STUDY OF HOT TEARING EVALUATION METHODS AND QUANTIFICATION OF CONTRACTION FORCES IN DIE CASTING ALLOYS DISSERTATION

Transcription

1 STUDY OF HOT TEARING EVALUATION METHODS AND QUANTIFICATION OF CONTRACTION FORCES IN DIE CASTING ALLOYS DISSERTATION Presented in Partial Fulfillment of the Requirements for the Degree Doctor of Philosophy in the Graduate School of the Ohio State University By Shri Nath Dubey, M.S. Graduate Program in Industrial and Systems Engineering The Ohio State University 2015 Dissertation Committee: Professor Jerald R. Brevick, Adviser Professor Allen Yi Professor Jose Castro

2 Copyright by Shri Nath Dubey 2015

3 ABSTRACT Hot tearing is the undesired formation of irregular cracks in metal castings that develop during solidification and cooling; typically while the casting is still inside the mold or die cavity. The cause of hot tearing is generally attributed to the development of thermally induced tensile stresses and strains in a casting as the molten metal contracts during solidification and solid state shrinkage. Hot tearing often occurs at the inside corners or fillets of casting geometries, where casting shrinkage is constrained by the relatively rigid mold cavity. In the past years, several methods have been employed to evaluate the hot tearing propensity of casting alloys. Specifically, constrained rod casting, crack-ring, horizontal Bar, T-Shape, and Ring mold designs have been used. Constrained rod casting (CRC) has evolved as the most common method, and has been used in several studies to evaluate hot tearing for die casting alloys. The results from CRC methods are not universally applicable, because there is no standardized CRC mold design and the test results are also semi-quantitative at best. ii

4 This research is focused on the design and manufacture of a new CRC mold (Enhanced Constrained Rod Casting-ECRC) for quantitative measurement in determining hot tearing of A206.2, A380, Test-A, and AT72 alloys. A new feeding concept for the ECRC was developed for streamlined flow and for reducing fill time using numerical simulations (Magmasoft). ECRC used four constrained rods at various lengths and with bulbous ends for hot tearing evaluation. A measurement rod was used without bulbous end to measure thermal contraction forces as function of time during solidification with a load cell installed over a quartz rod of a low thermal expansion. Thermal contraction forces were measured as function of time and as hot tearing developed in the casting of A206.2, A380, Test-A, and AT72 alloys, which became evident (via a significant drop in load-time curves). Sprue design was optimized to improve the flow so no area of the casting freezes early. A transient thermal-mechanical FEA model was developed of enhanced constrained rod casting (ECRC) and simulations were performed for thermal strains at various pouring temperatures of A206.2, Test-A, A380, and AT72 alloys. It was observed that thermal strains were a function of time (and temperature) for various pouring temperatures. Hot tearing predictive models were used to study shrinkage porosity and strain rate of die casting alloy to validate the experimental studies. The Niyama criterion model emerged as a good predictor of hot tearing characteristics since hot tearing is influenced by thermal gradients and cooling rates in die casting alloys. iii

5 Dedicated to my parents, my wife, and my children. iv

6 ACKNOWLEDGMENTS First and foremost I would like to thank my advisor Professor Jerald R. Brevick for his guidance, encouragement, and patience during my research and study. Conducting research at the Department of Industrial and System Engineering was a very rewarding and excellent learning experience to me. My sincere thanks go to my thesis committee members Professor Allen Yi and Professor Jose M. Castro for encouragement, critical comments and guidance during research. I want to thank Dr. Allen Luo for opportunity to work on GM project. I would like to thank Mr. Andrew Klarner, Mr. Emre Cinkilic, and Mr. Weihua, Sun for their support during experimental studies of AT72 alloy. My sincere thanks to Mr. Bill Tullos, for his enormous support on experimental studies and casting simulation. I would like to thank Mr. Josh Hassenzahl for support in using casting lab and tools for experimental studies. I would like to thank Mr. Mike Zazon and Mr. Cedric Sze for IT support. Finally, my deepest gratitude to my family. I am grateful to my parents who believed in me. I would like to thank my wife Sadhna and my children - Priya, Ashish, and Nikhil for their understanding and cooperation during Ph.D. research. v

7 VITA EDUCATION Master of Science in Mechanical Engineering August, 1994 West Virginia University, Morgantown, West Virginia Bachelor of Science in Mechanical Engineering December, 1991 South Dakota School of Mines & Technology Rapid City, South Dakota Bachelor Degree in Mechanical Engineering July, 1989 Østfold University College, Sarpsborg, Norway PUBLICATION Dubey, S., N., Brevick. J., R., Overview of Recent Research Regarding Hot Tearing of Die Casting Alloys, Die Casting Congress & Exposition, October 8-10, 2012, at the Indiana Convention Center Halls A&B in Indianapolis, IN. FIELDS OF STUDY Major Field: Industrial and Systems Engineering vi

8 TABLE OF CONTENTS ABSTRACT... ii ACKNOWLEDGMENTS... v VITA... vi PUBLICATION... vi FIELDS OF STUDY... vi TABLE OF CONTENTS... vii LIST OF TABLES... xi LIST OF FIGURES... xiii CHAPTER 1: INTRODUCTION HOT TEARING EVALUATION METHODS SCOPE OF RESEARCH RESEARCH OBJECTIVES RESEARCH METHODOLGY ORGANIZATION OF THE DISSERTATION CHAPTER 2: LITERATURE REVIEW EXPERIMENTAL METHODS FOR EVALUATION OF HOT TEARING CONSTRAINED ROD CASTING (CRC) FOR EVALUATION OF Mg-Al ALLOYS CRACK-RING FOR EVALUATION OF MG-9AL-XZn ALLOY vii

9 2.1.3 HORIZONTAL BAR FOR EVALUATION OF Mg ALLOY T-SHAPE FOR EVALUATION OF A206 ALLOY RING MOLD FOR EVALUATION OF AA1050 VIA ACOUSTIC EMISSION RING MOLD FOR EVALUATION OF Al-Cu AND Mg ALLOYS NUMERICAL MODELING APPROACH AND SIMULATION HOT TEARING PREDICTIVE MODELS SUMMARY AND DISCUSSION CHAPTER 3: THEORETICAL BACKGROUND CHARACTERISTICS OF HOT TEARING STAGES OF CRYSTAL GROWTH SOLDIFICATION PROCESS FORMATION OF DENDRITE MICROSTRUCTURES STRESS-STRAIN RELATIONS AS A FUNCTION OF TEMPERATURE STRAIN THEORIES HOT TEARING CRITERIA HEAT TRANSFER IN CASTING SOLIDIFICATION FLOW OF HEAT INTERFACES WITH CHILL FLOW OF HEAT INTO CASTING HEAT RELEASED DURING SOLIDIFICATION CHAPTER 4 - EXPERIMENTAL STUDIES OF DIE CASTING ALLOYS DESIGNING OF A BOOK MOLD ASSEMBLY viii

10 4.1.1 DESIGNING OF A SPRUE AND RUNNER FOR METAL FLOW DEVELOPMENT OF ENHANCED CONTRAINED ROD CASTING EXPERIMENTAL STUDIES OF A206.2, TEST-A, A380, AND AT72 ALLOYS INSTRUMENTATION OF EXPERIMENTAL APPARATUSES EXPERIMENTAL MATRIX AND CHARATERTICS OF A206.2, TEST-A, A380, AT72 ALLOYS MELTING AND CASTING PROCESSES OF A206.2, TEST-A, A380, AND AT72 ALLOYS FORCE MEASUREMENTS AND DATA ANALYSIS OF A206.2 ALLOY FORCE MEASUREMENTS AND DATA ANALYSIS OF TEST-A FORCE MEASUREMENTS AND DATA ANALYSIS OF A380 ALLOY FORCE MEASUREMENT AND DATA ANALYSIS OF AT72 ALLOY CONCLUSIONS CHAPTER 5 NUMERICAL AND HOT TEARING PREDICTIVE MODELING NUMERICAL AND PREDICTIVE MOEDLING APPROACH FINITE ELEMENT MODELING AND SIMULATION FLOW MODELING AND SIMULATION COOLING AND SOLIDIFCATION OF ECRC MOLD SEQUENTIAL COUPLED TRANSIENT THERMAL STRUCTURAL METHODS ix

11 5.4 THERMO-MECHANICAL MODELING AND SIMULATION CONSTITUTIVE EQUATIONS AT SOLID STATE BOUNDARY CONDITIONS AND MATERIAL PROPERTIES OF CASTING ALLOYS HOT TEARING PREDICTIVE MODELING CONCLUSIONS CHAPTER 6: RESULTS DISCUSSION AND CONCLUSIONS AND FUTURE RESEARCH RESULTS AND DISCUSSION CONCLUSIONS FUTURE RESEARCH REFERENCES APPENDIX-A: EXPERIMENTAL DATA FOR A206.2 ALLOY APPENDIX-B: EXPERIMENTAL DATA FOR A380 ALLOY APPENDIX-C: EXPERIMENTAL DATA FOR TEST-A ALLOY APPENDIX-D: EXPERIMENTAL DATA FOR AT72 ALLOY x

12 LIST OF TABLES Table 1. Castability Index for Mg Die-casting Alloys [13] Table 2. Influence of Zn on Mg-Zn Alloy at Different Mold Temperatures [16] Table 3. B206 Alloys with Various Compositions [17] Table 4. Hot tearing Susceptibilities to Die Casting Alloys Table 5. Hot tearing Susceptibilities [35] Table 6. Hot tearing Ratings [13] Table 7. Characteristic of Hot tearing [37,39,42] Table 8. Types of Tools Applied in Development of Book Mold Assembly [57] Table 9. List of Apparatuses for Experimental Studies Table 10. Chemical Compositions of A206.2 Alloy Table 11. Chemical Compositions of Test-A Alloy Table 12. Chemical Compositions of A380 Alloy Table 13. Chemical Compositions of AT72 (Mg-7Al-2Sn) Table 14. Summary of Hot Tearing Results Table 15. Summary of Hot Tearing Results Table 16. Summary of Hot Tearing Results at 800 C Table 17. Summary of Hot Tearing Results at 750 C Table 18. Summary of Hot Tearing Results at 800 C Table 19. Shrinkage [42] xi

13 Table 20. Physical Properties of A206, A380, Test-A Alloy and AT72 [65, 66] Table 21. Physical Properties of A380 [66] Table 22. Physical Properties of P20 Tool Steel [66] Table 23. Pouring Temperatures for Alloys Table 24. Mold Temperatures for Alloys Table 25. Simulation Time Steps Table 26. Total Thermal and Mechanical Strain of A206 Alloy Table 27. Total Thermal and Mechanical Strain of Test-A Alloy Table 28. Total thermal and mechanical strains of A Table 29. Total Thermal and Mechanical Strain of AT Table 30. The summary of total maximum thermal strain xii

14 LIST OF FIGURES Figure 1. Hot tearing at junction points of constrained rod casting... 2 Figure 2. a) Hot tearing in aluminum ingot b) Hot tearing in extrusion billet [3]... 3 Figure 3. Shoulder cracks in the MC-HPDC recycled AM series magnesium scrap prior to optimization [4]... 4 Figure 4. Hot cracks in Heat Affected Zone (HAZ) of gas tungsten arc welding... 5 Figure 5. Dendritic solidification along with possible hot tearing phenomena in casting [6]... 6 Figure 6. Experimental set up of constrained rod casting [8]... 7 Figure 7. Air entrapment and unsteady flow in CRC Figure 8. A schematic illustration of methodologies for hot tearing evaluation methods 15 Figure 9. Hot tearing in the longest rod [13] Figure 10. Steel mold for constrained rod casting to determine hot tearing [14] Figure 11. Crack-ring mold to determine hot tearing [18] Figure 12. A model of horizontal bar with downsprue [20] Figure 13. Hot tearing effects at mold temperatures 210 and 250 C [20] Figure 14. Hot tearing at reflections (0 0 1) [21] Figure 15. Schematic of instrumented constrained T-shaped casting [22] Figure 16. Cooling curves T and dt/dt-t, hot tearing at 800 C [22] Figure 17. Experimental set up of ring mold casting [23] xiii

15 Figure 18. Hot tearing detection via AE method [23] Figure 19. Crack length as function of Cu content in Al-Cu alloys [24] Figure 20. Crack length as function of Mg content in Al-Mg alloys [24] Figure 21. A 3D model of ingot with bottom block and boundary conditions [25] Figure 22. Hot tearing modeling approach [26] Figure 23. Comparison between hot tear in casting and simulation results [26] Figure 24. A 2D model with mesh and boundary conditions [27] Figure 25. Determination of hot cracking susceptibility [30] Figure 26. λ (Lambda) curve [36] Figure 27. Hot tearing in Al-10Cu alloy (Spittle and Cushway 1983) [37,39] Figure 28. Hot tearing at various carbon contents in Fe-C diagram [41] Figure 29. Start of crystallization, crystal growth, and film stages in solidification [44] 54 Figure 30. Stages of hot tearing in binary Alloys [43] Figure 31. Modeling of dendrites at coherency states [38] Figure 32. Formation of grains and arm spacing [37] Figure 33. Phase changes in stresses-strain curves at different temperatures [42] Figure 34. Strain development from film stage to non-equilibrium solidification [41].. 62 Figure 35. Strain development leading to tear formations [41] Figure 36. Ratio of vulnerability ( tv) to stress relaxation ( tr) for hot tearing [46,47].. 67 Figure 37. A schematic illustration of enhanced constrained rod casting Figure 38. Fluid flow in a tapered sprue (circular) xiv

16 Figure 39. a) Bend radius for inner b) Bend radius for outer radius between sprue and runner [54] Figure 40. Eddies and aspiration [56] Figure 41. Formation of streamlined fluid flow [56] Figure 42. A 2D engineering drawing of left side mold assembly Figure 43. A 2D engineering drawing of right side mold assembly Figure 44. The details of installation for left mold cavity Figure 45. A 3D view of left mold cavity Figure 46. The details of installation for right mold cavity Figure 47. A 3D view of right mold cavity Figure 48. A schematic illustration of experimental set up for ECRC mold Figure 49. Book mold assembly of enhanced constrained rod casting Figure 50. Closed book mold assembly of enhanced constrained rod casting with K-type thermocouple Figure 51. Data acquisition system (Tracer DAQ) with donut load cell [59,61] Figure 52. Quartz rod for force measurement with shaft collar [60] Figure 53. USB-TC thermocouple data logger with K-type thermocouple [62,63] Figure 54. Load cell compression, flat, and ceramic washers [59,64] Figure 55. Instrumentation and installation of testing equipment for ECRC mold Figure 56. Installation of load cell along with quartz rod and washers Figure 57. Support bracket for quartz rod and load cell assembly xv

17 Figure 58. Casting of ECRC as solidified inside mold after pouring of Test-A-alloy at 700 C Figure 59. Contraction force and cooling curve of A206.2 at 700 C Figure 60. A photograph of cast part exhibited hot tearing for A206.2 C at 700 C Figure 61. Contraction force and cooling curve of A206.2 at 760 C Figure 62. A Photograph of cast part exhibited hot tearing for A206.2 C at 760 C Figure 63. Contraction force and cooling curve of A206.2 at 800 C Figure 64. A photograph of the ECRC showed hot tearing cracks at 800 C Figure 65. Contraction force and cooling curve of Test-A at 700 C Figure 66. A photograph of cast part did not exhibit hot tearing for Test-A at 700 C Figure 67. Casting of ECRC as solidified inside mold at 700 C of Test-A-alloy Figure 68. Contraction force and cooling curve of Test-A alloy at 750 C Figure 69. A photograph of cast part exhibited hot tearing for Test-A at 750 C Figure 70. Solidified casting inside mold showed hot tearing cracks at 750 C Figure 71. Contraction forces and cooling curve for Test-A at 800 C Figure 72. A photograph of cast part exhibited hot tearing for test-a at 800 C Figure 73. Contraction forces and temperatures as a function of time for A380 alloy at 700 C Figure 74. A photograph of cast part exhibited hot tearing cracks at 700 C Figure 75. Contraction forces and cooling curve for A380 alloy at 750 C Figure 76. A photograph of cast part did not exhibit hot tearing at 750 C Figure 77. Contraction forces and cooling curve for A380 alloy at 800 C xvi

