Rapid Infrared Thermal Processing of AA 2618 and AA 6061 Forgings a Dissertation Presented to the Faculty of the Russ College Of

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Rapid Infrared Thermal Processing of AA 2618 and AA 6061 Forgings

A dissertation presented to

the faculty of

the Russ College of Engineering and Technology of Ohio University

In partial fulfillment

of the requirements for the degree

Doctor of Philosophy

Vamadevan Gowreesan

November 2008

This dissertation titled

Rapid Infrared Thermal Processing of AA 2618 and AA 6061 Forgings

VAMADEVAN GOWREESAN

has been approved for

the Department of Mechanical Engineering

and the Russ College of Engineering and Technology by

Frank F. Kraft

Associate Professor of Mechanical Engineering

Dennis Irwin

Dean, Russ College of Engineering and Technology

ABSTRACT

GOWREESAN, VAMADEVAN, Ph.D., November 2008, Integrated Engineering

Rapid Infrared Thermal Processing of AA 2618 and AA 6061 Forgings (150 pp.)

Director of Dissertation: Frank F. Kraft

Application of rapid heating methods for the thermal processing of aluminum alloy forgings is investigated in this work. Differential scanning calorimeter analysis of rapid heating of wrought aluminum alloys and laboratory simulation of solution treatment of wrought aluminum alloys were performed. The results showed that rapid heating accelerated the solutionizing of second phases in the alloys. Based on the results of these studies and practical limitation at the shop floor, optimized thermal processing cycles with an infrared furnace for AA 2618 and AA 6061 forgings were determined. Prototype forgings were produced and thermally processed with the proposed infrared heating cycles in addition to prototype forgings processed with conventional heat treatment.

Mechanical testing was performed from samples obtained from the prototype forgings and measured mechanical properties were compared to understand the effect of thermal processing for shorter duration with rapid infrared heating. The results showed that short solution treatment with rapid infrared heating can produce forgings with mechanical properties comparable to that of conventionally processed forgings.

The energy savings resulting from the short solution treatment with rapid infrared heating are also analyzed. The acceleration of the dissolution of second phase particles with rapid infrared heating has been attributed to the thermal stresses arising from mismatch of thermal expansions of the aluminum matrix and the second phases during rapid heating.

Approved:______

Frank F. Kraft

Associate Professor of Mechanical Engineering

To my wife,

parents,

brothers,

And also to all those teachers starting with the one who transformed an alphabet-hating five year old boy into one who

could write,

to those who teach 800 level courses.

ACKNOWLEDGEMENTS

I am greatly indebted to my academic advisor Dr. Frank Kraft for his guidance

and encouragement in my endeavors. What I learned from him helps me in my

professional work very much.

I am grateful for Dr. Khairul Alam, Dr. Daniel Gulino, Dr. Hugh Richardson and

Dr. Horatio Castillo for agreeing to serve on the dissertation steering committee. Their

input and suggestions brought some additional insights into the work.

This research work would not have been possible without the financial support by

EMTEC (Edison Materials Technology Center), Forging Industry Association (FIA) and

John C. Baker Fund of Ohio University. Their financial contribution and support are greatly appreciated. I would like to acknowledge other project partners for their

contributions, namely Rob Mayer of Queen City Forging Company, Puja Kadolkar of

Oak Ridge National Laboratory, Percy Gros of EMTEC and George Mochnal of Forging

Industry Association. Rob Mayer deserves a special appreciation for sharing some of his

results with us. It had been a great privilege and pleasure working on this project. Dr.

Jay Gunasekera is also acknowledged for his involvement in this research work. I would

like to take this opportunity to thank Devon Polin for sharing part of the experimental work load in this project. Mihnea Anghelescu is also acknowledged for his help with

DSC work.

I wish to express my sincere appreciation to the ever-helpful Stephanie Walker and Randy Mulford.

TABLE OF CONTENTS Page

ABSTRACT ...... 3 ACKNOWLEDGEMENTS ...... 6 LIST OF TABLES ...... 11 LIST OF FIGURES ...... 12 CHAPTER 1: INTRODUCTION ...... 19 1.0 Introduction ...... 19 1.1 Research Summary ...... 19 1.2 Background Information ...... 20 1.2.1 Forging Process ...... 20 1.2.2 Forging Stock ...... 22 1.2.3 Preheating of the Forging Stocks ...... 22 1.2.4 Forging Steps ...... 23 1.3 Heat Treatment ...... 23 1.3.1 Precipitation Hardening ...... 24 1.3.2 Grain Refinement ...... 26 1.3.3 Conventional Heating ...... 26 1.3.4 Infrared Heating ...... 27 1.4 Scope of the Work ...... 28 1.4.1 Theoretical Basis of Infrared Thermal Processing ...... 29 1.4.2 Determination of Optimal Rapid Infrared Thermal Processing for the Forgings ...... 29 1.4.3 Comparisons of Mechanical Properties ...... 30 1.5 Significance of the Work ...... 31 1.6 Objectives ...... 32 1.7 Outline of the Dissertation ...... 33 CHAPTER 2: Literature Review ...... 34 2.0 Introduction ...... 34 2.1 Thermal Processing of Alloys with Rapid Heating ...... 35 2.1.1 Rapid Heating of Titanium ...... 35

2.1.2 Rapid Heating of Steels ...... 37 2.2 Recent Applications of Rapid Heating Methods for Thermal Processing of Aluminum Alloys...... 38 2.3 Solution Heat Treatment ...... 41 2.4 DSC Studies on Thermal Processing of Aluminum Alloys ...... 45 CHAPTER 3: Savings in energy and processing time ...... 49 3.0 Introduction ...... 49 3.1 Modeling of Heat Transfer...... 49 CHAPTER 4: DSC Experimentation and analysis ...... 51 4.0 Introduction ...... 51 4.1 Differential Scanning Calorimeter ...... 51 4.2 Experimentation ...... 52 4.3 Chemical Composition of the Alloy Samples ...... 54 4.4 Results and Analysis ...... 56 4.4.1 Net Energy Transfer at Different Temperatures ...... 57 4.4.2 Variation of Specific Heat Capacities ...... 62 4.5 Conclusions ...... 64 CHAPTER 5: Physical Simulation of Solution Treatment with Rapid Heating ...... 66 5.0 Introduction ...... 66 5.1 Materials ...... 66 5.2 Experimental Procedure ...... 67 5.3 Results ...... 72 5.3.1 Results of Solutionizing Tests of AA 6061 ...... 72 5.3.2 Results of Solutionizing Tests of AA 2618 ...... 76 5.4 Metallographic Evolution of AA 6061 with Solution Treatment ...... 78 5.5 Results ...... 84 5.6 Metallographic Evolution of AA 2618 with Solution Treatment ...... 85 5.7 Conclusions ...... 92 CHAPTER 6: Prototype forging production and sample extraction ...... 94 6.0 Introduction ...... 94 6.1 Aluminum Alloy Materials ...... 94

