresses and strains now follow a pattern similar to what is described above. As steady state is achieved after the 2 nd cycle, the paths for the following cycles are coincident with that of the 2 nd cycle. 41

56 Discussion of the Stress-Strain Behavior: High Cycle vs. Low Cycle Fatigue The strain in the test sample continues to increase while the sample is immersed in the aluminum bath even as the stresses decrease. Such behavior is not characteristic of Hooke s law: stress is proportional to strain. This is due to the fact that when temperature changes are present, Hooke s law must be modified to account for thermal expansion. The strains are due to temperature gradients and thermal expansion. Thus, even when there are no constraints, i.e. there are no stresses, strains, due to thermal expansion may still exist. The yield criteria are based on stress, hence, large strains are not necessarily indicative of plastic behavior. Figure 23 shows maximum strains of approximately 1%. Strains that are greater than 0.2% are usually associated with plastic behavior, however, examination of the von Mises equivalent stress reveals that the sample does not deform plastically in the steady state. Figure 24 shows the von Mises equivalent stress as a function of time for the first two cycles. 42

57 Von Mises Stress (Mpa) σ y (811 K) =855 MPa Time (Seconds) Figure 24. Von Mises equivalent stress (exterior corner) vs. time for the first two cycles of pure H13 geometry; yield strength at temperature = 811 K The maximum stress is reached by the exterior corner at the end of the immersion period when its temperature is 851 K. The yield strength at 811 K is 855 MPa as shown in figure 24. Hence, it is clear that no plastic deformation occurs after the first cycle. Thus, the large strains at the end of immersion are due to large thermal expansion and not caused by plastic deformation. Plastic deformation occurs only during the first cycle due to the initial temperature of the specimen (293 K). Only prior to the first cycle the specimen is this 43

58 cold resulting in such large temperature gradients causing large stresses resulting in yielding. For cycles 2, 3, etc. the sample is warmer than 293 K and the temperature gradients are not as large during immersion. The absence of plastic deformation after the first cycle indicates high cycle fatigue life. Conway and Sjodahl give the following criterion for differentiating between low cycle and high cycle fatigue life: In low cycle fatigue, the peak stresses are above the tensile yield strength, and hence the strains induced usually have a noticeable plastic component. In high cycle fatigue, the strains are confined, at least from a macroscopic point of view, to the elastic region (1991, 2). Hence, thermal fatigue of the Wallace test sample lies within the high cycle regime. Stress-Strain Performance of Bi-metallic Cu-H13 Models The thermal performance of Bi-Metallic Cu-H13 models was discussed earlier. The maximum temperature attained by the models decreased with increase in the amount of Copper. Figure 25 shows the von Mises stress vs. time traces for the exterior corner of the geometry for pure H13, 25% Cu, 50% Cu and 75% Cu models. From the figure we see that as the amount of copper increases, the maximum von Mises stress also increases. For all test samples, we see that the behavior is similar to pure H13 geometry, i.e. the maximum von Mises stress reduces after the first cycle. Plastic strains are observed only in the first cycle in the H13 region. Copper continues to deform plastically during each cycle. The maximum deformation observed in the Copper region is observed close to the H13 Copper interface. 44

59 Von Mises Stress (Mpa) Time (Seconds) σ y (700 K) =1000 MPa σ y (811 K) = 855 MPa 100% H13 25% Cu 50% Cu 75% Cu Figure 25. Von Mises stress vs. time for exterior corner element for pure H13, 25% Cu, 50% Cu and 75% Cu samples; yield strength at temperature = 811 K; yield strength at 700 K The maximum von Mises stress for the exterior corner of each of the four geometries is summarized in the table below. As the amount of copper in the geometry is increased, we see that the maximum stress at the exterior corner H13 element also increases. Also, with increase in Copper, the maximum von Mises stress in the Copper region decreases. 45

60 Geometry Maximum von Mises Stress (MPa) pure H % Cu % Cu % Cu 961 Table 2. Maximum von Mises stress observed for all four test samples Calculating the fatigue life using Method of Universal Slopes Cyclic stress range and strain range are often related, via empirical models, to fatigue life. Equation 5.1, originally due to Manson (Conway and Sjodahl 1991, 40), related cyclic strains ϵ and the number of cycles to failure N f, (5.1) where E is the elastic modulus of the material, and M, G, z, and are other materialdependent properties. Upon investigation of many materials, Mason related M and G to material properties that could be determined with a simple tensile test, (5.2) (5.3) Where D = -ln(1-ra) is the logarithmic ductility (RA = reduction in area), and σ u is the ultimate tensile strength. Furthermore, Mason postulated that the exponents z and γ could be approximated z = -0.6 and, irrespective of the material. Combining the above equations gives the following relation 46