18 Figure 78. A photograph of cast part exhibited hot tearing cracks at 800 C Figure 79. Contraction forces and cooling curve for AT72 alloy at 675 C Figure 80. A photograph of cast part exhibited hot tearing cracks at 675 C Figure 81. Contraction forces and cooling curve for AT72 alloy at 710 C Figure 82. A photograph of cast part exhibited hot tearing at 710 C Figure 83. Contraction forces and cooling curve for AT72 alloy at 750 C Figure 84. A photograph of cast part exhibited hot tearing at 750 C Figure 85. Model of current constrained rod casting (CRC) Figure 86. Model of enhanced constrained rod casting (ECRC) Figure 87. Filling velocity of CRC at 700 C for A Figure 88. Filling velocity of ECRC at 700 C for A Figure 89. Hotspot of A206 alloy at 700 C pouring temperature Figure 90. Hotspot of A206 alloy at 800 C pouring temperature Figure 91. Cooling and solidification of A206.2 alloy at 700C Figure 92. Solidification shrinkage of A206.2 alloy at 700 C Figure 93. Cooling and solidification of A206.2 alloy at 760 C Figure 94. Solidification shrinkage of A206.2 at 760 C Figure 95. Solidification shrinkage Al-Si alloy at 700 C Figure 96. Solidification shrinkage Al-Si alloy at 750 C Figure 97. Solidification shrinkage Al-Si alloy at 800 C Figure 98. Transient thermal analysis using ANSYS sequential coupling method [70] 150 xvii

19 Figure 99. Model of ECRC for thermal and structural analysis [70] Figure 100. Plane55 2D thermal solid element [70] Figure 101. Model with plane55 thermal solid element [70] Figure 102. ECRC model with boundary conditions [70] Figure 103. Thermal strain for A206 alloy at 700 C Figure 104. Thermal strain for A206.2 alloy at 760 C Figure 105. Thermal strain for A206 alloy at 800 C Figure 106. Thermal strain for Test-A Alloy at 700 C Figure 107. Thermal strain for Test-A alloy at 750 C Figure 108. Thermal strain for Test-A alloy at 800 C Figure 109. Thermal strain for A380 alloy at 700 C Figure 110. Thermal Strain for A380 alloy at 750 C Figure 111. Thermal strain for A380 alloy at 800 C Figure 112. Thermal strain for AT72 alloy at 675 C Figure 113. Thermal strain for AT72 alloy at 710 C Figure 114. Thermal strain for AT72 alloy at 750 C Figure 115. Shrinkage porosity of A206.2 alloy at 700 C Figure 116. Shrinkage porosity of A206.2 alloy at 760 C Figure 117. RDG criterion for die casting alloys Figure 118. Feurer s lamda model for Al-Si alloy [69] Figure 119. Hot cracking susceptibility for Mg-Al alloy [14] Figure 120. Fluidity of AZ91 alloy as a function of Sn content [72] xviii

20 Figure 121. Proposal for future research work xix

21 CHAPTER 1: INTRODUCTION Hot tearing (or hot cracking) is the formation of irregular cracks in a casting that develop during solidification and cooling; typically while the casting is still inside the mold or die cavity. The cause of hot tearing is generally attributed to the development of thermally induced tensile stresses and strains in a casting as the molten metal contracts during solidification and solid state shrinkage. As a result, hot tearing often occurs at inside corners or fillets of casting geometries, where casting shrinkage is constrained by the mold cavity. In die casting, the mold cavity is a comparatively rigid structure (usually steel), in compared to the relatively low strength aluminum, magnesium or zinc casting alloys at high temperature. One of the key castability attributes considered in the development of new metal casting alloys is a low propensity for hot tearing. In addition to casting design features, factors that influence hot tearing include both casting alloy (chemical composition and solidification characteristics), and casting process parameters. Therefore, hot tearing tends to be of greater concern in die casting processes, compared to sand casting processes where the mold cavity is typically lower in strength, and more compliant to casting shrinkage. Hot tearing cracks can be seen via in Constrained Rod Casting (CRC) shown in Figure 1 [1, 2]. 1

22 Figure 1. Hot tearing at junction points of constrained rod casting Also, hot tearing defects have been investigated in the continuous casting of steel and aluminum ingots. The rapid cooling processes cause higher thermal gradients that led to thermal contraction in the ingots, which result in hot tears as shown Figure 2 [3]. 2

23 Figure 2. a) Hot tearing in aluminum ingot b) Hot tearing in extrusion billet [3] Moreover, hot tearing exists in the other metal forming processes such as high pressure die casting and welding. In Melt Conditioned High Pressure Die Casting (MC-HPDC), the hot cracking phenomenon was investigated in the recycled AM series magnesium scrap by Tzamtzis et al., as shown in Figure 3 [4]. The MC-HPDC process has shown excellent potential as a physical recycling technology high grade Mg-alloy scrap. Intensive shearing was used to optimize the MC-HPDC process in order to eliminate the casting defects. To optimize the process, the parameters were altered by reducing the die filling time, changing the intensifier position, changing the die temperature to 180⁰C and melting temperatures range to ⁰C. Grain refinement and morphology changes were accomplished and as a result, the hot cracking was eliminated. 3

24 Figure 3. Shoulder cracks in the MC-HPDC recycled AM series magnesium scrap prior to optimization [4] Susceptibility to hot cracking in weldments of 230 super alloys was performed by Cheng et al. The specimen was welded by Gas Tungsten Arc Welding (GTAW) and Plasma Arc Welding (PAW) and then heat treatments were applied in two ways for stress relief: (1) rapid heating to 1245 C and maintain at the same temperature for 30 min, and then quenching in water (2) rapid heating to 1120C and maintain at the same temperature for 30 min, and then quenching in water. It was observed that all these cracks were in micro-sizes and parallel to grain boundary as shown the microstructure of HAZ (heat affected zone) in Figure 4 [5]. 4

25 Figure 4. Hot cracks in Heat Affected Zone (HAZ) of gas tungsten arc welding From previous studies, there have been several phenomena linked to hot tearing as shown in Figure 5. However, these two include lack of feeding, and tensile stress and strain in the solid state. First, when liquid flows into mushy zone and after dendrite formation, in the localized area, liquid becomes isolated and can no longer be compensated by the flow of liquid. Secondly, tensile stress and strain is caused by volumetric changes, which produced thermal contraction that led to hot tearing [6,7]. 5

26 Figure 5. Dendritic solidification along with possible hot tearing phenomena in casting [6] 6

27 1.1 HOT TEARING EVALUATION METHODS There are several methods that have been employed to evaluate hot tearing propensity of die casting alloys that include Constrained Rod Casting (CRC), Crack-Ring, Horizontal Bar, T-Shape, and Ring Mold. The CRC, Horizontal Bar, and Ring Mold Methods typically employed in the gravity die casting process. CRC has evolved as the most common method, and has been used in several studies to evaluate hot tearing for die casting alloys. The design of a typical CRC test mold consists of rods of varying lengths with each rod having a bulbous or restraint ball at the end as shown in Figure 6. Figure 6. Experimental set up of constrained rod casting [8] 7

28 Upon casting, the molten metal fills the runner, rods, and rod ends. As cooling, solidification and shrinkage occur, tensile stresses and strains develop along the length of the rods because of the constraint provided by the solidifying vertical runner bar and bulbous rod ends. The longer the rod, the greater the stresses and strains that develop. The hot tearing propensity of casting alloys is evaluated using these rods at various pouring temperatures as hot tearing is visually observed shown in Figure 1. The drawbacks of the CRC method are that results are not universally applicable, because there is no standardized CRC mold design (dimensions or number of rods), and the test results are semi-quantitative at best [8]. There is currently a great deal of effort being expended on the development of new light metal alloys (aluminum and magnesium based) for high pressure die casting of automotive, sporting goods, electronics, and aerospace components. However, evaluation of the hot tearing characteristics of new alloys using the CRC method does not yield a good quantitative measure of hot tearing propensity, and the CRC testing methodologies and mold designs published in the literature are inconsistent. Furthermore, there is no evidence that any engineering design has been employed in the design of the CRC molds described in the literature. 8

29 To address these issues, the objectives of this research are first to design and develop a new CRC mold (Enhanced Constrained Rod Casting) based on the engineering principles for molten metal flow in die casting alloys. The second objective is to conduct experimental studies on new light metal alloys (aluminum and magnesium based) using the Enhanced Constrained Rod Casting mold. 1.2 SCOPE OF RESEARCH Many researchers, which include Argo et al. [1], Cao et al. [2], Zhou et al. [3] etc., have used CRC as a method to evaluate the hot tearing propensity of die casting alloys. The fundamental issues with current liquid metal feeding concepts are when the liquid metal drops through a straight vertical sprue, the air is aspirated and often oxidized in contact with mold material (steel), which is different from liquid metal in a tapered downsprue. In addition to that, abrupt changes in direction of liquid metal flow at the bottom of the sprue causes turbulence (unsteady) and air aspiration resulting in formation of inclusions, as well as slowing the velocity of liquid metal as shown in Figure 7. The unsteady flow might develop bubbles that could promote porosity and other casting defects [7]. Cao et al., Li S., and others have used threaded steel screw to measure the contraction forces with a load cell during solidification of CRC mold casting. However, the concern is that steel has a high coefficient of linear thermal expansion, which varies at elevated temperatures. As liquid metal solidified at varying temperatures, the measured forces 9

30 differ due to thermal expansion with a steel threaded screw during contraction and shrinkage. These uncertainties in measurement of forces have caused some unreliability in previously collected force data [8,9,10]. Figure 7. Air entrapment and unsteady flow in CRC 10

31 1.3 RESEARCH OBJECTIVES The objectives of the current research are as follows: Design and manufacture a mold (Enhanced Constrained Rod Casting-ECRC) for hot tearing evaluation method that fills all rods as concurrently as possible to minimize the heat loss and avoid air entrapment during mold filling. It also solidifies at uniform cooling rate. Set up an experimental method to study hot tearing of different materials that include - A206.2, Test-A, A380, and AT72 alloys. The selection of A206.2 and A380 alloys were chosen as baseline alloys because there were existing hot tearing work reported in the literature. A206.2 is known to have high propensity to hot tearing. A380 alloy is widely used to produce automotive parts and considered to be good alloy in terms of hot tearing propensity. A206.2 and A380 alloys will validate the ECRC mold design. Test-A (Aluminum) and AT729 (Magnesium) alloys are new alloys that had never been tested for hot tearing before. Develop a new measurement technique and replace threaded steel screws has high coefficient of linear thermal expansion 7.3 (10 6 in/in F). Find rod that has low coefficient of linear thermal expansion to measure thermal contraction forces during contraction and shrinkage. 11

32 Run experiments on alloys A206.2, A380, Test-A, and AT72 and measure temperatures and contraction forces during solidification phases. Evaluate and analyze data to understand hot tearing characteristics of A206.2, Test- A-alloy, A380, and AT72. Investigate influence of temperatures at various pouring temperatures for A206.2 A380, Test-A-alloy, and AT72, keeping mold cavity at room temperature (25.5 C). Characterize hot tearing phenomena which vary for different alloys from the graph of cooling curves and contraction forces, which are a function of time. Develop numerical models for CRC and ECRC (proposed enhanced constrained rod casting) to optimize the flow and perform comparative analysis using MagmaSoft. Also, perform further simulation on ECRC model to study porosity shrinkage, hotspot, and micro-porosity. Develop a Finite Element model of ECRC mold and perform thermal and structural analysis to predict total thermal strain in casting using ANSYS. Evaluate existing hot tearing predictive models that include Niyama criterion, RDG (Rappaz Drezet Gremaud) criterion, and Clyne and Davies criterion to study shrinkage porosity, thermal strain, and Hot Tearing Susceptibility (HTS) index to validate experimental studies. Use hot tearing predictive models and numerical analysis and simulation results to validate experimental studies. 12

33 1.4 RESEARCH METHODOLGY The following methodologies are pursued in order to achieve the research objectives. A schematic illustration of methodologies is illustrated in Figure 8. Develop analytical and numerical models for CRC (constrained rod casting) and ECRC (proposed enhanced constrained rod casting) to design and optimize the flow and perform comparative analysis using MagmaSoft. Develop a solid model of newly designed casting part. Create left and right mold cavities of new part and engineering drawings for both cavities. Manufacture new book mold assembly of P20 tool steel to evaluate hot tearing of A206.2, A380, Test-A (Aluminum), and AT729 (Magnesium) die casting alloys. Develop an experimental plan and equipment for experimental studies. These consist of a book mold, load cell, K-type thermo-couples, control-switch, quartz rod with grooves, bracket, temperature data logger, and data acquisition system. Quartz rods with grooves have been selected for measuring thermal contraction forces, which has a very small coefficient of linear thermal expansion of (10 6 in/in F) compare to threaded steel rod of 7.3 (10 6 in/in F)[11]. Setup experimental devices to measure thermal contraction forces and temperatures with respect to time during solidification phases of liquid metal cooling. 13

34 Run experimental studies of die casting A206.2, A380, and Test-A (Al) alloys at pouring temperatures of 700 C, 750 C, and 800 C. Perform experimental studies of AT72 die casting alloy at pouring temperatures of 675 C, 710 C, and 750 C. Evaluate cooling temperatures and thermal contraction forces data as function of time for A206.2 A380, Test-A-alloy, and AT72. Also, characterize the hot tearing predictive models for shrinkage porosity and thermal strain. Examine cast parts for hot cracking and perform comparative analysis on measured data to predict hot tearing characteristics. Develop a Finite Element model of ECRC mold and perform thermal and structural analysis simulation to predict total thermal strain in die casting of A206.2, A380, Test- A-alloy, and AT72 alloys using ANSYS. Develop a finite element model of constrained rod casting (CRC) and Enhanced constrained rod casing (ECRC) using MagmaSoft in predict fluid flow velocity, hot spot, filling time, solidification times and porosity defects. 14

35 Figure 8. A schematic illustration of methodologies for hot tearing evaluation methods 15

36 1.5 ORGANIZATION OF THE DISSERTATION The dissertation is divided into 6 chapters and an appendix. Chapter 1 introduces of hot tearing and research approach. Chapter 2 describes literature research and previous research work, which covers hot tearing evaluation methods, influence of Cu, Si, Fe, and grain refiners in aluminum and magnesium die casting alloys. It also covers about the thermo-mechanical modeling and simulation of die casting alloys. Chapter 3 covers the theoretical background of hot tearing formation and characteristics of hot tearing in casting. Chapter 4 describes the development of book mold assembly, hot tearing evaluation methods, and experimental studies in predicting hot tearing of A206.2, A380, Test- A-alloy, and AT72 alloys. Chapter 5 covers the thermo-mechanical modeling of enhanced constrained rod casting (ECRC) using ANSYS. It also covers finite element modeling and simulation of constrained rod casting (CRC) and enhanced constrained rod casting (ECRC) using MagmaSoft to perform comparative analysis. Chapter 6 includes the conclusions and future research possibilities for improvements. Appendices include all of the supporting materials include drawings and data from experimental studies that were used in this research. 16