6.2 Forging ...... 96 6.3 Infrared Furnace ...... 99 6.4 Solution Treatment...... 99 6.5 Sample Extraction ...... 101 CHAPTER 7: Experimental methodology ...... 104 7.0 Introduction ...... 104 7.1 Hardness Testing ...... 104 7.2 Tensile testing ...... 105 7.3 High Cycle Fatigue Testing ...... 107 7.3.1 R.R. Moore Testing Machine...... 107 7.3.2 Sample Preparation ...... 109 7.3.3 Theory ...... 111 7.4 Low Cycle Fatigue Testing ...... 114 7.4.1 Loading Cycles ...... 116 CHAPTER 8: Results ...... 118 8.0 Introduction ...... 118 8.1 Results of Testing of AA 2618 Forgings ...... 118 8. 1. 1 Mechanical Property Comparisons ...... 119 8.1.2 High Cycle Fatigue Testing ...... 120 8.1.3 Low Cycle Fatigue ...... 121 8.2 Results of Testing of AA 6061 Forgings ...... 123 8.2.1 Mechanical Property Comparisons ...... 124 8.2.2 High Cycle Fatigue ...... 126 8.2.3 Low Cycle Fatigue ...... 128 8.3 Grain Size Measurement of Conventionally Treated and IR Solutionized AA 2618 Forging Prototypes ...... 130 8.4 Hardness and Grain Size Measurements of AA 2618 Prototype Part Production ...... 131 8.5 Conclusions ...... 134 CHAPTER 9: DISCUSSION ...... 136 9.1 Summary of the Work ...... 136

9.2 Discussion ...... 137 9.2.1 Direct Energy Savings ...... 137 9.2.2 Second Phase Particles and Control over Grain Size...... 138 9.2.3 Improvement of Mechanical Properties ...... 141 9.3 Conclusions ...... 142 9.4 Recommendation for Future Work ...... 143 REFERENCES: ...... 145

LIST OF TABLES

Table 4.1: Composition of AA2618 test material in weight percentage, via atomic emission spectroscopy. The balance is Al...... 55

Table 4.2: Composition of AA661 test material in weight percentage, via atomic emission spectroscopy. The balance is Al ...... 55

Table 4.3: Conventional solution temperatures and precipitates present in the alloys under investigation...... 56

Table 8.1: Tensile and hardness test values for infrared solution treated AA 2618 samples and conventionally solution treated AA 2618 samples. All samples underwent a standard aging cycle also. Each sample was taken from a different forging upset of that lot. The bottom entries in the table are handbook specifications for comparison (Brown and Setlak, Code 3213 2003)...... 119

LIST OF FIGURES

Figure 1: Microstructure of a conventionally processed AA 2618 alloy forging. The determined optimum rapid infrared thermal processing is to be applied in the commercial production of this part (Kadolkar 2004)...... 22

Figure 2: Phase diagram of a hypothetical system A-B (Brooks 1991)...... 24

Figure 3: Infrared heater at working condition (http://www.infraredheating.com/die_pre- heater.htm 2002,1996)...... 27

Figure 4: DSC with the liquid Nitrogen tank...... 52

Figure 5: Heat flow rate for unit mass of AA 2618 during the heating cycle with a heating rate of 5 ºC/min. The corresponding temperature history is also shown...... 58

Figure 6: Heat flow rate for unit mass of AA 2618 during the heating cycle with a heating rate of 20 ºC/min. The corresponding temperature history is also shown...... 59

Figure 7: Heat flow rate for unit mass of AA 6061 during the heating cycle with a heating rate of 5 ºC/min. The corresponding temperature history is also shown...... 60

Figure 8: Heat flow rate for unit mass of AA 6061 during the heating cycle with a heating rate of 20 ºC/min. The corresponding temperature history is also shown...... 61

Figure 9: Variations of change of cp of AA 2618 at two heating rates during thermal cycles with the same maximum temperature...... 63

Figure 10: Variations of change of Cp of AA 6061 at two heating rates during thermal cycles with the same maximum temperature...... 64

Figure 11: Extruded bars and coupons obtained from them for solutionizing trails. The coupons were 19mm in diameter and 4.8mm thick...... 68

Figure 12: Thermal cycle data for AA6061 sample that was solutionized for 1200 s at 532ºC and then quenched (Gowreesan, et al. 2006). K-type 30 AWG thermocouple wires were attached to the specimen surface to measure temperature during heating and quenching...... 69

Figure 13: Typical quenching cycle for samples heated to 532 ºC (Gowreesan, et al. 2006)...... 70

13 only includes measurements on “solutionized” samples. The conductivity of as- extruded/un-solutionized material was 29.8 MS/m...... 73

Figure 15: Average hardness of AA6061 as a function solutionizing time and test solution temperature (Gowreesan, et al. 2006). Error bars represent ±1 standard deviation. Each data point is an average of 8 measurements on the sample...... 74

Figure 16: Average electrical conductivity of AA 6061 as a function time and solution test temperature for samples aged at 175ºC for 8 hours (Gowreesan, et al. 2006). Error bars represent ±1 standard deviation. Each data point is an average of 5 measurements on the sample...... 75

Figure 17: Average hardness of AA6061 as a function solutionizing time and solution test temperature for samples aged at 175ºC for 8 hours (Gowreesan, et al. 2006). Error bars represent ±1 standard deviation. Each data point is an average of 8 measurements on the same sample...... 75

Figure 18: Average electrical conductivity of AA 2618 after solutionizing and after aging, as a function of solutionizing time and solution test temperature (Gowreesan, et al. 2006). Error bars represent ±1 standard deviation. Each data point is an average of 5 measurements on the sample...... 77

Figure 19: Average hardness of AA2618 after solutionizing and after aging, as a function solutionizing time and solution test temperature (Gowreesan, et al. 2006). Error bars represent ±1 standard deviation. Each data point is an average of 8 measurements on the sample...... 77

Figure 20: Micrograph of as-extruded AA 6061 coupon on plane parallel to the extrusion direction (Mayer, et al. 2007) ...... 79

Figure 21: Micrograph of as-extruded AA 6061 coupon on plane perpendicular to the extrusion direction (Mayer, et al. 2007) ...... 79

Figure 22: Micrograph of AA6061 coupon after solutionizing at 532 ºC for 20 seconds. Taken in plane parallel to the extrusion direction (Mayer, et al. 2007) ...... 80

Figure 23: Micrograph of AA6061 coupon after solutionizing at 532 ºC for 20 seconds. Taken in plane perpendicular to the extrusion direction (Mayer, et al. 2007) ...... 80

Figure 24: Micrograph of AA6061 coupon after solutionizing at 532 ºC for 900 seconds. Taken in plane parallel to the extrusion direction (Mayer, et al. 2007) ...... 80

Figure 25: Micrograph of AA6061 coupon after solutionizing at 532 ºC for 900 seconds. Taken in plane perpendicular to the extrusion direction (Mayer, et al. 2007) ...... 80

Figure 26: Micrograph of AA6061 coupon after solutionizing at 552 ºC for 20 seconds. Taken in plane parallel to the extrusion direction...... 81

Figure 27: Micrograph of AA6061 coupon after solutionizing at 552 ºC for 20 seconds. Taken in plane perpendicular to the extrusion direction...... 81

Figure 28: Micrograph of AA6061 coupon after solutionizing at 552 ºC for 900 seconds. Taken in plane parallel to the extrusion direction...... 81