61 (5.4) The first term in the equation represents fatigue life due to plastic strain range; the second term represents fatigue life due to elastic strain range, viz., (5.5) (5.6) This method is known as the method of universal slopes because the material independent exponents, z = -0.6 and, are the slopes of the curves when the relations 5.5 and 5.6 are plotted on log-log axes (Conway and Sjodahl 1991, 42-45). For low cycle failure, fatigue life is best estimated by the relation 5.5; for high cycle failure, relation 5.6 gives the better approximation. Using the properties of H13-steel (temperature = 294 K) enumerated in Appendix A, equations 5.5 and 5.6 are plotted in figure

62 (5.5) (5.6) Figure 26. Elastic and plastic strain range / fatigue life approximations for H-13 based on the material data in Appendix A (temperature = 294 K) and the method of universal slopes (Conway and Sjodahl 1991, 43) Analysis of the strain data for the four samples, pure H13, 25% Cu, 50% Cu and 75% Cu reveals that the largest strain throughout the steady state cycles, i.e. cycle 2, 3, etc. is observed at the end of immersion and the minimum strain is observed at the end of spray cooling. Table 3 shows the maximum total thermal and mechanical strain range experienced by all the four test samples at the exterior corner H13 element and the Copper element experiencing maximum strain range during the steady state cycle 2. As the amount of Copper is increased, the maximum temperature reached in the H13 region keeps on decreasing and the maximum temperature in the Copper region keeps on increasing. Hence, the thermal strain in the H13 region keeps on decreasing with increase in Copper whereas, the maximum thermal strain in the Copper region increases with 48

63 increase in amount of Copper in the test geometry. (Note: the total strain range is equal to the elastic strain range for H13 because, as stated earlier, plastic strain is absent during the steady state cycles in H13) Geometry Maximum Strain Maximum Strain Range in H13 Range in in region Copper region Pure H % Cu % Cu % Cu Table 3. Maximum strain range for pure H13, 25% Cu, 50% Cu and 75% Cu test samples Based on the method of universal slopes, the fatigue life for all the three geometries is calculated and summarized in the table below. Geometry Calculated cycles to failure (N f ) for H13 region Calculated cycles to failure (N f ) for Copper region pure H % Cu % Cu % Cu Table 4. Cycles to failure for pure H13, 25% Cu, 50% Cu and 75% Cu test samples 49

64 Figure 27 and 28 show the fatigue parameters used by Wallace to evaluate fatigue behavior; nd 2 and average maximum crack length are plotted vs. number of cycles. Extrapolation of the base H-13 curves to nd 2 equals to zero and average maximum crack length equals zero, suggests that fatigue cracks begin to appear around 6000 cycles. Our computational model suggests that for a pure H13 geometry the number of cycles to failure is 5827 cycles. Figure 27. Plot of nd 2 vs. number of cycles for various hot work tool steels from Wallace thermal fatigue test (Benedyk, Moracz, and Wallace, 1970, 201) 50

65 Figure 28. Plot of average maximum crack length vs. number of cycles for various hot work tool steels from Wallace thermal fatigue test (Benedyk, Moracz, and Wallace 1970, 201) Spray Time using Bi-metallic Components As the maximum temperature attained by the test sample decreases with increase in the amount of copper and due to the higher thermal conductivity of copper, the amount of spray lubricant required in cooling down the die to operating temperature also reduces significantly. The average operating temperature for Aluminum dies is about 561 K as stated by Schrader, George, Elshennawy, Ahmad, Doyle and Lawrence (Schrader, George, Elshennawy, Ahmad, Doyle, and Lawrence 2000). Figure 29 shows the 51

66 temperature vs. time graph for pure H13, 25% Cu, 50% Cu and 75% Cu test samples. For 75% Cu test sample, operating temperature was reached even before spray cooling. The heat flux value of 720 KW/m 2 used for the sprays was found by averaging the heat flux values at 588 K for 7 commercially available die casting lubricants as determined by Leff (Leff, 1999). The cycle consisted of a 12 second immersion in molten aluminum, a 2 second dwell time where it is air cooled, followed by lubricant spraying to cool the die to operating temperature as mentioned earlier in chapter 3. From the graph, it is clear that the amount of spray cooling required is significantly less with the H13-Cu bi-metallic test sample. For the pure H13 sample, the lubricant had to be sprayed for 2.4 seconds to bring the temperature of the die below 561 K, for 25% copper test sample this time was 0.6 seconds and for the 50% Cu test sample this time was only 0.2 seconds. The results are summarized in the table 5. 52