37 CHAPTER 2: LITERATURE REVIEW This section summarizes the information published during the past years about hot tearing defects in die casting of aluminum and magnesium alloys. Included is a brief description of hot tearing, and an overview of experimental methods for evaluating the hot tearing propensity of alloys, process-related factors reported to influence hot tearing, and alloy characteristics which affect hot tearing. Additionally, numerical and analytical models proposed for evaluating and predicting hot tearing are reviewed. 2.1 EXPERIMENTAL METHODS FOR EVALUATION OF HOT TEARING Several methods have been employed to experimentally evaluate the hot tearing propensity of die casting alloys. These method include constrained rod casting (CRC), crack-ring, horizontal bar, T-shape, and ring mold. The CRC, horizontal bar, and ring mold methods typically employ the gravity die casting process. The crack ring and T-shape methods discussed utilized sand molds. The CRC method is often used for evaluation of die casting alloys [11, 12]. An overview of the alloys evaluated for hot tearing (by author), and the experimental methods employed, follows. 17

38 2.1.1 CONSTRAINED ROD CASTING (CRC) FOR EVALUATION OF Mg-Al ALLOYS Argo et al. [13] investigated casting process parameters of AJ52 high temperature Mg-Al- Sr alloy using constrained rod casting (CRC). The investigation included castability index, creep resistance, and hot tear ratings for the Mg-Al alloys shown in Table 1. The authors reported that the main factors that affected the castability were: freezing range (difference between liquidus and solidus temperatures on the phase diagram), specific heat, fluidity, wettability, and thermal conductivity of the alloy. The freezing range of the alloys considered had a greater positive influence on castability than other casting characteristics. However, it was discovered that longer freezing range alloys may be more susceptible to hot tearing than alloys having a short freezing range. In addition to freezing range, another factor they reported as influencing hot tearing is the amount of eutectic present at grain boundaries. It was observed that AZ91D alloy had longer freezing range and less susceptibility to hot tearing than AJ50x alloys with smaller freezing range. 18

39 Table 1. Castability Index for Mg Die-casting Alloys [13] An oil pan and valve cover were die-cast of AJ52x alloys for thin wall (< 2mm), medium wall (5 mm) and thick wall (> 5 mm) using a cold chamber machine. Hot tearing was exhibited on the thin wall (< 2mm) wall cover of AJ50x and AJ51x alloys. No hot tears were found on the medium wall and thick wall (> 5 mm) valve cover of AJ52x. Considering all of the casting characteristics, the castability index was developed for thin walled, medium walled, and thick walled castings as shown in Table 1. Figure 9 shows one of the constrained rod casting (CRC) experimental castings and the associated hot tears at the transition from the runner bar to the rods (center and right of the photo) used by Argo, et al. A larger overview of a typical constrained rod casting die cavity is shown in Figure 10. The design of a typical CRC cavity consists of rods of varying length, with each rod having a bulbous end larger in diameter than the rod. Upon casting, the cavity fills the runner, rods, and rod ends. As cooling, solidification and 19

40 shrinkage occurs, tensile stresses and strains develop along the length of the rods because of the constraint provided by the solidifying runner bar and bulbous rod ends. The longer the rod, the greater the stresses and strains that develop. The hot tearing propensity of an alloy is evaluated relative to the CRC rod length at which hot tearing can be observed. Alloys that demonstrate hot tearing even in the short rods have a high propensity for hot tearing; those that only hot tear at the longest rod lengths are considered to have a low propensity for hot tearing. There does not appear to be a standard geometry for CRC mold cavities. Therefore, the CRC test is typically a comparative test only; it is not numerically absolute, or quantitatively very precise. Figure 9. Hot tearing in the longest rod [13] Cao et al. [14] also studied the hot tearing of Mg-Al alloys using the gravity die casting process with a steel mold for constrained rod casting (CRC). In addition to the conventional CRC die cavity, this tool was cleverly instrumented with a load cell and 20

41 thermocouples to obtain a cooling curve (temperature vs. time) and constrained rod load (load vs. time) data. Magnesium-Aluminum alloys ranging from 0.25 to 8 wt% Al content were used to study the hot tearing effects on binary Mg-Al alloys. Hot tearing susceptibility occurred sharply with Mg-1wt% Al alloys and as contents of Al alloy increased, the susceptibility decreased significantly and did not exhibit any hot tearing beyond 8 wt% Al. Boron Nitride (BN) coating of 100 μm reduced the hot tearing susceptibility significantly compare to BN coating of 40 μm. Figure 10. Steel mold for constrained rod casting to determine hot tearing [14] Zheng et al. [15] investigated a quantitative method for determining the hot tearing in Mg- Al binary alloys. This study is similar to Cao et al. and constrained rod casting was employed. The steel mold was coated with Boron Nitride (BN) and preheated at temperatures between 250 and 500 C. The pouring temperature was kept 80 C above the 21

42 liquids temperature. The contraction stresses induced during solidification shrinkage were evaluated during the CRC trials. Hot tears observed were correlated to the plots of contraction forces and temperatures as a function of time. Hot cracks were measured by a wax penetration method, and crack size was determined using volume of total cracks. For Mg-1wt% Al alloy, when mold temperatures were increased, the hot tear crack sizes decreased. When mold temperature reached 500 C, no cracks were found. For Mg-3wt% Al alloy, no cracks appeared at mold temperature of 350 C. There were no cracks for Mg-9 wt% Al at mold temperature of 250 C. Higher cooling rates developed larger temperature gradients over uneven shapes of the casting, which promoted thermal stresses resulting in hot tears. Zhou et al. [16] evaluated the influence of Zn on hot tearing susceptibility of Mg-Zn binary alloys and effects of the friction between mold walls and constrained rod casting. The pouring temperature was used 80 C above liquidus temperature. A thin layer of boron nitride coating was used before pouring. The mold temperatures were used from 200 to 550 C. It was observed that varying mold temperatures and content of Zn had considerable effects on hot tearing as shown in Table 2. The Mg-6 wt% Zn did not exhibit hot tearing at mold temperature of 450 C. As the content of Zn increased, no cracks were found at mold temperature of 450 C. Higher the cooling rates mean lower the mold temperature. These authors concluded that at higher cooling rates, alloys will be more prone to hot tearing [16]. 22

43 Table 2. Influence of Zn on Mg-Zn Alloy at Different Mold Temperatures [16] Kamga et al. [17] investigated the hot tearing of aluminum-copper B206 alloys with iron and silicon additions using constrained rod casting (CRC) with nominal length of 50.8 mm (Bar-A), 88.9 mm (Bar-B), 127 mm (Bar-C), and mm (Bar-D). Also, the purpose was to understand the characteristic of B206 with higher iron contents and combined effects of iron and silicon. The B206 alloys was modified using Al-1020 and commercial master alloys (Al-50%Si, Al-25%Fe,Al-25%Mn, Al-50%Mg, and Al-50%Cu). After chemical analysis using optical emission spectrometer, the compositions are shown in Table 3 for experimental studies [17]. 23

44 Table 3. B206 Alloys with Various Compositions [17] In this experimental work, the constrained rod casting (CRC) mold was made of cast iron. For each operation, the mold was cleaned and heated up to 200 C and coated with graphite. In the casting alloys, Al-5wt%Ti and 1wt%B were added as grain refinement. The melting temperatures of alloys were maintained at 750 C. Castings were removed from the mold in temperature ranging 405 to 410 C. Three castings were produced for each alloy. Unaided visual and microscopic inspections were performed for hot tears. Hot tears were categorized based severity into four parts with assigned number for visual inspections, these include surface tear (1), light tear (2), sever tear (3), and complete tear (4). A hot tearing sensitivity (HTS) index was used along with visual inspections index to determine the level of tears and is defined as [17]: 24

45 HTS C / L D i A i i D i A 4/ Li (2.1) Where, C =Hot tear severity values that are assigned using index i=a,b,c, and D i HTS=1 means all the bars have cracks and HTS=0 means no bars have cracks It was observed that hot tearing was influenced by the iron content in the situations where there were coarse grains. Alloys with more than 0.01 wt% Ti showed fine grains microstructure; alloys with less than 0.01 wt% Ti showed coarse microstructure. The longest bar exhibited cracks; there were no cracks on two shortest bars [17] CRACK-RING FOR EVALUATION OF MG-9AL-XZn ALLOY Wang et al. [18] used a crack-ring mold cavity design in sand casting to investigate hot tearing in Mg-9Al-xZn alloys as shown in Figure 11. They discovered that zinc additions decreased the solidus temperature and increased hot tearing susceptibility coefficient (HSC) in alloys. The contents of Al and Zn lowered the melting point and alloys solidified at eutectic stage. They reported that the grains could not form a dendrite network due to low amount of remaining eutectic liquid. This promoted lack of strength during solidification shrinkage and promoted the development of the thermal stress-induced 25

46 cracking. Hot-tearing susceptibility increased at 0.8 wt% Zn (Mg 9Al 0.8Zn) and decreased when content exceeded 0.8 wt%zn. Wang et al. [19] further studied the effects of Zn and RE (rare earth elements) additions in Mg-9Al alloys using crack ring mold. They found that Zn additions lowered the solidifying temperature while RE additions had little effect. Hot tearing susceptibility (HSC) increased of Mg-9Al alloys as Zn additions with low quantity of RE content. Eutectic temperature decreased at grain boundaries due to Zn content [18, 19]. Figure 11. Crack-ring mold to determine hot tearing [18] 26

47 2.1.3 HORIZONTAL BAR FOR EVALUATION OF Mg ALLOY Bichler et al. [20] experimented with the onset of hot tearing in AE42 alloy. This alloy has high temperature strength compare to magnesium alloy (AZ91). AE42 magnesium alloy was poured into a gravity die casting mold that consisted of a down-sprue and a long horizontal bar as shown in Figure 12. The experiment was carried out using different mold (140, 220, 300, 340, and 390 C) and pouring temperatures (720, 740, and 765 C). The liquid metal started solidifying (directionally) from the horizontal bar through the downsprue and then into the mold cavity. The pouring cup solidified last. Hot tearing occurred for all mold temperatures below 300 C and pouring temperatures 720 C and 740 C at the transition between down-sprue and horizontal bar. Axial contraction from horizontal bar and down-sprue caused a resultant stress concentration at transitional point (90 C), where hot tearing was observed. As pouring temperature and mold temperature were increased to 765 and 390 C respectively, the hot tearing cracks vanished [20]. Bichler et al. [21], employed neutron diffraction techniques to measure the residual stresses and strain in magnesium alloy (AZ91D) at the onset of hot tearing. The model is used from their previous study shown in Figure 13. Pouring temperature was held at 720 C. Hot tearing formed at a mold temperature of 210 C (Figure 5, Part a). When mold temperature increased to 250 C, the casting was less susceptible to hot tearing, (Figure 5, Part b). Tensile residual strain was observed for (1 0 0) and (0 0 1) reflections, measured 27

48 by Neutron diffraction at mold temperature of 210 C. The mixed strains were recorded for (1 0 1) and (1 0 2) reflections. Similarly, Neutron diffraction strain mapping were observed at mold temperature of 250 C. The average strains were recorded for (1 0 0) and (0 0 1) reflections. These two average strain values for casting with hot tear and without hot tear are shown in Figure 14 [21]. Figure 12. A model of horizontal bar with downsprue [20] Figure 13. Hot tearing effects at mold temperatures 210 and 250 C [20] 28

49 Figure 14. Hot tearing at reflections (0 0 1) [21] T-SHAPE FOR EVALUATION OF A206 ALLOY Esfahani et al. [22] studied the hot tearing of A206 aluminum alloy using an instrumented constrained T-shaped casting (ICTC) method, see Figure 15. Four pouring temperatures (675, 700, 750, and 800 C) were selected in this study. The cooling curve (temperature vs. time) was used to analyze the events such as liquidus, solidus, and coherency, etc. during the solidification. Tear formations were displayed with a Load-time curve. At pouring temperature of 800 C, a change in slope was observed between two curves T-t and dt/dt-t, as shown in Figure 16. Hot tears were observed initially at pouring temperature of 700 C and largest cracks were noticed at 800 C at the T-junction of the casting. 29

50 Figure 15. Schematic of instrumented constrained T-shaped casting [22] Figure 16. Cooling curves T and dt/dt-t, hot tearing at 800 C [22] RING MOLD FOR EVALUATION OF AA1050 VIA ACOUSTIC EMISSION Pekguleryuz et al. [23] performed Investigation of hot tearing in aluminum alloy AA1050 via Acoustic Emission (AE) and cooling curve analysis methods using a ring mold. The experimental set up for a ring shaped mold is shown in Figure 17. AE method is an approach which can detect the hot tearing using elastic waves. When a specimen is in non- 30

51 equilibrium, it goes through deformation and results in releasing the elastic strain energy, which is detected as stress waves. The AE sensor converts the stress wave into a voltage proportional to the magnitude of the stress wave. Also, during plastic deformation of materials, there are many stress waves moving at the same speed, which superimposes (amplifies) these waves, and then it is easy to detect by the AE sensor. Hot tearing information was collected via AE and cooling curve analysis. The aluminum alloy AA1050 has liquidus and solidus equilibrium temperatures of 659 C and solidus 630 C, respectively. The non-equilibrium liquidus ranges from 652 to 659 C. The solidus temperature was 630 C. The non-equilibrium freezing range varied from 43 to 99 C. Hot tearing initiated where the AE energy was over 600 energy units (e.u.) and frequency range was from 110 to 140 khz in zone II. Hot cracking occurred when AE energy was over 650 energy units and average frequency range was from 111 to 145 khz in zone III. Hot tearing occurred at temperature ranging from 636 to 653 C. The solidus ranged from 556 to 614 C as shown in Figure 18. The fraction solid at the onset of hot tearing ranged from 0.71 to Addition of Fe in alloys slightly increased hot tearing. 31

52 Figure 17. Experimental set up of ring mold casting [23] Figure 18. Hot tearing detection via AE method [23] 32

53 2.1.6 RING MOLD FOR EVALUATION OF Al-Cu AND Mg ALLOYS Figure 19 shows the results obtained from ring mold testing for Al-Cu alloys. It indicates that hot tearing susceptibility appeared to maximum at 0.5 % wt Cu. Smaller crack length had 20 C of melt superheat. The larger crack length had 100 C of melt superheat. Hot tearing susceptibility occurred to be a maximum at 3.5% wt. Cu content [24]. Figure 19. Crack length as function of Cu content in Al-Cu alloys [24] Results from a ring mold test are presented in Figure 20. In Al-Mg alloys, the maximum value for hot tearing appeared at 1% Mg content. Due to melt superheating, the larger cracks were found at 100 C and showed crack length as a function of magnesium content [24]. 33

54 Figure 20. Crack length as function of Mg content in Al-Mg alloys [24] From Figure 19 and Figure 20, it can be concluded that as wt% of Cu and Mg content increases, the hot tearing reaches a maximum and then starts decreasing. Alloys after certain wt% will have smaller solidification range and more liquid available to feed the high volume faction of solid. Generally, smaller solidification range will have less thermal contraction in casting. As a result, a casting with better microstructures will be produced [24]. 2.2 NUMERICAL MODELING APPROACH AND SIMULATION Sengupta et al. [25] studied hot tearing defects in ingot using ABAQUS software. They simplified the model to a quarter section of geometry because of symmetry for simulation of thermal and stress analysis. The 3D model consists of an ingot and a bottom block. For the mesh of model, an 8-nodes gauss integration brick element was used. The element 34

55 sizes were carefully selected to provide constant thicknesses and refined mesh to the model. Also, incremental time intervals have been used for casting speed and mold filling rates. Since the casting process is time dependent, the heat transfer model considers the transient and the varying temperature effect on the thermal and physical properties of alloys. In this model, it is assumed that there is no flow of liquid metal to fill the mold but heat energy is transferred through diffusion method. The nodes are set to solution domain set at initial temperature of alloys. Figure 21 shows the 3D model with mesh and boundary conditions for thermal analysis. The model has predicted temperature and displacement measurements obtained from two 711 mm 1680 mm AA5182 ingots, cast under different start-up conditions. These conditions are produced by varying the bottom block filling rate and flow rate to achieve a non-typical cold and a non-typical hot. The model predicted that both hot and cold cast had the variation in plastic strain and stresses with respect to time. These two casts developed tensile stresses due to impingement of the secondary cooling. Accumulation of tensile plastic strains could form pores that lead to hot tearing. Future work is recommended such as mesh refinements for optimization and elastic and plastic strain behavior with respect to time [25]. 35