Figure 29: Micrograph of AA6061 coupon after solutionizing at 552 ºC for 900 seconds. Taken in plane perpendicular to the extrusion direction...... 81

Figure 30: Micrograph of AA6061 coupon after solutionizing at 572 ºC for 20 seconds. Taken in plane parallel to the extrusion direction (Mayer, et al. 2007) ...... 82

Figure 31: Micrograph of AA6061 coupon after solutionizing at 572 ºC for 20 seconds. Taken in plane perpendicular to the extrusion direction (Mayer, et al. 2007) ...... 82

Figure 32: Micrograph of AA6061 coupon after solutionizing at 572 ºC for 900 seconds. Taken in plane parallel to the extrusion direction (Mayer, et al. 2007) ...... 82

Figure 33: Micrograph of AA6061 coupon after solutionizing at 572 ºC for 900 seconds. Taken in plane perpendicular to the extrusion direction (Mayer, et al. 2007) ...... 82

Figure 34: Variation of grain size with time at different solutionizing temperatures. Grain sizes were measured in planes perpendicular to extrusion direction...... 83

Figure 35: Variation of intercept or thickness of fibrous grains with solutionizing time. The grain thickness was measured from micrographs from planes parallel to the direction of extrusion...... 84

Figure 36: Micrograph of AA2618 coupon after solutionizing at 530 ºC for 20 seconds (Mayer, et al. 2007). Taken in plane parallel to the extrusion direction...... 86

Figure 37: Micrograph of AA2618 coupon after solutionizing at 530 ºC for 20 seconds (Mayer, et al. 2007). Taken in plane perpendicular to the extrusion direction ...... 86

Figure 38: Micrograph of AA2618 coupon after solutionizing at 530 ºC for 900 seconds (Mayer, et al. 2007). Taken in plane parallel to the extrusion direction...... 86

Figure 39: Micrograph of AA2618 coupon after solutionizing at 530 ºC for 900 seconds (Mayer, et al. 2007). Taken in plane perpendicular to the extrusion direction...... 86

Figure 40: Micrograph of AA2618 coupon after solutionizing at 535 ºC for 20 seconds. Taken in plane parallel to the extrusion direction...... 87

Figure 41: Micrograph of AA2618 coupon after solutionizing at 535 ºC for 20 seconds. Taken in plane perpendicular to the extrusion direction...... 87

Figure 42: Micrograph of AA2618 coupon after solutionizing at 535 ºC for 20 seconds. Taken in plane perpendicular to the extrusion direction...... 87

Figure 43: Micrograph of AA2618 coupon after solutionizing at 535 ºC for 900 seconds. Taken in plane perpendicular to the extrusion direction...... 87

Figure 44: Micrograph of AA2618 coupon after solutionizing at 540 ºC for 20 seconds (Mayer, et al. 2007). Taken in plane parallel to the extrusion direction...... 88

Figure 45: Micrograph of AA2618 coupon after solutionizing at 540 ºC for 20 seconds (Mayer, et al. 2007). Taken in plane perpendicular to the extrusion direction...... 88

Figure 46: Micrograph of AA2618 coupon after solutionizing at 540 ºC for 900 seconds (Mayer, et al. 2007). Taken in plane parallel to the extrusion direction...... 88

Figure 47: Micrograph of AA2618 coupon after solutionizing at 540 ºC for 900 seconds (Mayer, et al. 2007). Taken in plane perpendicular to the extrusion direction...... 88

Figure 48: Micrograph of AA2618 coupon after solutionizing at 545 ºC for 20 seconds (Mayer, et al. 2007). Taken in plane parallel to the extrusion direction...... 89

Figure 49: Micrograph of AA2618 coupon after solutionizing at 545 ºC for 20 seconds (Mayer, et al. 2007). Taken in plane perpendicular to the extrusion direction...... 89

Figure 50: Micrograph of AA2618 coupon after solutionizing at 545 ºC for 900 seconds (Mayer, et al. 2007). Taken in plane parallel to the extrusion direction...... 89

Figure 51: Micrograph of AA2618 coupon after solutionizing at 545 ºC for 900 seconds (Mayer, et al. 2007). Taken in plane perpendicular to the extrusion direction...... 89

Figure 52: Variation of grain size on planes perpendicular to the extrusion direction ..... 90

Figure 53: Variation of vertical intercept with solutionizing time. The grain intercept was measured from micrographs from planes parallel to the direction of extrusion...... 91

Figure 54: Variation of horizontal intercept with solutionizing time. The grain intercept was measured from micrographs from planes parallel to the direction of extrusion...... 92

Figure 55: Longitudinal macro-structure of AA2618 forging billet in the as-extruded condition (Mayer, et al. 2007). The photo shows a fine, recrystallized grain structure. Extrusion direction is horizontal to page. Photo was taken at ORNL...... 95

Figure 56: Longitudinal and cross sectional micro-structure of AA6061 forging billet in the as-extruded condition (Mayer, et al. 2007). The photo shows a fine, elongated grain structure in the direction of extrusion. Extrusion direction is horizontal to page...... 96

Figure 57: Continuous flatbed infrared furnace (right) and forging billet (left-foreground) at the Queen City Forging Company (QCF). The forging press with die heater is shown in the background (Used with permission of QCF)...... 97

Figure 58: Longitudinal macro-structure of AA2618 billet after upsetting 50% at 427 ºC (800 ºF) (Mayer, et al. 2007). The grain structure remained similar to that of the as- extruded condition. Pre-heating of this sample was via conventional methods. Extrusion and forging directions are vertical to the page. Photo was taken at ORNL...... 98

Figure 59: Longitudinal macro-structure of AA6061 billet after upsetting 50% between 430 ºC-480 ºC (806 ºF-896 ºF) (Mayer, et al. 2007). The grain structure remained similar to that of the as-extruded condition. Pre-heating of this sample was via conventional methods. Extrusion and forging directions are vertical to the page. Photo was taken at ORNL...... 99

Figure 60: Sample extraction from the prototype forging (Mayer, et al. 2007). One 12.7mm (0.5 inch) diameter rod for a tensile sample and nine 11 mm (7/16 inch) diameter rods for fatigue samples were extracted from the forging prototype...... 102

Figure 61: Sample extraction from the prototype forging (Mayer, et al. 2007). One 12.7mm (0.5 inch) diameter rod for tensile sample and eight 11 mm (7/16 inch) diameter rods for fatigue samples were extracted from the prototype forgings...... 102

Figure 62: Photo shows rods that have been wire-EDM machined from an AA 2618 forging upset that was conventionally heat treated (Mayer, et al. 2007). Rod samples were further machined into samples for tensile testing, and low and high cycle fatigue testing...... 103

Figure 63: Polished flat surface of the forging being tested for hardness...... 105

Figure 64: Tensile sample extracted and machined from the forging prototype. It has treaded sections at both ends to facilitate attaching the sample to the grips for loading. 106

Figure 65: Photograph of Instron, R.R. Moore fatigue testing machine...... 108

Figure 66: The rotation counter in the R.R. Moore fatigue tester. The black knob at the lower left is for changing the speed of revolution...... 109

Figure 67: Geometry of the high cycle fatigue samples...... 110

Figure 68: Measuring device to check the minimum diameter of the sample without damaging the surface. Nylon anvils attached to end of edges prevent damage of the aluminum sample surface...... 111