67 Temperature (K) Time (Seconds) Maximum Die Temp. Pure H13 Maximum Die Temp. 25% Cu Maximum Die Temp. 50% Cu Maximum Die Temp. 75% Cu Figure 29. Spray time required to cool test sample below operating temperature for pure H13, 25% Cu and 50% Cu test samples Test Sample Lubricant Spray time for maximum die temperature to drop below 561 K Percentage reduction in spray time and amount lubricant as compared to pure H13 Pure H % Copper % 50% Copper % Table 5. Spray cooling required for pure H13, 25% Cu and 50% Cu test samples It is important to note that the above results only illustrate the impact a bi-metallic H13-Cu composite can make on the amount of lubrication required. The spray time is very small is the above tests due to the geometry of the test sample which is very simple 53

68 and has thin sections. Real die casting dies can be very complex and have thick walls requiring more than 20 seconds of spraying. Mesh Size and Solution Accuracy As the accuracy of finite element solutions depend upon element size, multiple mesh sizes were tested and compared to ensure that accurate values were calculated. Both the thermal and displacement problem were calculated using an element edge length of 0.6 mm for all the geometries. A finer mesh produced similar results and hence an element edge length of 0.6 mm was determined to be adequate. A coarser mesh would also probably yield fairly accurate results. Figure 27 shows the mesh used for 25% Cu model. Figure 30. A 1013 element mesh used for 25% Cu model 54

69 Physical Testing Manufacturing the Bi-metallic Sample Based on the above results, a 50 H13: 50 Cu sample was chosen to be manufactured and tested physically. This geometry was shown to provide the best balance between fatigue life and cooling performance. A 2 x 2 x 7 pure Copper (C11000) square rod was machined down to x x 7 and a 1.5 diameter hole was drilled in the center for the cooling channel at The Ohio State University. The Copper was machined to x and not x which would provide the ideal 50H13: 50Cu ratio to account for the laser melting Copper while LENS deposition of H13. The machined Copper piece was then sent to Northern Illinois University for LENS deposition of H13 powder under the supervision of Dr. Federico Sciammarella. The H13 powder used for LENS deposition was -80/+270 mesh H13 powder from Carpenter Technologies. The powder was spherical and is called Micro- Melt(R) H13 CPP H13 ( M) by Carpenter Technologies. Following was the chemical analysis of the powder in wt. %: C: 0.41 Si: 0.91 Mn: 0.49 Cr: 5.20 Mo: 1.51 V: 1.04 P: S: Fe: Balance A 0.15 thick layer of H13 steel was deposited onto the Copper core and it was then machined down to the final 2 x 2 x 7 dimension. The deposition rate for the H13 was 55

70 approximately 20 hours mostly due to slow heating rates and impeded purge cycles. The interface microstructure for the test specimen showed complete coalescence between the Copper core and the H13 deposit Figure 31 shows the sample with H13 LENS deposited. No surface cracks were detected in any of the layers during intra-layer visual inspection. The hardness of the H13 tool steel deposit was 58 HRC. Figure 31 shows the OPTOMEC LENS machine used for H13 deposition at Northern Illinois University. Figure 32 shows the copper core, the core with H13 deposited and the final machined sample. Figure 31. The Optomec LENS machine used for H13 deposition at Northern Illinois University 56

71 Figure 32. Manufacturing of the test sample. (a) Copper Core. (b) Copper Core with H13 deposited. (c). Final machined test piece Results from Physical Dunk Test The physical testing of the sample was done at Case Western University by Dr. David Schawm. The dunk test developed by Wallace (Benedyk, Moracz, and Wallace 1970) has been covered in Chapter 2. Figure 33 shows the dunk test apparatus used at Case Western University and Figure 34 shows a schematic diagram of the same. 57

72 Figure 33. Dunk test apparatus used at Case Western University (Case Western University Website) 58

73 Figure 34. Schematic diagram of dunk test apparatus (Case Western University Website) The test piece lasted for 2,890 cycles before the first cracks were noticed. Figure 35 shows the dunk test specimen with cracks. Figure 35. Dunk test specimen post testing. (a) Side 1. (b) Side 2. (c) Side 3. (d) Side 4 59