56 Figure 21. A 3D model of ingot with bottom block and boundary conditions [25] Lin et al. [26] discussed predicting hot tearing in steel using a damaged porosity based model concept. In this study, MAGMA software was used to calculate the temperature and feeding results (Porosity). These data were incorporated in FEM model as shown in Figure 22 about the flow chart and details of the approach. Figure 22. Hot tearing modeling approach [26] 36

57 The hot tear is predicted by using damaged porosity equation g p,d t = g t s [ε xx + ε yy + ε zz]dt (2.2) f Where, g s = solid fraction, ε xx, ε yy, ε zz= visco-plastic strain rates, and t f = feeding cut off time. The above equation can calculate the volume fraction of porosity based on visco-plastic model. The solid fraction cut off is The mold material is important factor to consider in steel casting since the casting interacts with the mold and may apply resistance during contraction. The model used surface elements (SANDSURF) to replace the sand mold. The surface elements transfer normal forces to casting surface and predict the displacement and plastic strain. The damaged porosity model can predict hot tearing at some locations but it did not fully correlate with cracks that were found at actual casting. See below hot cracks in casting and modeling simulation results in Figure 23 [26]. 37

58 Figure 23. Comparison between hot tear in casting and simulation results [26] Ridolfi et al. [27] studied the formation of cracks in steel casting using software MSC Marc. A thermo-mechanical model of 2D was developed based on traveling slice method. Quadrilateral plain strain elements and eight-nodes were used in modeling and analysis. The model was divided into mold domain and steel domain. Due to symmetry of the part, one quarter of Model with boundary conditions and mesh are presented in Figure 24. The analysis predicted the temperature distribution per cooling rate, stress, and strain during the solidification process [27]. Figure 24. A 2D model with mesh and boundary conditions [27] 38

59 2.3 HOT TEARING PREDICTIVE MODELS Feurer [28] proposed a hot tearing criterion model to predict tear formation in the casting. He investigated the feeding characteristics and hot tearing properties and found that the poor feeding is caused by solidification shrinkage because of the interlinking of dendrites. However, the dendrites act as porous filters. The residual liquid is fed through these dendrites porous filters. Insufficient feeding to the casting may promote the volumetric shrinkage. If the shrinkage velocity exceeds the maximum flow rate of feeding, hot tear may occur. However, Feurer postulates that hot cracking susceptibility (HCS) is possible if [28]: SPV SRG (2.3) The maximum volumetric flow rate per volume is defined by SPV = 1 V ( V t ) = ( lnv t ) (2.4) Where, SPV= the maximum volumetric flow rate through dendrite networks, V = Volume, t = time The volumetric solidification shrinkage is caused by density difference between liquid and solid. It is defined by SRG = ( lnv ) = 1 ρ t ρ t (2.5) 39

60 Where, SRG = Volumetric shrinkage during solidification, ρ = ρ L g L + ρ S (1 g L )V V= volume of solidify elements, ρ L = density of the liquid, ρ S = density of solid phase Clyne and Davies [29] defined the hot cracking index as a ratio of a vulnerability time (t V ) when hot cracking may develop (liquid fraction from 0.01 to 0.10) to the time of stress relief time (t R ) during mass feeding (liquid fraction from 0.1 to 0.6) as shown in Figure 25. The hot cracking susceptibility (HCS) index is defined [29, 30]: H. C. S. = t V t R = t 99 t 90 t 90 t 40 (2.6) Where, (t V ) = the vulnerable time period when cracks can propagate, (t R ) = the time period for stress-relaxation during liquid feeding, f S = volume fraction of solid, t 99 = time at f S is 0.99, t 90 = time at f S is 0.90, t 40 = time at f S is 0.4. Figure 25. Determination of hot cracking susceptibility [30] 40

61 Katgerman [31] continued work on the hot cracking index based on Feurer [28] and Clyne et al. [29]; if after feeding is inadequate from meniscus, then results of volume reduction could cause stresses that promote cracks. The vulnerable time proportionality was changed from t 99 t 90 to t 99 t cr. Hot cracking (HC) index defined as [31]: H. C. = t 99 t cr t cr t 40 (2.7) Where, t cr = time at feeding becomes inadequate, t 99 = time at f S is 0.99, t 90 = time at f S is 0.9, t 40 = = time at f S is 0.4. The hot tearing index improved using coherency temperature defines where dendrite network begins to form. Considering the coherency temperature (t Coh ), then hot cracking index is defined by: H. C. = t 99 t cr t cr t Coh (2.8) Where, t Coh = coherency temperature, t cr = time at feeding becomes inadequate, t 99 = time at f S is Hatami et al. [32] proposed new criteria based on theories of Feurer [28], Clyne et al. [29], and Katgerman [31]. The new theory considers a volume element in mushy zone, called volume solid fraction for hot cracking simulation. If there is no flow considered, then the 41

62 zero flow point in mushy zone is known as solid fraction for rigidity. This new criteria postulates that if a neighboring element has a solid of fraction less than the rigidity point, the local liquid feeding is possible and no cracks will occur. And if they exceed the solid fraction of rigidity, then liquid feeding is not possible and hot cracks may occur. Such hot cracking susceptibility (HCS) is determined on each element: H. C. = t 99 t cr t 99 t cr (2.9) Where, T 99 T cr = critical time at which element transmits stresses, T cr = time at feeding becomes inadequate, T 99 = time at f S is Campbell [33] suggested a modified version of hot tearing criterion for susceptibility as the product of accumulated thermal strain at a casting hot spot and ratio of vulnerability to stress relief. It was only suitable for qualitative measurement in different alloys as defined below. CSC = α TLa l 2 t V t R (2.10) Where, α= Coefficient of thermal expansion, T= undercooling temperature, L= Casting length, a= grain size, l= Length of hot spot, α TLa =Accumulated thermal strain at hot spot l 2 t V t R = Ratio of vulnerability (t V ) to stress relief (t R ) 42

63 Kamga et al. [34] modified the hot tearing index proposed by Katgerman [31]. The new hot tearing index is defined as follows [34]: H. C. = T cr T 0.01 T Coh T cr (2.11) Where, T Coh = dendrite coherency temperature, T cr = = temperature below at feeding becomes inadequate, T 0.01 = = temperature at volume fraction liquid to The Niyama criterion was developed by E. Niyama in 1982 to predict shrinkage porosity. The Niyama is defined as [73] N y = G T (2.12) Where G is the thermal gradient and T is the cooling rate. Niyama criterion is based on strain rate. If a casting is going under applied tensile strain, the pressure drop in mushy zone is increased and then pores will grow without shrinkage and without deformation of casting. This shrinkage porosity will be perpendicular to applied strain which will promote hot tearing. Carlson et al., investigated shrinkage Porosity of WCB steel and correlated simulation results from MagmaSoft with experimental studies. AZ91D simulation showed shrinkage porosity. This demonstrates that Niyama criterion can be applicable to a wide variety of alloys in predicting shrinkage porosity [74]. 43

64 In RDG (Rappaz Drezet Gremaud) criterion, deformation is perpendicular to thermal gradient regardless whether it corresponds to columnar or equiaxed dendrites. Hot tearing will not occur if mushy zone can sustain some deformation and its strain rates remain low enough in order to permit liquid feeding. In the RDG hot tearing criterion [75], the depression pressure, p, over the mushy zone: p = psh + pmec ρgh (2.13) Where, psh and pmec are the pressure drop contributions in the mushy zone associated with the solidification shrinkage and the deformation caused by fluid flow, respectively, ρ is density, g is gravitational constant, and h is the distance below the liquid melt level. Hot cracking susceptibility (HCS): = 1 ε max (2.14) Drezet et al. used RDG criterion to determine HCS for welding of aluminum alloys. Influence of filler content on HCS values and solidification path determined. However, the disadvantage is that this method does not account for development of localized strain at microscopic level [75]. 44

65 2.4 SUMMARY AND DISCUSSION Hot tearing susceptibilities of the casting alloys investigated experimentally are summarized in Table IV, listed sequentially based on evaluation techniques. Casting characteristics such as specific heat, thermal conductivity, longer freezing range, pouring temperature, and mold temperatures influence the hot tearing. Hot tearing susceptibility for Mg-Al increased at 1wt% Al due to the development of larger temperature gradients in constrained rod casting (CRC). As the mold temperatures increased, the crack sizes decreased smaller temperature gradient. Similarly, for Mg-Zn alloys, the hot tearing rose sharply at 0.8 wt% Zn and decreased when content exceeding 0.8 wt% Zn. For AZ91D alloys, the hot tearing formed at 210 C of mold temperature. However, at 250 C of mold temperature, casting was less susceptible to hot tearing. A206 alloys had largest crack at pouring temperature of 800 C. 45

66 Table 4. Hot tearing Susceptibilities to Die Casting Alloys 46

67 Eskin et al. [35] summarized some of the measured hot tearing susceptibilities of Al Cu, Al Mg, Al Si, Al Fe, Al Mn, and Al Zn alloys as shown in Table V. Most alloys at a certain composition have a tendency of maximum hot cracking. This phenomenon is represented by λ (lambda) curve as shown in Figure 26. For Al-Si alloys, hot tearing rises to a maximum at 0.7 wt% Si [35, 36]. Table 5. Hot tearing Susceptibilities [35] 47

68 Figure 26. λ (Lambda) curve [36] Argo et al. [13] discussed the freezing range that influences solidification processes. From the experimental studies it was found that freezing range of AZ91D was the longest compare to AM50 as shown in Table 6. Mg-Al-Sr had larger freezing range than AM50A. Because longer freezing range promotes hot tearing due to surface tension and eutectic present in dendrite network. In general, an alloy with longer freezing range may be prone to hot tearing than alloy with shorter freezing range. However, experimental results demonstrated that AZ91D had lower hot tearing rating than AM50A. A380 alloy had the lowest hot tear rating [13]. 48

69 Table 6. Hot tearing Ratings [13] It was observed from hot tearing evaluation studies that hot tearing is typically experienced at corners or fillets of castings. This is caused by development of localized strain, which is the result of combined thermal contraction from down-sprue and longest bar or horizontal bar (in CRC experiments). For complex or irregular shape of geometries, higher cooling rates may develop larger temperature gradient, which generates a greater propensity for hot tearing. Experimental studies evaluating the hot tearing propensity of die casting alloys in the high pressure die casting process were not found. Models developed for predicting hot tearing lend insight into the process conditions and alloy characteristics which influence hot tearing. 49

70 CHAPTER 3: THEORETICAL BACKGROUND In this chapter, the theoretical background of hot tearing formation, characteristics of hot tearing, and strain theory that develop in the casting are presented. 3.1 CHARACTERISTICS OF HOT TEARING Hot tearing is a casting defect that occurs near solidus temperature at late film stage, where the liquid films are still present at grain boundaries despite the completion of solidification as shown in Figure 28. Hot tearing is easily identified with one or more characteristics that are presented in Table 7 [37, 39, 42]. 50

71 Table 7. Characteristic of Hot tearing [37,39,42] Hot Tearing Visual appearance Surface failures by microscope Time Characteristics Forms a ragged, branching crack Main tear and its numerous minor off shoots generally follow the intergranular paths Reveals a dendrite morphology (see Fig. 9) Heavily oxidized with higher temperature alloys such as steel When the material is still at incoherent stage Close to completion of solidification Location It is often at a hot spot Where contraction strain from adjoining extensive thinner sections may be concentrated Temperature At or just above solidus temperature (Fig. 7) Highly specific to certain materials Defects Uniaxial tensile failure in weak material Feeding problem related to hydrostatic stresses causing pores in liquid phase. 51

72 Campbell (2003) described that dendrites open a pathway for draining of eutectic liquid, which had successfully formed the tear in Al-10Cu alloys. Figure 27 shows a scanning electron microscope view of the hot tearing surface of an Al 3% Cu alloy [37,39]. Figure 27. Hot tearing in Al-10Cu alloy (Spittle and Cushway 1983) [37,39]. Pellini (1952) used the radiographic method and thermal measurements to obtain the data in order to predict the hot tearing in steel casting of various carbon contents ranging from 0.03 to 1.00 percent. The casting (the length of casting was reduced to 16 from 24 to minimize the contraction) was poured in a restraining bar and sulphur or phosphorus content was added. High quantities were included to see the effects produced by these elements in the casting. It was observed that each element produced tears in the shortened casting, which began approximately 50-75⁰C below the solidus (Fe-C) temperature as shown in Figure 28 [37,41,42]. 52

73 Note: solid circles no tears, x hot tears Figure 28. Hot tearing at various carbon contents in Fe-C diagram [41]. 3.2 STAGES OF CRYSTAL GROWTH SOLDIFICATION PROCESS There are various stages in the solidification processes, as liquid metal cools, it starts with crystallization at liquidus temperature as shown in Figure 29. From the start of the crystallization process to a complete solid occurs in between liquidus and solidus temperatures. Immediately after the initial stage, as dendrite networks have been formed and crystals cannot move around in liquid metal, then early films are formed. As liquid metal cools down further, the late film stages are formed. At this stage, hot tearing 53

74 develops due to lack of feeding in some parts of castings. The tensile strength and elongation describe the effects of the resulting microstructure. The shrinkage starts decreasing from the early film stage to the late film stage. Hot tearing tendency at late stage of dendrite formation can be influenced and varied by die casting alloys and its composition. Graphical descriptions of this process are shown in Figure 29 and Figure 30 [31,44]. Figure 29. Start of crystallization, crystal growth, and film stages in solidification [44] 54

75 There are four stages of solidification. At stage 1, the primary dendrites are freely dispersed in liquid. Both dendrites and liquid are capable of relative movement. At stage 2, the liquid is moving freely at coherency stage between interlocking dendrites and liquid. Healing is possible if cracks develop. At stage 3, the grain boundaries are in an advanced stage of development, and this is considered to be at critical solidification range (semisolid), where the liquid is not moving freely. No healing of cracks is possible if significant strain developed in the casting. At stage 4, the casting alloy is fully solidified. The a-c represents the coherent temperature and a-e is critical temperature [43]. Figure 30. Stages of hot tearing in binary Alloys [43] 55

76 The material is susceptible to hot tearing once it has reached the coherency temperature at stage 2. However, further developments of microstructures prevent the free flow of liquid through interlocking dendrites. The accommodation of strains within mushy zone is not possible, since the healing could not occur with remaining liquid at stage 3, which is considered a critical solidification range. The relative movement of dendrite and liquid is not possible; it can only accommodate low tensile strain, see Figure 31. The other possibility is that the solidus temperature may depressed by undercooling and lack of diffusion, which will promote hot tearing in the casting [37, 38, 52]. Figure 31. Modeling of dendrites at coherency states [38] 56

77 3.3 FORMATION OF DENDRITE MICROSTRUCTURES Formation of a dendritic microstructure is characteristic of aluminum alloys as solidification takes place. Secondary dendrite arm spacing is known as dendrite arm spacing (DAS). Secondary dendrite arm spacing is one of the most important length parameters, other than grain size in the casting. One of the most important factors that may influence the casting is DAS. DAS determines the microstructures of cast parts. The more dendrite arm spacing per unit volume means better mechanical properties of alloys, see Figure 32. The secondary dendrite arm spacing is controlled by the coarsening process and cooling rate. During the coarsening process, the dendrite arms grow at the tip of the dendrites. The surface energy is reduced only if dendrites surface area is reduced. As a result, the small arms go into solution while the larger arms grow independently and increase the spacing between the arms. Finally, DAS is controlled by solidification time [37]. The relationship between DAS and local solidification time can be described as follows: n 2 K *t f (3.1) Where, is dendrite arm spacing in micrometer, t 2 f is local solidification time in seconds, K is a proportionality to coefficient and n is between For Al-Cu alloys, K is 7.5 and n is 0.39 respectively [1,3]. 57