Figure 69: The free-body diagram of the high cycle fatigue sample and sample grips (Operating Instructions, Instron Model R.R.Moore High Speed Rotating Beam Fatigue Testing Machine n.d.). The sample is held with two sample grips at the ends and the free-body diagram of the combination of the sample and the grips is considered here. Two equal and parallel loads (W) are applied to the sample and sample grips combination. R corresponds to reactions at the supports...... 112

Figure 70: Low cycle fatigue sample fixed to the grips in the MTS machine. An extensometer is attached to the gauge section of the sample to measure and feedback the strain in the sample...... 115

Figure 71: A computer screen example of the dynamic variation of applied force and measured strain during low cycle fatigue testing...... 116

Figure 72: High cycle fatigue data of AA 2618 (Mayer, et al. 2007). Graph shows a comparison of high-cycle fatigue data for T61 treated specimens that were conventionally solution treated, IR solution treated and data from the Aerospace Structural Metals Handbook (Brown and Setlak, Code 3213, page 16, 2003). ASM Handbook endurance limit (Bray 2002) is also shown...... 121

Figure 73: Low cycle fatigue data for conventionally heat-treated and rapid infrared treated AA 2618 samples (Mayer, et al. 2007)...... 122

Figure 74: Cyclic stress-strain for AA 2618 T61. The stress and strain measurements were taken at half life loading cycle...... 123

Figure 75: High cycle fatigue data for AA6061-T6 (Mayer, et al. 2007). Graph shows a comparison of high-cycle fatigue data for T6 treated specimens that were conventionally solution treated, infrared solution treated and data from the Aerospace Structural Metals Handbook (Brown and Setlak, Code 3206, page 15, 2003). The data from ASM Handbook corresponds to drawn rod AA 6061-T6...... 127

Figure 76: Low cycle fatigue data for conventionally heat-treated and rapid infrared treated AA 6061 samples (Mayer, et al. 2007)...... 128

Figure 77: Cyclic stress- strain for AA 6061 T6. The stress and strain measurements were taken at half life loading cycle...... 129

Figure 78: AA 2618 micro-graphs showing the grain size in (a) an IR solution treated sample (530 ºC for 40 min) after ageing, and (b) a sample that was conventionally solution treated and aged (Mayer, et al. 2007). The section shown in (a) is parallel to the

18 forging direction and the section in (b) is transverse to the forging and extrusion directions...... 130

Figure 79: Hardness in forged and IR solution treated prototype parts as a function of time at temperature (Mayer, et al. 2007). Parts were given a conventional artificial ageing cycle. Error bars represent ±1 standard deviation. (Data provided by QCF) ..... 132

Figure 80: Variation of grain size in forged and IR solution treated prototype parts (Mayer, et al. 2007). Error bars represent ±1 standard deviation. (Data provided by QCF) ...... 133

CHAPTER 1: INTRODUCTION

1.0 Introduction

This dissertation presents work on thermal processing of Aluminum Association

(AA) alloys AA 2618 and AA 6061 forgings with rapid heating. The study includes

exploration of the solutionizing process with rapid heating, differential scanning

calorimetry (DSC) studies of rapid heating of the aluminum alloys, physical simulation of

the effect of rapid heating of aluminum alloys and comparisons of mechanical properties

of prototype forgings that underwent thermal processing with rapid heating with that of

conventionally processed prototype forgings.

This introductory section presents the background information on the research,

the objectives of the work and an overview of the rest of the dissertation.

1.1 Research Summary

Application of infrared heating for the thermal processing of aluminum forgings

was investigated in this research. Previous investigations in this area (Blue, et al. n.d.)

(Lu, et al. 2004) have shown that using rapid infrared heating for the processing of

aluminum alloy forgings can result in considerable energy savings and enhanced

mechanical properties for aluminum alloy AA 2618. However the full potential of rapid

infrared heating for thermal processing of aluminum forgings has not been realized due to the lack of experimental data and fundamental understanding. In other words, the

thermal processing has not been optimized to obtain improved mechanical properties at

the cost of minimum energy and time. In addition, lack of experimental data supporting

the benefits of using infrared heating for the thermal processing of aluminum forgings

20 hinders popularizing of this method; data supporting the benefits of infrared thermal processing needs to be obtained. Understanding the process at the fundamental level has to be achieved.

Hence the current work can be roughly divided into three areas of focus. The first one is the analysis and understanding of the process of solution heat treatment with high heating rates, at the thermodynamic/metallurgical level. The second objective is to determine the optimum rapid infrared heating processing parameters for two selected alloys, AA 2618 and AA 6061. Demonstration of enhancement of mechanical properties due to rapid infrared heating is the third objective.

In order to present a clear picture of this work, a review of the forging process, aluminum alloys, and heat treatment is presented in the following sections. This is followed by the specific objectives of this work and a short overview of the dissertation.

1.2 Background Information

1.2.1 Forging Process

In the forging process, the workpiece is plastically deformed in the solid state to the required shape by impacting it into dies having cavities of the required shape. Other

than flash, this process requires minimum material removal. The deformation associated

with the process refines the material, which in turn makes it desirable over other

fabrication methods. This work is concerned with hot forging, in which the workpiece is

heated to over about 60% of the absolute melting temperature of the alloy. Hot working

is done to improve the formability of the work piece.

Two types of forging processes are related to the current work. They are impression die forgings and open die forgings.

Impression die forging (closed-die forging): Dies with impression cavities of the desired shapes are used. As the forging press brings the dies together, the workpiece or forging stock is squeezed into the cavities of the dies. Complex shapes can be forged using these types of dies. Based on the size, shape and material, the forging process can be done at different temperatures.

Open die forging: Flat dies (anvil and hammer) without any impression cavities are used for this process. Open die forging is performed as an initial and intermediate process.

Developing desired grain flow is one of the reasons for this step. Shafts, rings, saddle/mandrel rings and hollow sleeve type forging are some of the other applications of open die forging (Forging Industry Association 1985).

Figure 1 shows a section of a commercial forged part for which infrared thermal processing will be applied in the near future. Even though, this particular product was produced by impression die forging, the prototype forgings for the current work are from open die forging. The forging stock for the part shown in Figure 1 was an extruded bar such as that to be used for the prototype forgings for this research. It shows the grain structure variation of a part that is made from AA 2618. This forging was processed with conventional thermal processing.

1.2.2 Forging Stock

The forging stock (or preform) can be an as-cast block, cast and machined block, an extruded bar or even a stock obtained from a previous forging process. As mentioned above, the forging stock used in this work is extruded bar.

1.2.3 Preheating of the Forging Stocks

Billets are preheated prior to the forging process to achieve the required formability in the material; the alloy’s flow stress decreases with increasing temperature.

The standard preheating temperatures are between 415 ºC – 455 ºC for AA 2618, and

430 ºC – 480 ºC for AA 6061 (Kuhlman 2002).