78 Figure 32. Formation of grains and arm spacing [37] 3.4 STRESS-STRAIN RELATIONS AS A FUNCTION OF TEMPERATURE Flinn established the stress-strain relation as a function of temperature by considering a metal bar of aluminum (7% Cu, 2% Zn, balance Al), which is applied to a tension load starting from the liquid state. The phase of metal changes as temperature was falling, and five stages were observed from this investigation as shown in Figure 33 [42]. 58

79 Figure 33. Phase changes in stresses-strain curves at different temperatures [42] Step 1 Completely Liquid: The liquid metal follows the motion of end plate. Hot tear is not possible to form in liquid, since the cavity would be filled quickly. Any shrinkage would be healed by liquid. Step 2 Mostly Liquid with some solid: At this stage, the bar is under stress and may develop ruptures. Adequate liquid is available to fill those ruptures. Step 3 Mostly Solid with some liquid: Solid crystals start forming a network at a certain stage of solidification, which carry some strength. The material becomes coherent, where the solid content is about 50% to 90% depending upon crystallization growth. The casting ruptures under low stresses, while some liquid metal is still present. All these tears cannot 59

80 be filled by isolated patches of liquid metal. The first defect is encountered, which is called hot tearing. This occurs just above or at the solidus. Step 4 Solid (Plastic range): At this stage, the ductility is high and material deformation takes place at low stresses. No cracking takes place, unless it is brittle at this temperature. Another phenomenon is creep, if a steady load is applied to the bar, the grains will elongate as a function of time. Step 5 Solid (Elastic range): At this stage, as temperature is lowered, the grains elongate. Re-crystallization does not take place but flow stress increases, However, when transitioning from plastic to elastic, the temperature varies for different materials. Different temperature ranges are shown for stress-strain curves in Figure 16. Curve 1 and 2 are in liquid phase, therefore nothing is happening for stress and strain. Curve 3 represents the coherent temperature, where the metal ruptures at low stresses. Plastic behavior at low stresses is indicated by Curve 4. Finally, curve 5 shows the elastic deformation at high stresses [51]. 60

81 3.5 STRAIN THEORIES Pellini (1952) strain theory is based on the liquid films concept that exists at grain boundaries at, above, or in the region of solidus temperature. The strain theory provides a mechanism of hot tearing in terms of the time-rate of extension developed in the liquid film regions. It defines that total strain developed during film life period depends on strain rate and time of film life in the casting. Figure 34 described all possible forms of hot tearing including the critical amount of strain in the casting. The rate of extension may vary and depends on length of contraction, cooling rate, and width of hot spot extension. The effect of segregates is critical due to increased film life [41]. Figure 35 shows the various stages of casting in a solidification system that contains a hot spot. In case A, the hot zone is uniform in mushy stage where strain is weak and could cause separation. In case B, the hot zone is liquid film stage extension and highly concentrated in film regions. This results high strains which could cause separation. In case C, the hot zone in solid stage extension causes uniform creep of ductile solid metal [37,44]. 61

82 Figure 34. Strain development from film stage to non-equilibrium solidification [41] 62

83 Figure 35. Strain development leading to tear formations [41] 63

84 Pellini s theory can be quantified by assuming if the length of casting is L and it has coefficient of thermal expansion, during the cooling process T is temperature difference of liquid metal, and the casting will contract TL. If the contraction is at hot spot of length l, then the strain is as follows [37,41] ε T = α ΔT L/l (3.2) Where, =strain, = thermal expansion, T = temperature difference of liquid metal during cooling, L/ l = Change in length ratio due strain contraction in casting. Fine grains may contain many grain boundaries at a hot spot. Considering that numbers of grains are in length l and the diameter of grain is a. The number of grains at hot spot is l/a. Strain per boundary is [37]: ε b = α ΔT L a/l 2 (3.3) Thermal Stress: σ = ε T. E (3.4) Where E is the elastic modulus. 64

85 3.6 HOT TEARING CRITERIA Feurer (1976) proposed criterion to predict hot tear formation in castings. He investigated the feeding characteristics and hot tearing properties and found that the poor feeding is caused by solidification shrinkage because of the interlinking of dendrites. However, the dendrites act as porous filters. The residual liquid is fed through these porous dendrite filters. Insufficient feeding to the casting may result in the volumetric shrinkage. If the shrinkage velocity exceeds the maximum flow rate of feeding, hot tearing may occur. However, Feurer postulates that hot tear formation (HTF) is possible if [45]: SPV SRG (3.5) The maximum volumetric flow rate per volume is defined by: SPV 1 V V t ln V t (3.6) Where, SPV = the maximum volumetric flow rate through dendrite networks V = Volume, t = time 65

86 The volumetric solidification shrinkage is caused by the density difference between liquid and solid. It is defined by: ln V SRG t 1. t (3.7) Where, SRG= the volumetric shrinkage during solidification. Where, g 1 g ) (3.8) L L s ( L V= volume of solidify elements, = density of the liquid, L S = density of solid phase Clyne and Davies (1979 & 1981) defined a hot tearing criteria (HTC) as the ratio of the interdendrite separation between grains during a time period (liquid fraction from 0.01 to 0.10) of vulnerability, to the time period for stress relaxation as liquid feeding occur (liquid fraction from 0.1 to 0.06) as shown in Figure 36. The hot tearing susceptibility (HCS) is defined by [46,47]: Where, H. C. S. = t V t R = t 0.99 t0. 90 t 0.90 t 0.40 (3.9) t v = the vulnerable time period when cracks can propagate t = the time period for stress-relaxation during liquid feeding r t 0.99= the time at which the solid fraction is

87 t 0.9 = the time at which the solid fraction is 0.9 t 0.4 = the time at which the solid fraction is 0.4. Figure 36. Ratio of vulnerability ( t v ) to stress relaxation ( t r ) for hot tearing [46,47] Katgerman (1982) combines the hot tearing criteria of Feurer (1976) and Clyne and Davies (1979). His criterion is based on the concept that if feeding is inadequate the resultant volume reduction causes stresses to develop. The vulnerable time proportionality changes t to t tcr from 90 t Hot tearing formation (HTF) index defined as [48]: HTF t t t t 99 cr (3.10) cr 40 Where, t cr =distance from where after feeding is inadequate 67

88 The hot tearing formation index improved using coherency temperature, which defines where dendrite networks begin to form. Considering the coherency temperature ( t Coh ), then hot tearing formation is defined by HTF t t cr 99 (3.11) cr t t Coh Hatami et al. proposed a new criteria based on theories of Feurer (1976), Clyne and Davies (1979), and Katgerman (1982). The new theory considers a volume element in the mushy zone, called volume solid fraction of fs for hot cracking simulation. If there is no flow considered, then the zero flow point in mushy zone is known as solid fraction for rigidity. New criteria postulates that if a neighboring element has a solid of fraction less than the rigidity point, local liquid feeding is possible and no cracks will occur. If neighboring elements exceed the solid fraction of rigidity, then liquid feeding is not possible and hot cracks may occur. Such hot tearing formation (HTF) is determined on each element [49]: HTF T t T t cr cr (3.12) Where, t0.99 t cr = critical time at which element transmits stresses 68

89 Campbell suggested a significantly different hot tearing predictive model, suitable for qualitative assessment of different alloys as defined below [37,45]. CSC = α TLa l 2 t V t R (3.13) Where, α= Coefficient of thermal expansion, T= undercooling temperature, L= Casting length, a= grain size, l= Length of hot spot, α TLa =Accumulated thermal strain at hot spot l 2 t V t R = Ratio of vulnerability (t V ) to stress relief (t R ) The Niyama criterion was developed by E. Niyama in 1982 to predict shrinkage porosity. The Niyama is defined as [73] N y = G T (3.14) Where G is the thermal gradient and T is the cooling rate. Niyama criterion is based on strain rate. In the RDG hot tearing criterion [75], the depression pressure, p, over the mushy zone: p = psh + pmec ρgh (3.15) 69

90 Where, psh and pmec are the pressure drop contributions in the mushy zone associated with the solidification shrinkage and the deformation caused by fluid flow, respectively, ρ is density, g is gravitational constant, and h is the distance below the liquid melt level. The Hot Cracking Susceptibility (HCS) = 1 ε max (3.16) 70

91 3.7 HEAT TRANSFER IN CASTING SOLIDIFICATION The rate of solidification of liquid metal is controlled by the excessive heat in the liquid metal at the time of pouring and the rate of heat dissipation from the casting [37] FLOW OF HEAT INTERFACES WITH CHILL The heat flow from liquid metal interfaces with chill and it can be approximated as one dimensional heat transfer problem. Therefore, during the unsteady state (transient) conduction heat transfer in one dimensional, the flow of heat from liquid metal poured at melting temperature Tp against the mold wall at temperature T m, then the partial differential can be written as follows [37] 2 T( x, t) 1 T ( x, t ) 2 x t (3.17) The boundary conditions are as follows: At x, T T 0 And at x L, T Tm c k (3.18) c Where, is the thermal diffusivity of solid material, x is the Cartesian coordinates, k is the thermal conductivity, and c is the specific heat capacity. 71

92 3.7.2 FLOW OF HEAT INTO CASTING When heat is flowing into the casting, the latent heat of solidification is added to equation (14), then new equation is described by [37]: T x t T x t k 2 (, ) (, H c ) 2 (3.19) x t Where, H is the latent heat of solidification HEAT RELEASED DURING SOLIDIFICATION The rate of heat is released from mold during solidification is [37]: T ( x, t) Q HA (3.20) t The flow of heat is transferred to the mold. Considering the heat transfer coefficient is h and the unit per area of the mold is A, then the rate of heat transferred Q can be defined by Q ha( T ) c Tm (3.21) Where the mold has constant thickness and also the temperature difference T T ) is contact across the mold. ( c m 72

93 CHAPTER 4 - EXPERIMENTAL STUDIES OF DIE CASTING ALLOYS This section summarizes the development of a gravity die casting book mold assembly for hot tearing evaluation methods to conduct experimental studies in predicting hot tearing characteristics of A206.2, A380, Test-A-alloy, and AT72 alloys. The book mold is designed to be used as a tool for casting industries so that new alloys can be tested for hot tearing before production release. Both aluminum and magnesium die casting alloys are selected to examine hot tearing defects in casting. The experimental studies were conducted on several samples of casting at various temperatures so the casting could be analyzed for defects. Pouring and mold temperatures influence hot tearing propensity of die casting alloys. The contraction forces and cooling temperatures are measured in real time using a load cell (maximum rating: 500lb) and K-Type thermo-couples, respectively. The cast parts are analyzed for hot cracking and casting defects. 73

94 4.1 DESIGNING OF A BOOK MOLD ASSEMBLY The book mold assembly is designed as a gravity die casting mold based on concepts from constrained rod casting. Argo et al. and Cao et al. used constrained rod casting to investigate hot tearing phenomena in die casting alloys. The new book mold assembly is called Enhanced Constrained Rod Casting (ECRC). The ECRC contains five rods. The longest constrained rod is used for the measurement of thermal contraction forces during solidification. The remaining four constrained rods (A,B,C,D) with diameter of 9.5 mm and lengths of 51, 89, 127, and 165 mm are used for hot tearing evaluation. Each constrained rod has a bulbous ball end (restrain ball) with diameter of 19 mm as shown in Figure 37. The new book mold assembly for ECRC has a tapered sprue to facilitate liquid metal feeding through curved bend to constrained rods [2,14]. A schematic illustration of enhanced constrained rod casting is presented in Figure

95 Figure 37. A schematic illustration of enhanced constrained rod casting DESIGNING OF A SPRUE AND RUNNER FOR METAL FLOW The function of a sprue system is to facilitate the liquid metal into mold cavity without generating turbulence. A tapered sprue is a very important feature in a good gating system. It is developed following the Law of Continuity (conversation of mass) for fluid flow. This principle is derived from the fact that mass is always conserved in fluid systems regardless of the duct/pipeline complexity or direction of fluid flow. If fluid flow exits in a channel and the principles of mass flow are applied to the system, there exists a continuity of flow. This is defined as: The mean velocities at all across sections having equal areas are equal, and if areas are not equal the velocities are inversely proportional to respective cross sections [53]. 75

96 Consider a section of tapered sprue (circular) as specified in a control volume (CV) where the inlet area A 1 is larger than the outlet area A 2, thus per continuity of flow, the inlet velocity V 1 will be inversely proportional to outlet velocity V 2 as shown in Figure 38. Figure 38. Fluid flow in a tapered sprue (circular) Flow rate(q) = Area(A) velocity(v) (1) Volume flowing in (Q 1) = Volume flowing out(q 2) A 1 V 1 = A 2 V 2 (2) From the continuity equation (2), the velocity of liquid metal flow increases at outlet. A sharp bend between the sprue and the runner may develop eddies and lower pressure in localized regions, resulting in the reduction of fluid flow velocity. 76

97 The new design of tapered sprure and curved pipe would increase the fluid velocity. The inner and outer bend radius between tapered sprue and runner are designed by changing the bend geometries half of runner thickness (0.5 * t) and one and half times of runner thickness (1.5 * t) respectively. These changes will reduce stress concentration and form streamlined flow as shown in Figure 39 (a) and (b) respectively [54]. R = t 2 (3) R = 1.5 t (4) Where, R = inner radius of runner, R = outer radius of runner, and t = thickness of runner Figure 39. a) Bend radius for inner b) Bend radius for outer radius between sprue and runner [54] 77

98 The new feeding concepts are designed to feed mold cavities of enhanced constrained rods from the bottom through a tapered sprue and curved runner, which permits liquid metal to flow as streamlined with minimum hindrance. Additionally, it eliminates eddies, aspiration, and turbulent flow in the molten casting alloy as shown in Figure 40 and Figure 41 [55, 56]. Figure 40. Eddies and aspiration [56] Figure 41. Formation of streamlined fluid flow [56] 78

99 4.1.2 DEVELOPMENT OF ENHANCED CONTRAINED ROD CASTING The book mold assembly of Enhanced Constrained Rod Casting was developed using P20 tool steel. Haas VF3SS and Kitamura H400 CNC machines were used for machining of mold cavities. The surface finish of left and right mold cavities were measured in the ranges of micro inches. The following tools were used to perform milling, grinding, and surface finishing operations of book mold cavities as shown in Table 8 [57]. 79

100 Table 8. Types of Tools Applied in Development of Book Mold Assembly [57]. A 3D model was developed using SolidWorks software for both left and right side of the book mold assembly. From a 3D solid model, 2D engineering drawings were created as all the detailed dimensions are shown in Figure 42 and Figure

101 Figure 42. A 2D engineering drawing of left side mold assembly 81

102 Figure 43. A 2D engineering drawing of right side mold assembly 82

103 The book mold assembly consists of the left and right side of the mold which are connected with two separable hinges on the sprue side. The other end is free to open and close during casting operations. The book mold assembly contains four constrained rods (A, B, C, D). Each constrained rod has a bulbous or restrain ball at the end. An additional rod is added without bulbous end for thermal contraction force measurements. The views of the left and right molds are shown in Figure 44, Figure 45, Figure 46, and Figure 47 respectively. The dimensions for book mold assembly of enhanced constrained rod casting are given below [14]. Dimension: Length: 12 inch, Width: 9 inch, Thickness: 1.25 inch (half mold) The length of constrained rods: A: 6 inch (152.4 mm), B: 4.5 inch (114.3), C: 3 inch (76.2 mm), and D: 2 inch (50.8 mm). The diameter of rod: 0.38 inch (9.65 mm) The diameter of restrain ball: 0.76 inch (19.30 mm) The length of measurement rod for load cell: 7.5 inch (190.5 mm) 83