1.2.4 Forging Steps

Based on the complexity of the geometry of the part to be forged and the capacity of the available forging press, single or multiple forging steps may be required. That is, some geometries are so complex that they cannot be forged from the initial forging stock to the final shape in a single operation with a single set of dies. In some cases, the force required to forge the final product from the initial billet in a single step might be greater than the available force capacity of the forging press. In these cases, multiple step forging processes are employed. In multiple step forging, the initial billet is forged to an intermediate shape first. Then the intermediate forging is forged to another intermediate shape or the final shape. Depending on the temperature reduction during each step and the forging temperature range of the alloy, the intermediate forging might be reheated before the next forging process.

1.3 Heat Treatment

Heat treatment is a thermal process used to change the microstructure of the alloy and thereby to change the mechanical properties of the alloy. The selected two aluminum alloys used for this work are precipitation-hardening alloys. For precipitation-hardening aluminum alloys, the heat treatment consists of three processes: solution heat treatment, quenching and aging (Chandler 1996). The phenomenon behind the enhancement in the mechanical properties after heat treatment is precipitation hardening and this is treated in detail in the next section.

1.3.1 Precipitation Hardening

To provide a better understanding of the relationship between heat treatment and the resulting enhancement of the mechanical properties, a simple example is presented below (Brooks 1991). Figure 2 shows a portion of a phase diagram of a system A-B.

Figure 2: Phase diagram of a hypothetical system A-B (Brooks 1991).

Phase diagrams represent the relationship between equilibrium phases in an alloy and the state variables of the system such as the temperature and alloy composition. In

Figure 2, the percentage of component B in the system is represented by the abscissa while temperature is represented by the ordinate. Alloy systems represented by the state variables corresponding to the points in the region below the solvus line are of phases

α + θ whereas systems represented by the state variables corresponding to the points in the region above the solvus is of phase α only. The α-phase has the same crystallographic structure as element A, while the θ phase is an intermetallic phase of A

25 and B atoms. A system with 10% of B is taken as an example here. The vertical dotted line represents the phase variation of the selected system with the variation of the temperature. Provided the system is held at a temperature above T2 for sufficient amount

of time, a single α phase system will be achieved by dissolution of the θ phase (and B

atoms) into the primary solid α phase. If this system is then cooled slowly below T2, it

will transform into a system of α + θ as shown by the equilibrium phase diagram.

However if the system is rapidly cooled to ambient or room temperature (i.e., T1),

θ phase (a.k.a. 2nd phase) formation is suppressed because the dissolved solute atoms do

not have time to precipitate from solution. The resulting system is supersaturated with

the solute. This process is called solution treatment. The resulting system is also not in

equilibrium, but it is metastable. When this system is heated to an elevated temperature

below T2 or possibly kept at room temperature for sufficient time, precipitation of θ phase

dispersions (A + B intermetallic particles) will commence. When this process takes place

at a temperature above the ambient temperature, it is called “artificial aging”. When it

takes place at the ambient temperature, it is called “natural aging”. During the initial stage of the aging, the formation of fine participates dominates. The fine precipitates demonstrate coherency with the matrix (α) phase, increase the resistance to dislocation movement and thereby increase alloy strength. The reduction of strength due to the

elimination of B atoms in the solid solution is less than the increase in strength due the

formation of these fine second phase precipitates. If the aging “reaction” is allowed to

progress, the fine coherent precipitates will grow (or coalesce) and eventually lose

coherency with the matrix and a loss of material strength will ensue. This is called over- aging. This illustrates the necessity of optimizing the heat treatment.

1.3.2 Grain Refinement

Another type of strengthening mechanism is associated with grain size refinement. Note that strengthening or resistance to plastic deformation is achieved by anything that impedes the movement of dislocations in the crystal lattice. Grain boundaries impede dislocation motion. Thus smaller grain size, which results in more boundaries, can also increase material strength. If the deformation temperature is high enough and the forming conditions are favorable, recrystallization might occur during the deformation process. This is called dynamic recrystallization. If the solution treatment temperature is sufficient and the conditions are favorable to nucleation, then recrystallization may occur just after plastic deformation or during solution heat treatment. This is called static recrystallization. According to the Hall-Petch relationship, the strength is inversely proportional to the square root of the grain size.

Hence grain refinement will result in strength increase.

Since rapid infrared thermal processing is expected to enhance the mechanical properties, it is reasonable to expect precipitation hardening and/or grain refinement to be the mechanisms behind the enhancement.

1.3.3 Conventional Heating

Conventionally, gas fired furnaces have been used in the forging industry for preheating and heat treatment. Comparatively lower efficiency, long cycle time and poor control over the temperature are the disadvantages of these systems. Current forging and heat treatment standards practice is based on this heating technique.

1.3.4 Infrared Heating

Infrared heating is essentially a non contact heating method. High intensity heat lamps emit the heating radiation in this method. Ceramic emitters, quartz emitters, quartz lamps, quartz tubes, metal sheathed tubular and panel heaters are different types of heaters used. Figure 3 shows an infrared heater at its working condition.

Figure 3: Infrared heater at working condition (http://www.infraredheating.com/die_pre-heater.htm 2002,1996).

The efficiency of the energy conversion of electrical energy into heat radiation is

reported to be more than 90% in an infrared heater

(http://www.infraredheating.com/die_pre-heater.htm 2002,1996). The rate of heat

transfer for radiation is proportional to the fourth power of the temperatures (Stefan-

Boltzmann law) while the rate of heat transfer for convection involves only the first

28 power of the temperatures. In addition, in convection heating, heat transfer has to occur twice, first from the heat source to the air and then from the air to the part to be heated.

Hence infrared heating can be much faster and less costly than convection heating.

Additionally, infrared heaters have some operational advantages compared to other types of furnaces (http://www.infraredheating.com/die_pre-heater.htm 2002,1996).

These include short start-up time, cleaner method, and uniformity in heating. Since the electrical energy is directly converted to radiant heat, no by-products are produced.

Unlike combustion, there is no problem concerning environmental pollution at the plant level. The longer life of the heating element (above 5000 hours), reduced start up time, uniform and faster heating rate make infrared heaters very economical to operate at a commercial plant level.

1.4 Scope of the Work

The entire project can be roughly divided into three tasks. The first task was to review the theoretical basis of the phenomena of thermal processing with infrared heating. This includes identification of mechanisms responsible for mechanical property enhancement due to rapid infrared heat treatment. The effect of higher heating rates on solution heat treatment of aluminum forgings using differential scanning calorimetry

(DSC) as well as light microscopy is also investigated. Secondly, the optimum solution cycles with rapid heating had to be determined for each of the aluminum alloys selected.

Since the conventional solution treatment cycles had been designed for heating methods with slower heating rate, the new “optimum” solution treatment process for the selected alloys had to be determined. The third task was the comparison of the mechanical

29 properties of conventionally processed forgings with that of prototype forgings processed with the “optimal” rapid infrared solution cycles.

1.4.1 Theoretical Basis of Infrared Thermal Processing

As an emerging technology, the application of infrared thermal processing of aluminum alloy forgings has not been understood properly. The reason for superior enhancement of mechanical properties due to rapid infrared heat treatment had to be explored and understood. Since it was suspected that the enhancement of mechanical properties due to infrared heating was caused by the high heating rate associated with infrared heating, an extensive literature view was undertaken on the works related to the effect of heating rate and heating parameters on the solutionization of aluminum alloys.