104 Figure 44. The details of installation for left mold cavity Figure 45. A 3D view of left mold cavity 84

105 Figure 46. The details of installation for right mold cavity Figure 47. A 3D view of right mold cavity 85

106 4.2 EXPERIMENTAL STUDIES OF A206.2, TEST-A, A380, AND AT72 ALLOYS Experimental studies are conducted for hot tearing evaluation of A206.2, Test-A, A380, and AT72 alloys. A schematic illustration of the experimental set up is shown in Figure 48. The apparatuses for experimental studies include a book mold assembly of enhanced constrained rod casting (ECRC), data acquisition system, donut load cell, load cell washers, VAC Input/15 VDC Output Power Supply for load cell, quartz rod with grooves, shaft collar, flat washers, ceramic washers, K-type thermocouples with ceramic beads, TC Thermocouple Data Logger, and support bracket as listed in Table 9. The photographs for apparatuses and test equipment are as presented in Figures 49, 50, 51, 52, 53, and 54, respectively. 86

107 Figure 48. A schematic illustration of experimental set up for ECRC mold 87

108 Table 9. List of Apparatuses for Experimental Studies Number Description of Apparatuses Quantity References 1 2 A book Mold Assembly for Enhanced Constrained Rod casting using P20 tool Steel LTH300, lb, Donut Load Cell, Standard, 1/4" [6.35 mm] ID - OT: -60 F to 200 F 1 N/A LTH300 (L2760), 1/4" ID, Top Washer, A2 Tool Steel Material LTH300 (L2760), 0.40" ID, Bottom Washer, A2 Tool Steel IAC180, Power Supply Kit for CSG110, VAC Input/15 VDC Output Power Supply CSG110, Strain Gauge Amplifier, Standard, Enclosed With Din Rail Mount, With DB9 Connectors, ABS-94HB Black Enclosure, Analog Output, +/-5 VDC, +/-10 VDC K-Type thermocouples 8 Ceramic Beads, Dole holes (DH ) 1 for each test one pack (QTY-45) 9 Clear Fused Quartz Rod, 6mm diameter 4 feet USB data-acquisition-1208fs [Tracer DAQ] EL-USB-TC Thermocouple Data Logger Ceramic washers 2 N/A 13 Steel Washers 2 N/A 14 Support Bracket 1 N/A 15 Shaft Collar 1 N/A 16 Hex head Bolts (3/8"X6") long 2 N/A 17 Double pole Single Through Control Switch 1 N/A 18 Heat Sink tube 1/2 feet N/A 19 Alignment Pin for Book Mold Assembly 1 N/A

109 Figure 49. Book mold assembly of enhanced constrained rod casting Figure 50. Closed book mold assembly of enhanced constrained rod casting with K-type thermocouple 89

110 Figure 51. Data acquisition system (Tracer DAQ) with donut load cell [59,61] Figure 52. Quartz rod for force measurement with shaft collar [60] 90

111 Figure 53. USB-TC thermocouple data logger with K-type thermocouple [62,63] Figure 54. Load cell compression, flat, and ceramic washers [59,64] 91

112 4.3 INSTRUMENTATION OF EXPERIMENTAL APPARATUSES Figure 4.19 shows instrumentation and installation of experimental apparatuses for Enhanced Constrained Rod Casting (ECRC) mold. A quartz rod with grooves is installed in ECRC mold of the longest measurement rod to measure contraction forces during casting solidification. The quartz rod assembles with a donut load cell, compression washers, ceramic washers, and steel washers are shown in Figure 55. The quartz rod is supported by a support bracket, which is attached to the exterior of the ECRC mold as shown in Figure 57. The operating temperature of load cell is -60 F to 200 F. Therefore, to protect the load cell from high temperatures, the ceramic and steel washers are used to form a heat sink, so heat would be dissipated properly to keep the load cell under operating temperature, as shown in Figure 56. The donut load cell is connected to the USB dataacquisition-1208fs [Tracer DAQ] and an output cable is routed from the USB dataacquisition-1208fs and connected to the PC for data collection. Two compression load washers (Hardened washer - A2 Tool Steel) are used in front and back of the load cell mounted over the quartz rod. When compression loads are applied on the load cell due to contraction/shrinkage of casting during cooling, the load cell transmits the contraction forces to an amplifier. The output thermal contraction forces are amplified and then displayed on the data acquisition system in volts. A C-clamp is used to tighten and hold the book mold assembly closed during casting. 92

113 To measure the temperatures of casing during solidification, the EL-USB-TC Thermocouple Data Logger is used. It is connected with a K-thermocouple and installed through a hole at the left side of the book mold, adjacent to the longest measurement rod for measurements in real time. The data cables of the load cell and K-type thermocouples were connected to a control switch (double pole single throw) so these two apparatuses can be turned On/Off simultaneously. Contraction forces and temperatures data were measured and recorded on the PC. This data was then converted into an Excel or spreadsheet format (from *.csv and *.sch) for post-processing and evaluation of casting defects. The conversion for thermal contraction forces are: one volt per fifty pounds (1Volt = 50 lb), which was then converted from pounds to Newton (N) units. Figure 55. Instrumentation and installation of testing equipment for ECRC mold 93

114 Figure 56. Installation of load cell along with quartz rod and washers Figure 57. Support bracket for quartz rod and load cell assembly 94

115 4.4 EXPERIMENTAL MATRIX AND CHARATERTICS OF A206.2, TEST-A, A380, AT72 ALLOYS For experimental studies, A206.2, Test-A (Aluminum alloy), A380, and AT72 (Magnesium alloy) alloys were selected for hot tearing evaluation. Aluminum alloy A206.2 possesses several characteristics that are suited for automotive, aerospace, and military applications. A206.2 is a heat treatable die casting alloy, employed to have high tensile, yield strength, and high fracture toughness. A206.2 is subjected to corrosion problems due to high wt% of copper content. In spite of its excellent properties, A206.2 alloy is not often used, primarily because of its propensity for hot tearing. A 35 kilogram ingot of A206.2 alloys was sourced from Trialco Inc. for experimental studies. The actual data for chemical compositions were provided by Trialco Inc. for A206.2, and is presented in Table 10 [65]. Table 10. Chemical Compositions of A206.2 Alloy A206.2 Alloy (wt%) Si Cu Mn Mg Fe Ti Zn Al Standard Actual

116 Test-A alloy is a new alloy, which was developed by an automotive component supplier to have excellent corrosion resistance properties. In this study, Test-A alloy contains a high percentage of Silicon (9.4 wt% of Si), which is one of the most important alloying elements used in die casting alloys. Higher percentages of Silicon content are responsible for decreases in thermal expansion coefficient of the casting. The Si content increases the fluidity, allows for better mold filling, reduces shrinkage, and does not diminish corrosion resistance of aluminum. The actual data for Test-A alloy chemical composition is presented in Table 11. This data was used for experimental studies. Table 11. Chemical Compositions of Test-A Alloy A380 die casting alloy is widely used in the manufacture of aerospace and automotive components which includes engine parts, gear boxes, electronic enclosures, support brackets, and power equipment. A380 has excellent fluidity due to high Si content, and resistance to hot tearing. Due to its high thermal conductivity, it dissipates heat quickly during the solidification process. Table 12 contains the chemical composition of A380 that were used for experimental studies. 96

117 Table 12. Chemical Compositions of A380 Alloy Table 13 contains the Chemical Compositions of AT72 (Mg-7Al-2Sn) that were used for experimental studies. AT72 is a new magnesium alloy, which has been developed for thin wall and light weight casting parts. The addition of 2% Sn content enables the alloy to be heat treatable. AT72 alloy is close to AZ91D alloy except 0.5% of Zn content. Table 13. Chemical Compositions of AT72 (Mg-7Al-2Sn) 97

118 4.5 MELTING AND CASTING PROCESSES OF A206.2, TEST-A, A380, AND AT72 ALLOYS The melting of aluminum A206.2, Test-A, and A380 alloys were conducted in a silicon carbide (SiC) crucible using an induction furnace. The alloys were heated above C liquidus for each pouring temperature. It took minutes to complete the melting process. The mold cavities of Enhanced Constrained Rod Casting (ECRC) were properly cleaned and two coats of Boron Nitride (BN) were applied 24 hours prior to pouring. The ECRC mold was heated up to 150 C for an hour in an electric oven and then removed for cooling. Then the instrumentation was assembled to mold for experimental studies. The initial temperature of ECRC mold was kept at room temperature (25.5 C) before pouring of A206.2, A380, and Test-A alloys for experimental studies. Three castings were carried out for each alloy at pouring temperatures of 700 C, 750 C, and 800 C. The selection of these samples and pouring temperatures were based on previous studies and the evaluation of die casting alloys at elevated temperatures. Three pouring temperatures were not chosen using liquidus plus super heat temperatures because it may vary for alloys. But three pouring temperatures of 700 C, 750 C, and 800 C used for consistency. 98

119 In the first test, the liquid metal of A260.2 alloy was poured at pouring temperature of 700 C into the tapered sprue through the runner, which feeds the measurement and constrained rods. It took approximately 3 seconds to fill the entire the mold cavities of Enhanced Constrained Rod Casting. After each cast, a time interval of minutes were used for cooling the book mold assembly. These casting processes were repeated for the A206.2 alloy at pouring temperatures of 760 C and 800 C, respectively. Similar casting processes were followed on the A380 and Test-A die casting alloys and the liquid metal was poured at pouring temperatures of 700 C, 750 C, and 800 C. The melting of die casting AT72 (magnesium) alloy was conducted in a low carbon steel pouring cup, in an Electric Resistance Furnace. The melt was protected with a tube inserted into the furnace with sulfur hexafluoride and carbon dioxide gas layer. During the pouring of AT72 alloy, the layer of sulfur hexafluoride and carbon dioxide gas were used to prevent exposures of the Mg to oxygen, which can cause fires. The ECRC book mold assembly was heated up to 156 C, 147 C, and 165 C for pouring temperatures of 675 C, 710 C, and 750 C, respectively. Before pouring, the ECRC mold was set up for measuring the contraction forces and temperatures. These processes were repeated for each pouring of 710 C and 750 C. After each run, a time interval of 50 minutes was used for cooling and removing the cast from the mold. It took an hour to heat up the mold in the furnace for the next run. After three experimental runs of AT72, the castings were 99

120 evaluated for hot tearing defects, and measured data were analyzed. All the test data can be found in appendices for A206.2, A380, Test-A, and At72 alloys FORCE MEASUREMENTS AND DATA ANALYSIS OF A206.2 ALLOY The data collection devices were turned on before pouring and contraction forces and temperatures were measured and displayed on a PC in real-time using the data acquisition system. The measured data was recorded for post processing to evaluate hot tearing formation and solidification stages. The switch for data acquisition was turned off after 6-8 minutes of each test. Finally, castings were removed from the book mold assembly after minutes for observation and analysis. Recorded data was analyzed and graphs were plotted for each casting. Figure 58 shows the casting of ECRC as it solidified inside the mold. Figure 58. Casting of ECRC as solidified inside mold after pouring of Test-A-alloy at 700 C 100

121 Testing of the Test-A-alloy was performed at pouring temperatures of 700 C, 750 C, and 800 C. The collected data from experimental studies were used for post processing and understanding of the formation of hot tears. As the casting solidified after pouring each alloy, the cooling curve and contraction forces were measured as a function of time HOT TEARING EVALUATION OF A206.2 AT 700 C The first sample of A206.2 alloy was poured at pouring temperature of 700 C. The contraction forces and temperatures were recorded as a function of time as shown in Figure 59. The maximum casting temperature was recorded at 643 C. At the initial stage of solidification, the thermal contraction force rose for a short period due to solidification shrinkage. The contraction force was measured to be 20.9 N, corresponding to a temperature of 559 C. At this point the load dropped because the long rod developed hot tear and fractured. After 5 seconds, the contraction force rose for 21 seconds due to linear contraction of short fractured end of the long rod. The force was measured to be N, corresponding to a temperature of 499 C. After 40 seconds, the contraction force curve reached a steady state, where the temperature mold and casting achieved the equilibrium as shown in Figure

122 The cast part was removed from the mold and inspected for hot tearing failures. A photograph of ECRC showed cracks due to hot tearing at restrain ball of constrained rod A, which is partially connected as shown in Figure 60. Restrain ball is fully separated from constrained rod B, whereas constrained rod C is separated from feeder. There are no cracks observed due to hot tearing at constrained rod D. A summary of observation analysis and results for hot tearing cracks are presented in Table 14. Figure 59. Contraction force and cooling curve of A206.2 at 700 C 102

123 . Figure 60. A photograph of cast part exhibited hot tearing for A206.2 C at 700 C Table 14. Summary of Hot Tearing Results Constrained Rod A Constrained Rod B Constrained Rod C Constrained Rod D Measurement Rod for Load cell Feeder to constrained rods Hot tears at restrain ball Hot tears at restrain ball and fully seperated Hot tears at constrained Rod C No hot tears Hot tears away from middle toward load cell Big cracks at feeder 103

124 HOT TEARING EVALUATION OF A206.2 AT 760 C The second sample of A206.2 alloy was poured at 760 C. The graphs of contraction force and temperature as a function of time are shown in Figure 61. The maximum casting temperature was recorded to be C. At the initial stage of solidification, the casting started solidifying where the contraction force rose to point where a noticeable change occurred due to the development of hot tearing in the longest rod. At this stage, the contraction force was measure to be 78.6N, which corresponds to a temperature of 651 C. The contraction force rose smoothly for 15 seconds due to friction between broken rod and mold. The contraction force reached to a maximum point and it was measured to be 86 N, which corresponds to a temperature of C. The contraction force curve declined and reached to a thermal equilibrium where mold and casting have the same temperature as shown in Figure 62. The photograph of the ECRC process showed cracks due to hot tearing at the restrain ball, which was separated from the constrained rod A as shown in Figure 62. Constrained rod B and C were fully separated from the feeder. There were no cracks observed due to hot tearing at constrained rod D. A summary of the observation analysis and results for hot tearing cracks are presented in Table

125 Figure 61. Contraction force and cooling curve of A206.2 at 760 C 105

126 Figure 62. A Photograph of cast part exhibited hot tearing for A206.2 C at 760 C Table 15. Summary of Hot Tearing Results Constrained Rod A Constrained Rod B Constrained Rod C Constrained Rod D Measurement Rod for Load cell Feeder to constrained rods Hot tears at restrain ball and fully seperated Hot tears at the junction of feeder and Constrained Rod Hot tears at the junction of feeder and Constrained No hot tears Hot tears at the junction of feeder and Meassurement Rod Big cracks at feeder 106

127 HOT TEARING EVALUATION OF A206.2 AT 800 C The third sample of A206.2 alloy was poured at 800 C. The graph of the Contraction Force and Temperature vs time is shown in Figure 63. The maximum casting temperature was recorded to be 750 C. At the initial stage of liquidus temperature, the contract force rose smoothly and measured to be N, corresponding to a temperature of 627 C. The contract force dropped quickly due to the development of a hot tearing in the long rod. After 5 second, cleavage was complete. The contraction force rose smoothly for 15 seconds due to solid state contraction of rod. The contraction forces reached to a maximum of 76.6 N, which corresponding to a temperature of 505 C. At this stage, the contraction force curve declined to a point where the casting reached a thermal equilibrium state with the mold. A photograph of ECRC showed cracks due to hot tearing at restrain ball, which was separated from constrained rod A as shown in Figure 64. Hot tearing at restrain ball of constrained rod B occurred but it remained connected. There were no cracks observed due to hot tearing at constrained rod C and D. A summary of observation analysis and results for hot tearing cracks are presented in Table

128 Figure 63. Contraction force and cooling curve of A206.2 at 800 C 108

129 Figure 64. A photograph of the ECRC showed hot tearing cracks at 800 C Table 16. Summary of Hot Tearing Results at 800 C Constrained Rod A Hot tears at restrain ball and fully seperated Constrained Rod B Hot tears at restrain ball Constrained Rod C No hot tears Constrained Rod D No hot tears Measurement Rod for Load cell Hot tears at the junction of feeder and Meassurement Rod Feeder to constrained rods Big cracks at feeder 109