Several existing models relating the heat treating parameters and solutionization were carefully scrutinized. In order to understand the effect of heating rate in a solution treatment, DSC analysis was performed on the selected alloys with two solution cycles with 5 ºC/minute and 20 ºC/minute heating rates. The results of these experiments were reviewed with respect to the results of the literature review.

1.4.2 Determination of Optimal Rapid Infrared Thermal Processing for the

Forgings

Increasing the solutionizing temperature closer to the melting temperature of the alloy will accelerate and improve the solutionizing process, however poor temperature control of conventional furnaces increases the risk of incipient melting during heat treatment. Infrared heating systems can offer better control of temperature. Hence, the

30 infrared furnace can extend the solution treating temperatures to higher values. The solutionizing behavior of AA 2618 and AA 6061 at these different temperatures has to be considered when selecting the suitable solutionizing parameters for infrared thermal processing. Hence, a set of physical simulations of solution treatment of AA 2818 and

AA 6061 with rapid heating was performed using a laboratory furnace. Aluminum alloy coupons were subjected to an array of solution cycles in a laboratory furnace. With the smaller mass of the coupon, high heating rates similar to that of infrared furnaces on larger parts were achieved. The degree of solutionizing due to these solution cycles were estimated with hardness and electrical conductivity measurements. The results were used to determine the ‘optimal’ infrared thermal processing cycles for the two alloys.

1.4.3 Comparisons of Mechanical Properties

A simple upsetting operation was performed to produce prototype parts for infrared thermal processing and subsequent testing. Billets were cut from 57 mm (2.25 in) extruded round bars. These billets were initially 152 mm (6 in) long and were upset to 76 mm (3 in) by axial compression. The pre-determined “optimum” infrared thermal

processing cycles were applied to the upsets. Mechanical properties and electrical conductivity of samples from these upsets were tested. First, uniaxial tensile tests were

performed to obtain the yield strength, elongation and ultimate tensile strength. Then, based on the yield stress values, the stress levels for fatigue tests were determined. An

R.R.Moore rotating beam fatigue tester was used for high cycle fatigue testing. A servo- hydraulic MTS machine was used for strain controlled low cycle fatigue testing. In addition, hardness testing was also performed. In a similar manner, conventionally

31 processed upsets were produced and their corresponding properties were measured for comparison.

1.5 Significance of the Work

Infrared heating has advantages over conventional heating methods. High efficiency and the speed with which uniform steady state is reached make it attractive from the stand point of energy costs. A cost savings up to 40-50% have been projected from a full scale industrial production set-up (Kadolkar 2004). Considering the magnitude of the global forging industry sales, a very significant energy and cost savings are expected as a result of commercialization of this method. However no standard process is available as opposed to established procedures with standard conventional heating. The available processes are relevant to heating methods with relatively slower heating rates and are not applicable to heating methods with relatively higher heating rates such as infrared heating.

The optimal solutionizing process includes optimum heating temperature and the optimum duration. When solutionizing at relatively higher temperatures, the required solutionizing duration is lower. However imprecise controllability of furnace temperature and the effects associated with temperatures near the eutectic temperature such as incipient melting prevent venturing near the eutectic temperature. The infrared furnace has shown the potential of more controllability of the temperature. The determination of optimal solutionizing is expected to make use of this aspect of infrared heating.

An understanding of the thermodynamics involving the infrared heating method will also help to determine the optimal infrared solutionizing cycles for other alloys without performing extensive experiments. For example, the understanding of the thermodynamics will make it possible to develop data for the development of new alloys using Material by Design® (http://www.questek.com/matbydesignhead.shtml 2007).

1.6 Objectives

The objectives of the work can be summarized as follows:

¾ To investigate the effects of heating rate on solution treatment of aluminum alloy

forgings.

¾ To determine optimal infrared solution heat treatment cycles for forged AA 2618

and AA 6061 alloys.

¾ To compare mechanical properties of conventionally processed forgings and

prototype forgings processed with the “optimum” rapid infrared solutionizing.

¾ To determine the reasons for the enhancement due to rapid infrared heat treatment

over conventional processing. Differential scanning calorimetry studies and light

microscopy will be used to evaluate the effect of higher heating rate on solution

treatment of the selected alloy forgings.

¾ To model the heat treatment of conventional gas fired furnace and rapid infrared

heating based on basic heat transfer principles. This will explore the time and

energy savings due to infrared heating.

1.7 Outline of the Dissertation

The dissertation is organized into nine chapters. The first chapter presents the

introduction. The next chapter covers literature review related to thermal processing with rapid heating, analytical modeling of solutionizing with rapid heating and DSC study of

rapid heating of aluminum alloys. In Chapter 3, the energy and time savings resulting

from the use of infrared heating is analyzed. The proposed effect of heating rate on solutionizing of aluminum alloys is investigated using DSC studies of aluminum alloys in

Chapter 4. Coupons of the two aluminum alloys were heated at two different heating rates and the effect of heating rates on the energy absorption/rejection by the alloys were compared and studied in this chapter. This is followed by a chapter on the physical simulation of the solution treatment of aluminum alloy forgings with rapid heating.

Based on the results from the physical simulation, ‘optimized solution treatment cycles’

were determined. This chapter is followed with a chapter on the prototype forging

production and sample extraction. The next chapter is on experimental methodology to

compare the mechanical properties of aluminum alloy prototype forgings processed with

rapid infrared heating with that of conventionally processed prototype forgings. The

mechanical properties evaluated include hardness, tensile properties, high cycle fatigue

and low cycle fatigue. The results of these tests are presented in Chapter 8. And the final chapter interrelates and discusses the results. In addition, it summarizes the conclusions of the work and recommendations for future work.

CHAPTER 2: LITERATURE REVIEW

2.0 Introduction

The driving factors behind the research in this field are the desire to reduce process energy consumption, to improve mechanical properties, and to develop new methods of heating and modifying alloys. For example, the possibility of using short solution treatments for strontium modified aluminum alloy has been suggested (Zhang,

Zheng and StJohn 2002). The research presented herein is one of the first attempts to use the infrared rapid heating method to optimize the thermal processing of aluminum alloy forgings. However the applications of rapid heating for processing of other metal parts have been explored earlier. Hence the application of rapid heating in the processing of titanium alloys and steels are reviewed first, followed by works on rapid heating in the processing of aluminum alloys.

Thereafter a detailed review of recent works on optimization and analysis of the heat treatment process of aluminum alloys are also presented, since the current work aspires to optimize thermal processing. Since this work involves experimentation, the experimental approach is also presented where appropriate. The heat treatment process involves annealing, solution treatment and ageing and the corresponding works are reviewed separately. The findings from this review were used for the experimental aspect of this research.

Differential scanning calorimetry (DSC) was used to investigate the effect of heating rate on the solutionizing of aluminum alloys. Hence previous works utilizing

DSC for investigation of thermal processing of aluminum alloys are also reviewed in this chapter.

2.1 Thermal Processing of Alloys with Rapid Heating

This research is one of the first attempts to use a rapid heating method for processing aluminum alloy forgings. However the application of rapid heating for processing of other alloys has been explored earlier.

2.1.1 Rapid Heating of Titanium

Ivasishin and Oshkaderov (1982) explored the effect of heating rate on hardening of titanium alloys VT23 and VT6. Their studies showed that the heating rate affected the phase transformation temperature and resulting grain size. Further, they found that the increase in heating rate resulted in a reduction of grain size.