130 RESULTS AND DISCUSSION In the solidification process as heat is removed from Enhanced Constrained Rod Casting through conduction, thinner section of the casting which include constrained rods (A,B,C,D) solidified faster than feeder and sprue. Because, rods are smaller in diameter than the feeder, the rods probably solidified slightly before the feeder. The first sample of A206.2 casting alloy poured at 700 C, exhibited significant hot tearing. As the casting temperature increased from 700 C to 760 C on second sample of A206.2 casting, a greater amount of hot tearing in the casting was observed. The third sample of A206.2 was poured at 800 C, and the hot tearing in the casting decreased slightly. Low volumetric shrinkage was observed due to heat maintained in the casting. Cu addition increases shrinkage of alloy (4.7% Cu) due to poor interdendritic fluidity and having a wide solidification interval where less amount of residual eutectic at grain boundaries. A206.2 alloy exhibited higher hot tearing at 760 C and 800 C than 700 C of pouring temperatures. Thus it is obvious that the volumetric shrinkage is influenced by both content of %Cu in alloy and pouring temperatures. 110

131 The A206.2 alloy generated a basic shape in the time vs load plots. However, each shape extended at each pouring temperature FORCE MEASUREMENTS AND DATA ANALYSIS OF TEST-A The studies of Test-A alloy are performed at three different pouring temperatures (700 C, 750 C, and 800 C) using the enhanced constrained rod casting (ECRC) mold. The data collection apparatuses were turned on before pouring and contraction forces and temperatures were collected and displayed on the PC in real-time. The measured data was recorded for post processing to determine hot tearing formation and solidification curves. The control switch for data acquisitions were turned off after 6-8 minutes of casting. After each cast, a time interval of minutes allowed for cooling the cast and book mold assembly. The castings were removed from book mold assembly for evaluation. Recorded data were analyzed and graphs were plotted for each casting to determine the formation of hot tearing defects. 111

132 HOT TEARING EVALUATION OF TEST-A AT 700 C The first sample of Test-A alloy was poured at 700 C. The contraction forces and temperatures were as a function of time, as shown in Figure 65. The maximum casting temperature was recorded to be 650 C. At the initial stage of liquidus, the contraction force curve lifted slightly however there was no hot tearing occurred in the casting. The contraction force curve rose smoothly and reached to a maximum of N, which corresponded to a temperature of 552 C. The contraction force curve declined and reached a thermal equilibrium state with the mold. A photograph of ECRC casting shows no cracks in constrained rod A, B, C, D, and measurement rod for load cell as shown in Figure 66. Figure 67 shows casting solidified inside mold at 700 C pouring of Test-A-alloy. 112

133 Figure 65. Contraction force and cooling curve of Test-A at 700 C 113

134 Figure 66. A photograph of cast part did not exhibit hot tearing for Test-A at 700 C Figure 67. Casting of ECRC as solidified inside mold at 700 C of Test-A-alloy 114

135 HOT TEARING EVALUATION OF TEST-A AT 750 C The second sample of Test-A alloy was poured at 750 C. The contraction forces and temperatures were as a function of time, as shown in Figure 68. After pouring, the maximum casting temperature was recorded at 692 C. At the initial solidification of casting, the contraction force rose sharply. There was no crack occurred in the longest rod. Formation of small triangle is caused by high temperature thermals strain. The contraction force was measured to be N, which corresponds to a temperature of C. At this point, the casting is fully solidified. Then the contraction force rose vertically to a maximum of N, corresponding to a temperature of C. At the maximum point, the contraction force curve declined slightly and reached a thermal equilibrium with the mold as shown in Figure 68.. A photograph of ECRC casting shows cracks due to hot tearing at restrain ball, which was separated from constrained rod A as shown in Figure 69. Hot tearing occurred at the junction of the constrained rod D and the feeder. There were no cracks observed due to hot tearing at constrained rod B and C. A summary of observation analysis and results for hot tearing cracks are presented in Table 17. Figure 70 shows the solidified casting inside the mold and hot tearing that occurred at Constrained Rod A and Rod D during solidification. 115

136 Figure 68. Contraction force and cooling curve of Test-A alloy at 750 C 116

137 Figure 69. A photograph of cast part exhibited hot tearing for Test-A at 750 C Figure 70. Solidified casting inside mold showed hot tearing cracks at 750 C 117

138 Table 17. Summary of Hot Tearing Results at 750 C Constrained Rod A Constrained Rod B Constrained Rod C Constrained Rod D Measurement Rod for Load cell Feeder to constrained rods Hot tears at restrain ball and fully seperated No hot tears No hot tears Hot tears at the junction of feeder and Rod No hot tears No hot tears HOT TEARING EVALUATION OF TEST-A AT 800 C The third sample of Test-A alloy was poured at 800 C. The contraction forces and temperatures as a function of time were recorded as shown in Figure 71. After pouring, the maximum casting temperature was recorded at 729 C. At the initial solidification of the casting, the contraction force rose and dropped quickly forming a semi-circle due to high temperature thermal strain. No hot tearing exhibited in the longest rod. The contraction force rose smoothly to a maximum of N, corresponding to a temperature of 563 C. The contraction force curve declined slightly and reached a thermal equilibrium. 118

139 A photograph of the ECRC part shows cracks due to hot tearing at the restrain ball, which was separated from constrained rod A as shown in Figure 72. There were no cracks observed due to hot tearing at constrained rod B, C, and D. A summary of observation analysis and results for hot tearing cracks are presented in Table 18. Figure 71. Contraction forces and cooling curve for Test-A at 800 C 119

140 Figure 72. A photograph of cast part exhibited hot tearing for test-a at 800 C Table 18. Summary of Hot Tearing Results at 800 C Constrained Rod A Constrained Rod B Constrained Rod C Constrained Rod D Measurement Rod for Load cell Feeder to constrained rods Hot tears at restrain ball and fully seperated No hot tears No hot tears No hot tears No hot tears No hot tears 120

141 RESULTS AND DISCUSSION There was no hot tearing occurred at 700 C of pouring temperature. The longest did not show hot tearing at the pouring temperatures of 750 C and 800 C of pouring temperatures. Therefore, Test-A alloy is a good alloy and did not show hot tearing in the longest rod at different pouring temperatures. Because, Si addition increases volumetric contraction in the formation of a silicon phase, which expands and feed dendrites through surface tension during solidification, hence reduces shrinkage porosity. Also, it contains high heat of fusion which maintain the heat and fluidity FORCE MEASUREMENTS AND DATA ANALYSIS OF A380 ALLOY Similarly, casting experiments of A380 alloy were conducted. Testing of A380 alloy was performed at three different pouring temperatures (700 C, 750 C, and 800 C) as liquid metal was poured into the mold cavities of enhanced constrained rod casting (ECRC). The recorded data were analyzed and graphs were plotted for each casting to characterize hot tearing defects. 121

142 HOT TEARING EVALUATION OF A380 ALLOY AT 700 C The first sample of A380 alloy was poured at 700 C. The graph of contraction forces and temperatures as a function of time (Cooling curve) is shown in Figure 73. After pouring, the maximum cooling temperature was recorded at 657 C. At the initial solidification, the contraction force rose vertically and declined sharply due to high temperature thermal strain in the casting. At this stage, the contraction force measured to be N, which corresponds to a temperature of C. Thermal contraction force reached to a maximum of N, corresponded to a cooling temperature of C. No hot tearing occurred in the longest rod. The contraction force curve declined constantly and reached a thermal equilibrium with the mold. 122

143 Figure 73. Contraction forces and temperatures as a function of time for A380 alloy at 700 C. A photograph of first sample casting shows hot tearing cracks at Constrained Rod A and B as shown in Figure 74. There were no cracks observed due to hot tearing at Constrained Rod C and D. 123

144 Figure 74. A photograph of cast part exhibited hot tearing cracks at 700 C HOT TEARING EVALUATION OF A380 ALLOY AT 750 C The second sample of A380 alloy was poured at 750 C. The graph of contraction forces and temperatures as a function of time (Cooling curve) is shown in Figure 75. After pouring, the maximum cooling temperature of casting was recorded at C. At the initial solidification of casting, the contraction force rose smoothly and reached to a maximum of N, which corresponds to a cooling temperature of 656 C. The contraction force curve declined constantly and reached a thermal equilibrium state. Due to low pouring temperature this did not show similar graphs pattern as other test. Additionally, the smallest rod D did not fill. 124

145 Figure 75. Contraction forces and cooling curve for A380 alloy at 750 C A photograph of second sample casting is shown in Figure 76. There were no hot tearing cracks observed of A380 alloy when poured at 750 C. However, the restrain ball cavity of constrained rod D was not filled. 125

146 Figure 76. A photograph of cast part did not exhibit hot tearing at 750 C HOT TEARING EVALUATION OF A380 ALLOY AT 800 C The third sample of A380 alloy was poured at 800 C. The graph of contraction forces and temperatures as a function of time (Cooling curve) is shown in Figure 77. After pouring, the maximum cooling temperature of casting was recorded at 703 C. At the initial solidification of casting, the contraction force rose linearly and reached to a point where hot tearing developed. At this point, the contraction force measured to be N, which corresponds to a cooling temperature of 554 C. The longest rod did not show hot tearing. The contraction force dropped to a value of N and rose again 126

147 linearly to a maximum of N, which corresponds to a cooling temperature of 482 C. This happened due to heat of fusion and solidification of the casting. The contraction force curve declined constantly and reached a thermal equilibrium. Figure 77. Contraction forces and cooling curve for A380 alloy at 800 C A photograph of third sample casting is shown in Figure 78. The third sample of casting which exhibited hot tearing at the junction of sprue and constrained rod B. All other constrained rods did not show hot tearing defects. 127

148 Figure 78. A photograph of cast part exhibited hot tearing cracks at 800 C RESULTS AND DISCUSSION The first sample of A380 alloy did not exhibit hot tearing at 700 C of pouring temperature in the longest rod. The second sample of A380 do not correlate with other testing data due to low pouring temperature as it showed contraction force curve shifted to thermal equilibrium state. The third sample of A380 at the pouring temperature of 800 Cdid not exhibit hot tearing in the longest rod. Pouring temperature and alloy compositions (Cu addition of 3.5%) influence hot tearing despite having 9.1%Si content. 128

149 4.5.4 FORCE MEASUREMENT AND DATA ANALYSIS OF AT72 ALLOY The casting experiment was conducted on AT72 (magnesium alloy) alloy to determine hot tearing. The casting was poured at three pouring temperatures (675 C, 710 C, and 750 C). For each run, the liquid metal was heated above C of pouring temperature. The mold was heated to three temperatures of 156 C, 147 C, and 165 C in an electric furnace (Kiln power rated watts). Three sample of cast parts were examined for casting defects. Measured data were analyzed and graphs were plotted for each casting to understand hot tearing defects HOT TEARING EVALUATION OF AT72 ALLOY AT 675 C The first sample of AT72 alloy was poured at 675 C pouring temperature into ECRC mold. At the time of pouring, the mold temperature was at 156 C. The graph of contraction forces and temperatures are shown as a function of time (Cooling curve) in Figure 79. After pouring, the maximum cooling temperature of casting was recorded at 653 C. At the initial solidification of casting, the contraction force rose non-linearly and reached to a point where a noticeable change occurred due to high thermal strain. At this point, the contraction force was measured to be 53.4 N, which corresponds to a cooling temperature 129

150 of 545 C. At this stage, it was measured to be 61.5 N, which corresponds to a cooling temperature of 415 C. The contraction force curve declined and reached a thermal equilibrium state. A photograph of first sample of casting is shown in Figure 80. The first sample of cast part exhibited hot tearing at the retrain Ball of constrained rod A. All other constrained rods did not exhibit hot tearing defects. Figure 79. Contraction forces and cooling curve for AT72 alloy at 675 C 130

151 Figure 80. A photograph of cast part exhibited hot tearing cracks at 675 C HOT TEARING EVALUATION OF AT72 ALLOY AT 710 C The second sample of AT72 alloy was poured at 675 C pouring temperature in the ECRC mold. At the time of pouring, the mold temperature was at 147 C. The graph of contraction forces and temperatures is shown as a function of time (cooling curve) in Figure 81. After pouring, the maximum cooling temperature of casting was recorded at 684 C. At the initial solidification of casting, the contraction force rose linearly and dropped due to high thermal strain. At this point, the contraction force measured to be 43.8 N, which corresponds to a cooling temperature of 677 C. The contraction force rose again non- 131

152 linearly to a maximum of 79.6 N, corresponding to a cooling temperature of 572 C. From this stage, the contraction force curve declined and reached a thermal equilibrium state. Figure 81. Contraction forces and cooling curve for AT72 alloy at 710 C A photograph of second sample of casting is shown in Figure 82. Hot tearing exhibited between Constrained Rod A and feeder. There were no casting defects found on other constrained rods. 132

153 Figure 82. A photograph of cast part exhibited hot tearing at 710 C HOT TEARING EVALUATION OF AT72 ALLOY AT 750 C The third sample of AT72 alloy was poured at 750 C pouring temperature into ECRC mold. At the time of pouring, the mold temperature was at 165 C. The graph of contraction forces and temperatures are shown as a function of time (cooling curve) in Figure 83. After pouring, the maximum cooling temperature of casting was recorded at 712 C. At the initial solidification of casting, the contraction force rose linearly and shifted due to high thermal strain. At this transition point, the contraction force was measured to be 37.5 N, which corresponds to a cooling temperature of 677 C. The contraction force continued 133

154 to an maximum of N, which corresponds to a cooling temperature of 592 C. From this stage, the contraction force curve declined and reached a thermal equilibrium state. A photograph of second sample of casting is shown in Figure 84. Hot tearing exhibited between Constrained Rod A and feeder. There were no casting defects found on other constrained rods. Figure 83. Contraction forces and cooling curve for AT72 alloy at 750 C 134

155 Figure 84. A photograph of cast part exhibited hot tearing at 750 C RESULTS AND DISCUSSION The first sample of AT72 did not exhibit hot tearing at 675 C of pouring temperature. At pouring temperature 710 C, the second sample of AT72 did not show hot tearing. At pouring temperature 750 C, the third sample of AT72 did not show hot tearing in the longest rod. 135

156 Pouring temperatures and mold temperatures influence hot tearing. By increasing the mold temperature, hot tearing will be reduced due to small thermal gradient which lower the strain contraction rate. 4.6 CONCLUSIONS In the solidification process as heat is removed from Enhanced Constrained Rod Casting through conduction, thinner section of the casting which include constrained rods (A,B,C,D) solidified faster than feeder and sprue. Because, rods are smaller in diameter than the feeder, the rods probably solidified slightly before the feeder. The first sample of A206.2 casting alloy poured at 700 C, exhibited significant hot tearing. As the casting temperature increased from 700 C to 760 C on second sample of A206.2 casting, a greater amount of hot tearing in the casting was observed. The third sample of A206.2 was poured at 800 C, and the hot tearing in the casting decreased slightly. Low volumetric shrinkage was observed due to heat maintained in the casting. Cu addition increases shrinkage of alloy (4.7% Cu) due to poor interdendritic fluidity and having a wide solidification interval where less amount of residual eutectic at grain boundaries. A206.2 alloy exhibited higher hot tearing at 760 C and 800 C than 700 C of pouring temperatures. Thus it is obvious that the volumetric shrinkage is influenced by both content of %Cu in alloy and pouring temperatures. 136

157 Test A-alloy did not show hot tearing occurred at 700 C of pouring temperature. At the pouring temperatures of 750 C and 800 C, Test-A alloy did not show hot tearing in the longest rod. Therefore, Test-A alloy is a good alloy. Because, Si addition increases volumetric contraction in the formation of a silicon phase, which expands and feed dendrites through surface tension during solidification, hence reduces shrinkage porosity. Also, it contains high heat of fusion which maintain the heat and fluidity. The first sample and third sample of A380 alloy did not exhibit hot tearing poured at 700 C and 800 C. However, hot tearing occurred at the junction of constrained rod A and rod B. The second sample of A380 alloy did not generate a basic shape in the time vs load plots due to low pouring temperature. Pouring temperature and alloy compositions (Cu addition of 3.5%) influence hot tearing despite having 9.1%Si content. The AT72 alloy did not show hot tearing in the longest rod at the pouring temperatures of 675 C, 710 C, and 750 C. Pouring temperatures and mold temperatures influence hot tearing. 137

158 CHAPTER 5 NUMERICAL AND HOT TEARING PREDICTIVE MODELING 5.1 NUMERICAL AND PREDICTIVE MOEDLING APPROACH This section discusses numerical and hot tearing predictive models to study the casting and simulation. The metal casting process is a complex process due to the cooling of liquid metal in a rigid mold, thermal properties of liquid metal and mold, and uneven cooling of liquid metal at elevated temperatures. The modeling and simulation of casting due to nonlinear behavior of materials is a challenge which require simplifications in modeling, setting up boundary conditions, and simulation in order to correlate with experimental studies. Due to the complexity of physical phenomena and limitations of finite element software, it is necessary to make assumptions and simplifications in a particular area of mold cavities for casting simulation. In the numerical modeling approach, the finite element model of enhanced constrained rod (ECRC) was developed to study the flow and shrinkage in the casting using MagmaSoft. A comparative analysis was performed between constraint rod casting (CRC) and enhanced constrained rod casting (ECRC). 138

159 The simplified version of 2-D ECRC model cavity was used to simulate temperature distribution and thermal strain during phase changes using ANSYS. Sequential Coupled Transient Method was used to study casting simulation of enhanced constrained rod casting (ECRC) at 700 C, 750 C, and 800 C pouring temperatures for A206.2, Test-A, A380, and AT72 alloys. The hot tearing predictive models were used to predict shrinkage porosity and hot tearing susceptibility based on thermal strain rates. 5.2 FINITE ELEMENT MODELING AND SIMULATION MagmaSoft was used to model the current constrained rod casting (CRC) and enhanced constrained rod casting (ECRC) to perform simulations as shown in Figure 85 and Figure 86 respectively. In modeling of CRC and ECRC, vent holes were incorporated so casting would not have porosity. 139

160 Figure 85. Model of current constrained rod casting (CRC) Figure 86. Model of enhanced constrained rod casting (ECRC) 140

161 The modeling of current constrained rod casting (CRC) and enhanced constrained rod casting (ECRC) are to perform a comparative analysis to determine casting velocity and solidification porosity. Tracers were added at the top of sprue to simulate and predict casting velocity. The Casting filling velocity is critical in filling the mold cavities since even millisecond improvements are advantageous due to rapid cooling and solidification of casting. To produce good casting, casting industries optimize the filling velocities to make sure liquid metal reaches to corners and cavities farthest in a short amount of time FLOW MODELING AND SIMULATION The casting filling velocity of constrained rod casting (CRC) of m/s in comparison to enhanced constrained rod casting (ECRC) is m/s as shown in Figure 87 and Figure 88. The new mold design of ECRC provides 42.4% higher filling velocity over current CRC. The ECRC model has been optimized with tapered sprue, bend, and runner, which develops streamline flow without forming turbulent with increased filling velocity. 141

162 Figure 87. Filling velocity of CRC at 700 C for A206.2 Figure 88. Filling velocity of ECRC at 700 C for A

163 Figure 89 and Figure 90 show Hot Spot of A206 alloy at 700 C and 800 C of pouring temperatures. Hot spots for enhanced constrained rod casting are at the transitional points of feeder and constrained rods. At these junctions points, casting solidified last where hot tearing develops. At the pouring temperature of 800 C, Hot spot shifted to constrained A and B whereas at the pouring temperature of 700 C, the Hot Spot influenced on Constrained Rods A,B,C, and D. Figure 89. Hotspot of A206 alloy at 700 C pouring temperature 143

164 Figure 90. Hotspot of A206 alloy at 800 C pouring temperature COOLING AND SOLIDIFCATION OF ECRC MOLD Figure 91 and 92 show that constrained rods solidified at higher cooling rates than feeder as observed from simulation at pouring temperatures of 700 C and 760 C. Figure 92 shows that volumetric solidification shrinkage was observed at 7.68 % at the end of solidification for Al-Cu alloy at 700 C. At 760 C, the shrinkage was observed to be 8.33 % at the end of solidification as shown in Figure 94. When the pouring temperature was 144

165 increased, the solidification increased by 7.8%. Thus, the pouring temperature influenced the shrinkage. The theoretical value of Al-4.5%Cu showed a shrinkage value of 6.3%. A206.2 alloy exhibited higher shrinkage due to addition of 4.7%Cu content. The shrinkage of 7.68% is close to a theoretical value of 6.3%. Figure 91. Cooling and solidification of A206.2 alloy at 700C 145

166 Figure 92. Solidification shrinkage of A206.2 alloy at 700 C Figure 93. Cooling and solidification of A206.2 alloy at 760 C 146

167 Figure 94. Solidification shrinkage of A206.2 at 760 C Al-10%Si showed shrinkage of 4.61% at casting temperature of 700 C as shown in Figure 95. When the pouring temperature was increased to 750 C, the shrinkage was only increased by 4.97% as shown in Figure 96. At the pouring temperate of 800 C, the shrinkage was by 5% as shown in Figure 97. The theoretical value of shrinkage, 3.8%, compare to the shrinkage obtained from simulation, 4.61%, is very close. The shrinkage of 4.61% correlates with theoretical value of 3.61% because of the variation in the Si content. Table 20 shows the theoretical values of Shrinkage for different alloys. A higher % of Si content reduces the shrinkage porosity in Al-Si alloy. 147

168 Table 19. Shrinkage [42] Figure 95. Solidification shrinkage Al-Si alloy at 700 C 148

169 Figure 96. Solidification shrinkage Al-Si alloy at 750 C Figure 97. Solidification shrinkage Al-Si alloy at 800 C 149

170 5.3 SEQUENTIAL COUPLED TRANSIENT THERMAL STRUCTURAL METHODS The sequential coupling method for Transient Thermal Analysis were used in simulating the Enhanced Constrained Rod Casting (ECRC) as schematic illustrated in Figure 98 [70]. Figure 98. Transient thermal analysis using ANSYS sequential coupling method [70] 150

171 5.4 THERMO-MECHANICAL MODELING AND SIMULATION For the Modeling of ECRC, the assumptions were made to simply the model. The longest constrained rod was modeled for casting simulation since as it takes longer to solidify. The dimensions and thickness of Enhanced Constrained Rod Casting were used in modeling the 2D thermo-mechanical model in order to represent the actual model in a simulation of the casting. The simulation model was divided into regions A1 and A2 as shown in Figure 99. The meshing of the model was done with Plane55 2-D thermal solid elements as shown in Figure 100. The Plane55 has 4 nodes with a single degree of freedom and temperature at each node. It works as a 2D thermal conduction element and has orthotropic material properties [70]. Figure 99. Model of ECRC for thermal and structural analysis [70] 151

172 Figure 100. Plane55 2D thermal solid element [70] The material properties were assigned to elements of both regions A1 and A2. The material properties of steel (H13 Tool Steel) were assigned to elements of region A of the book mold. Material properties for casting including aluminum A206, A380, and Test-A (Kally) were assigned to elements of region A2 of the mold cavity respectively. Magnesium alloy AT72 was assigned to element of region A2 for casting simulation as shown in Figure 101. Figure 101. Model with plane55 thermal solid element [70]. 152

173 5.5 CONSTITUTIVE EQUATIONS AT SOLID STATE Modeling of the mushy zone is a complex and two phase liquid-solid phenomena. The mechanical response in the casting depends on the microstructural evolution, which integrates heat of fusion and solidification casting between liquidus and solidus temperatures. To overcome these issues, the study of the mushy zone is simplified while thermal and structural studies are conducted at solid state. The Energy equation was solved in order to determine the temperature distributions of the casting and mold during casting process. Assuming the pouring velocity of liquid metal is equal to zero. The temperature, (T) with respect to time (t) in two dimensional for mold and cavity, the energy balance equation is [66,70] [K T x x ] + [K T x y ] + Q = ρ C T y p t (1) Where, Q = Heat Source K x, K y = Thermal Conductivity in x and y-directions C p ρ = Specific heat = density 153

174 For thermal analysis, the compatibility equation relates to strains with incremental displacements. The mechanical equations relate to incremental forces which resulted in incremental stresses. The constitutive equations relate to incremental stress and strain. These equations are given as follows [ 66,70]: ε = [A] u (2) Where, x [A] = 0 [ 0 y] σ = [D]ε e (3) Where, A, D = stifness matrix ε e = elastic strain [D] = [ K11 K12 K13 K14 ] The total strain is given by the following equation: ε T = ε el + ϵ in + ε th (4) Where, ε T ε el = Total strain = Elastic component 154

175 ε pl ε th = Plastic component = Thermal component 5.6 BOUNDARY CONDITIONS AND MATERIAL PROPERTIES OF CASTING ALLOYS Figure 102 shows the boundary conditions that were applied to the thermal and structural analysis simulations of A206, A380, Test-A-alloy, and AT72 alloys. For thermal analysis simulation, a thermal convection coefficient of 25 W/m².K was applied at the outer surface of the book mold for heat transfer between ambient temperature (300K) and steel for all thermo-mechanical simulation. The mold temperatures was applied to nodes of region A1 for steel properties. The casting temperatures were applied to nodes of region A2 for A206, A380, Test-A-alloy, and AT72 alloys as pouring temperatures. For structural analysis, point A was fixed for (X,Y) direction of coordinates assuming C-clamp was applied to hold both left and right book mold assembly in order to prevent any displacements during cooling and contraction of casting. Point B was fixed only in X- direction of the coordinate to resist linear contraction of casting as shown in Figure 102. In the experimental analysis, the restrain ball provided this similar function by anchoring the rod to resist the contraction force during casting solidification. 155

176 Figure 102. ECRC model with boundary conditions [70] Table 21 and Table 22 show the physical properties of A206, A380, Test-A-alloy, and AT72 alloys for thermal and structural analysis and simulation. Table 23 shows the physical properties of P20 Tool Steel that was used for mold in analysis and simulation. The casting and mold temperatures that were used in experimental studies and thermalstructural analysis for A206, A380, Test-A-alloy, and AT72 alloys are presented in Table 24 and Table

177 Table 20. Physical Properties of A206, A380, Test-A Alloy and AT72 [65, 66] Table 21. Physical Properties of A380 [66] Temperature ( C) Density (Kg/m³) Thermal Expansion Coefficient 2.20E E E E E E

178 Table 22. Physical Properties of P20 Tool Steel [66] Table 23. Pouring Temperatures for Alloys Alloys for Casting Pouring Temperatures ( C) A Test-A-alloy A AT

179 Table 24. Mold Temperatures for Alloys Alloys Mold Temperatures ( C) Pouring temperatures A206, Test-A-alloy, A AT After setting the boundary conditions for thermal analysis, the transient parameters are setup as follows: Table 25. Simulation Time Steps Transient Simulation Steps 3600 Seconds Time step 10 seconds Maximum Time Step 900 seconds Minimum Time Step 3.6 seconds From the thermal casting simulation of the A206 alloy at pouring temperatures of 700 C, 760 C, and 800 C, it was observed that the constrained rod solidified faster than the feeder because of thinner section of casting. For structural analysis simulation, the loads were used from thermal analysis. The constraints were applied as described in Figure 102 to secure the mold in studying thermal strain and material deformation. From simulation, 159

180 thermal strain and distributions at the junction point between the feeder and constrained rod were observed at pouring temperatures of 700 C, 760 C, and 800 C as shown in Figure 103, Figure 104, and Figure 105. The total thermal and mechanical strain are summarized in Table 27. Table 26. Total Thermal and Mechanical Strain of A206 Alloy Die casting Alloys Thermal and Mechanical Strain (700 C) Thermal and Mechanical Strain(760 C) Thermal and Mechanical Strain (800 C) A Figure 103. Thermal strain for A206 alloy at 700 C 160

181 Figure 104. Thermal strain for A206.2 alloy at 760 C Figure 105. Thermal strain for A206 alloy at 800 C 161

182 The thermal and structural simulations of Test-A (K-alloy) were performed at pouring temperatures of 700 C, 750 C, and 800 C. The thermal strain and distribution at the junction point between feeder and constrained rod were observed at pouring temperatures 700 C, 760 C, and 800 C as shown in Figure 106, Figure 107, and Figure 108. The total thermal and mechanical strains are summarized in Table 28. Table 27. Total Thermal and Mechanical Strain of Test-A Alloy Die casting Alloys Thermal and Mechanical Strain (700 C) Thermal and Mechanical Strain (750 C) Thermal and Mechanical Strain (800 C) Test-A-alloy Figure 106. Thermal strain for Test-A Alloy at 700 C 162

183 Figure 107. Thermal strain for Test-A alloy at 750 C Figure 108. Thermal strain for Test-A alloy at 800 C The thermal and structural simulations of A380 were performed at pouring temperatures of 700 C, 750 C, and 800 C. The thermal strains at junction point between feeder and constrained rod were observed at pouring temperatures 700 C, 760 C, and 800 C as shown 163

184 in Figure 109, Figure 110, and Figure 111. The total thermal and mechanical strains are summarized in Table 29. Table 28. Total thermal and mechanical strains of A380 Die casting Alloys Thermal and Mechanical Strain (700 C) Thermal and Mechanical Strain (750 C) Thermal and Mechanical Strain (800 C) A Figure 109. Thermal strain for A380 alloy at 700 C 164

185 Figure 110. Thermal Strain for A380 alloy at 750 C Figure 111. Thermal strain for A380 alloy at 800 C 165

186 The thermal and structural simulations of AT72 alloy were performed at pouring temperatures of 675 C, 720 C, and 750 C. The thermal strain at junction point between the feeder and constrained rod were observed at pouring temperatures 700 C, 760 C, and 800 C as shown in Figure 112, Figure 113, and Figure 114. The total thermal and mechanical strains are summarized in Table 30. Table 29. Total Thermal and Mechanical Strain of AT72 Die casting Alloys Thermal and Mechanical Strain (700 C) Thermal and Mechanical Strain (750 C) Thermal and Mechanical Strain (800 C) AT Figure 112. Thermal strain for AT72 alloy at 675 C 166

187 Figure 113. Thermal strain for AT72 alloy at 710 C Figure 114. Thermal strain for AT72 alloy at 750 C 167

188 From the simulation, the maximum total thermal and mechanical strain were observed at the junction between the sprue and constrained rod. The summary of thermal strains for A206, Test-A-alloy, A380, and AT72 alloys are presented in Table 31. Table 30. The summary of total maximum thermal strain Temperature Thermal Strain - Thermal Strain Test-A Thermal Strain - AT72 Thermal Strain - A380 A

189 5.7 HOT TEARING PREDICTIVE MODELING The Niyama criterion is the ratio of thermal gradient to the square root of cooling rates. The Niyama is defined as [73] N y = G T Where G is the thermal gradient and T is the cooling rate. Niyama criterion is based on strain rate. Figure 115 shows the shrinkage porosity of 5.0% at 700 C. The shrinkage Porosity of 6.0 % was observed at 760 C as shown in Figure 116. Shrinkage porosity occurred at transition between thicker and thinner sections (at the junction between rods and feeder) due to differential cooling. The simulation results showed that the shrinkage porosity increased when pouring temperature was increased. 169

190 Figure 115. Shrinkage porosity of A206.2 alloy at 700 C Figure 116. Shrinkage porosity of A206.2 alloy at 760 C 170