Baulin and Smirnov (1993) attributed a change in phase transformation behavior of titanium with heating rate to the α to β transformation mechanism. The authors propose that at higher heating rates, the formation of nuclei with higher activation energies controls the phase transformation instead of the growth rate of the nuclei.

Ivashishin and Teliovich (1999) explored the application of rapid heating rates for thermal processing of titanium and steel alloys. Their investigations included how the heating rates affected the mechanisms and kinetics of the microstructural evolution in titanium and steel alloys. Their results showed that rapid heating could be used to obtain titanium and steel alloys with superior mechanical properties than that of conventionally processed alloys. Refinement of grain size and fine microstructural components that formed due to the rapid heating had been attributed to the superior mechanical properties.

In other work, Ivashishin and Lutjering (1993) compared the structure and mechanical properties of two high temperature titanium alloys after rapid heating and conventional furnace heat treatment. In addition, the effect of the type of quenching, namely water quenching and air cooling, were also studied. The rapid heating was done using the direct resistance heating method. The rapid heating rate used was 50 ºC/s.

Thermal processing with rapid heating produced much smaller β grains than conventional furnace heating. Consequently the samples that underwent rapid heat treatment exhibited better mechanical properties than conventionally heat treated samples.

Elagina, et al. (1984) rapidly annealed fine-grained lameller titanium alloys of different compositions at different heating rates. The heating rates used were between 10 and 300 ºC/s. Their results showed that by changing the heating rate and temperature, it was possible to change the final grain size of VT9 titanium alloy between 10μm and

130μm. In the same paper, the authors advocated for the use of rapid heating prior to deformation for the following reasons. The oxidization of the alloy is reduced since the time at elevated temperature is reduced. As the results had shown, grain refinement occurs and correspondingly improving mechanical properties. These arguments could be applicable to the preheating of aluminum blanks before deformation, even though this work does not deal extensively with this aspect of infrared heating of aluminum alloys.

In the next section, works on the application of rapid heating to the thermal processing of steels are reviewed.

2.1.2 Rapid Heating of Steels

Semiatin, Williams, and Bryer (1987) used a Gleeble testing machine to simulate and evaluate the rapid annealing behavior of steel sheet samples. The materials tested in this work were cold-rolled low-carbon steel, cold-rolled electrical steel and austenitic stainless steel (AISI 304). The annealing of electrical steel was aimed not only on improving mechanical properties but also on obtaining favorable electrical properties to minimize core losses by minimizing interstitial contents and by developing proper grain size and crystallographic texture. The mechanical properties resulting from rapid annealing were tested and compared with those of conventional annealing. The required properties were obtained by rapid annealing a very low carbon content steel to a temperature higher than that used in conventional annealing. The results with 0.05% C steels showed slightly better response to conventional processing than rapid annealing.

Never the less, the results with AISI 304 showed that rapid annealing and conventional processing produced comparable mechanical properties. However, the authors did not explore the reasons for this behavior. While this work was on rapid annealing of steel alloys and our current work is on aluminum alloys, it may be reasonable to expect similar behavior with aluminum alloys.

Recent works investigating the effect of heat treatment parameters on the properties of aluminum alloys are reviewed in the next section to obtain some guidelines on the direction of current research work.

2.2 Recent Applications of Rapid Heating Methods for Thermal Processing of

Aluminum Alloys

The move toward lean manufacturing has moved researchers and manufacturers to look for faster heat treatment methods to shorten heat treatment times (Keist 2005).

The fluidized bed is one of the methods being explored in the aluminum industry. In this method fine hard media (such as sand) particles suspended partially by a gas acts as a fluid media to transfer the heat to the part to be heat treated. This method is being successfully used for heat treatment of steel. The advantages of this method are the higher heating rate and its excellent control over the temperature. The heat transfer coefficient of fluidized bed systems is between 500 and 700 W/m ºC. Further, this method of heating provides excellent temperature uniformity in the part being heated.

The increased temperature control facilitates heat treating at higher temperatures closer to partial melting temperatures of the alloys. It has been shown that by moving the heat treatment temperature closer to the partial melting temperature, the heat treating time could be reduced without compromising the effect of the heat treatment (Keist 2005)

(Apelian, Shivakumar and Sigworth 1990) (Shivkumar, et al. 1990). In addition, it is also reported that the uniformity in temperature within the part during quenching with fluidized bed is better than that during quenching with water due to the vapor blanket formed during quenching with water. The aging time can also be reduced by using fluidized bed.

Traditionally fluidized bed systems have been batch systems and therefore have been bottle-necks in the production lines. Construction and operation of continuous

39 fluidized bed systems are also currently being explored (Apelian, et al. 2003)

(Chaudhary, et al. 2003) (Keist and Bergman 2003).

Apelian, et al. (2003) performed research to look into the effect of fluidized bed heat treatment on aluminum. Cast aluminum alloy A356 was used for the work. They found more favorable changes in microstructure in alloys processed in a fluidized bed system than that processed in a conventional furnace. Mechanical testing showed a slight enhancement in the ultimate tensile strength and ductility of samples heat-treated in a fluidized bed system.

In an attempt to optimize the fluidized bed heat treatment process for A356.2 castings, tensile properties were obtained as a function of soaking time for solutionizing and aging (Keist and Bergman 2003). An array of samples was tested after solution treating at 540 ºC and 550 ºC with holding times from 15 minutes to 3 hours and then aging at 182 ºC for 2 hours. It should be noted that solution treating at 550 ºC has not been possible for A356 with conventional furnaces due to its proximity to the solidus temperature and the lack of control over temperature. Another set of samples (A356.2 aluminum alloy modified with 0.02% Sc) was tested after solution treating at 550 ºC for

30 minutes and then aging at 154 ºC, 182 ºC and 210 ºC for holding times from 15 minutes to 6 hours. The results showed that solution heat treatment at 550 ºC for shorter time resulted in mechanical properties similar to that obtained by conventional furnace solution heat treatment with longer heating durations. The review of this past work on a casting alloy offers some guidelines to this research, which involves optimizing the thermal processing of aluminum forgings with the rapid infrared heating method. In contrast, for this project it was decided to use physical simulation by rapidly heating

40 small samples in a laboratory furnace and then measure the resulting properties with measurements of microhardness and electrical resistance.

The microstructural evolution of the A356 alloy system with short cycle solution treatment with fluidized bed furnace was studied by Chaudhury, et al. (2003). A fluidized bed with ± 3 ºC temperature variation was used in this study. The solutionzing was performed at 538 ºC/540º C or at 554º C. Quenching was performed with water at room temperature. The ageing temperatures were 171 ºC, 190 ºC and 204 ºC and various ageing times were employed. Optical microscopy and scanning electron microscopy were performed on samples after solutionizing and after a complete heat treatment cycle including solutionizing and ageing. The analysis showed that the fibrous eutectic silicon particles fragmented, coarsened and spherodized with the thermal processing. This process was faster with solutionizing with a fluidized bed than with a conventional furnace. Also, the spheroids were more equiaxed when obtained with a fluidized bed furnace. The authors attributed this behavior to the difference in the thermal expansion coefficients between the silicon and the aluminum matrix. As a result, larger thermal stresses were induced on the silicon-aluminum matrix interface during rapid heating.

This gave rise to more dislocations/defects, assisting diffusion of atoms and enhancing the kinetics. Quantitative analysis showed that the number fraction of silicon particles with an aspect ratio lower than 1.5 achieved by 15 minutes of fluidized bed treatment was achieved with 60 minutes of conventional furnace treatment. A similar trend was shown by the number fraction of silicon particles with size greater than 1 μm illustrating the faster coarsening effect of fluidized bed heat treatment. The analysis of samples that underwent different complete ageing cycles revealed that the ageing of fluidized bed

41 solution treated samples produced spherical Mg2Si precipitates smaller (20-100 nm size

range) than that of conventional furnace solution treated samples (1000-2000 nm size

range). The most important conclusion derived from this work is the issue of thermal

stresses induced by the higher heating rate. The applicability of this theory to the rapid

infrared thermal processing of aluminum forgings will be discussed when the results of

the laboratory simulation of rapid solution treatment are discussed.

2.3 Solution Heat Treatment

Andreatta, Terryn, and de Wit (2003) studied the effect of solution heat treatment

on galvanic coupling between the intermetallics and matrix in AA7075-T6. The galvanic

coupling depends on dissolution of strengthening particles. Therefore even though the

focus of the investigators was on galvanic coupling, it is appropriate to review it as a

current research project related to the investigation of solution treatment of aluminum

forging with rapid infrared heating. Cold-rolled AA7075-T6 was used for the study. A

salt bath was used for the solution heat treatment of the samples. The solution treatment

temperature was 470 ºC. The solution treatment durations ranged from 60 minutes to 4

hours. The solutionized samples were quenched immediately. The electrical

conductivity and microhardness measurements were performed on the solution treated

samples to analyze the effect of the thermal processing. Volta potential maps were

obtained to study the Volta potential differences between intermetallics and matrices. An

SEM with EDAX facility was used to study the morphology and the composition of the

intermetallics obtained after the thermal treatment. Three types of intermetallics were

identified. Based on the microhardness and electrical conductivity measurements of

42 samples with different solutionizing times, it was concluded that 15 minutes of solution heat treatment at 470 ºC would be enough for the complete dissolution of MgZn2 strengthening particles. While the identification of the intermetallics on the reviewed work might not be specifically relevant to this work, the experimental methodology of using microhardness measurements and electrical conductivity measurements to study the effect of solution treatment of cold worked aluminum provides some guidelines for the physical simulation of thermal processing of aluminum with rapid heating for the current research work.

Davidson, Griffith, and Machin (2002) conducted a study to check whether the reduction in solution heat treatment time would affect fatigue properties. Strontium modified Al-7Si-0.6Mg (A356) alloy was used for the study. Samples from three different sources were used. The conventional heat treatment schedule consisted of 8 hours of solutionizing at 540º C followed by a hot-water quench. The solutionizing with reduced time was performed at 540 ºC for 4 hours. Quenching was performed in a water bath at 50 ºC and it was followed by aging at 162 ºC for 4 hours. The fatigue test samples were obtained from the samples and a final surface finish of ~1μm was obtained in the gauge sections of the fatigue samples. The fatigue life test was designed to replicate the service condition of the cast components. The loading cycles were of zero stress ratio with a frequency of 60Hz. To check the dependency of fatigue behavior on the loading cycle frequency, some of the samples were tested for loading cycles with a

1Hz frequency. The loading cycles were designed such that the test samples would fail within 105 number of cycles. In addition, tensile tests were conducted and the results

showed no significant change as a result of the reduction of solution heat treatment time.

The fatigue results were analyzed with statistical techniques. The statistical analysis did not identify any significant bias due to the reduction of solution heat treatment time at

95% confidence level.

Chen, et al. (2003) studied the changes in constituent dissolution and the resulting mechanical properties resulting from the stepped heat treatment in AA 7055. First, a single step homogenization and two step homogenization were carried out on cast ingots and their effects on the initial melting temperature of the ingots were studied. Another set of extruded bars was subjected to isothermal solution heat treatment or stepped solution heat treatment and quenched in ambient water followed by ageing at 120 ºC for

24 hours. Tensile properties, microhardness and fracture toughness were measured from samples obtained from the heat treated extruded bars. The homogenized ingots and heat treated extruded bars were studied at the microstructure level. In addition, as-cast microstructure was also analyzed. The fractrographs were obtained using an SEM. The analysis of the microstructure of the as-cast ingot and the heat treated ingot showed that a first step preheating to a temperature below the melting point of the as-cast alloy increased the initial melting temperature of the as-cast AA 7055. Similar behavior was observed with extruded bars. The increase in isothermal solution heat treatment made it possible to have a higher temperature solution treatment without the risk of overheating.

This resulted in improvement in mechanical properties of the alloy. SEM analyses of the tensile fracture surfaces indicated that the two- step treatment created less coarse constituent particles which enhanced the mechanical properties. While similar phenomena can be expected to occur in the thermal processing of aluminum alloy

44 forgings, due to limitations in the availability of the resources, the work presented herein did not look into this aspect.

Suni, Rouns and Shuey (2000) investigated high heating rate during annealing and its influence in the resulting grain size in cold rolled aluminum alloys. The samples used in this study were made with different solidification methods. Different alloy samples with different percentages of cold work reduction were used in the study. The faster heating was performed in a lead bath and the slower heating was performed in an air furnace. After annealing, the samples were quenched. The effect of these thermal processing was measured with electrical resistivity measurements. The grain sizes and related geometrical parameters were measured either from optical microscopy or orientation imaging microscopy. The results showed that for some of the samples, high heating rates slightly increased the recrystallized grain size while decreasing the aspect ratio. For other samples, faster heating rate produced finer equiaxied grains than slower heating rates. Considering this with the results of dispersoid measurement in the samples, it was concluded that the presence of fine particles would be required for the heating rate to have some effect as the dispersoids could have their ‘pinning’ effects.

This could be the reason for grain refinement with rapid infrared thermal processing.

The effect of resistance heating on recrystallization behavior of Al-7Si-0.6Mg alloy was investigated by Ferry and Johns (1998) with an electrical resistance heating furnace. By thermo mechanical processing, samples with similar grain sizes and textures were obtained first from a single phase alloy and a particle containing alloy. The thicknesses of these samples were reduced 70% by cold rolling. The rolled samples were then annealed with heating rates up to 1000 ºC/s. For comparison, some samples were

45 processed with same annealing process but at a heating rate of 50 ºC/hour. For this purpose, a conventional batch annealing furnace was used. X-ray diffraction, hardness measurements and optical microscopy were employed for the analysis of the effect of annealing. The results showed that higher heating rates reduced the incubation period for nucleation in the alloy tested.

2.4 DSC Studies on Thermal Processing of Aluminum Alloys

Previous work on the application of differential scanning calorimetry (DSC) for the thermal processing of aluminum alloys is reviewed in this section. The purpose of this review is to show how DSC techniques have been used for the analysis of thermal processing of aluminum alloys. In addition, the findings of these works are also discussed in this section.

DSC can be used to measure the heat energy required to increase the temperature of a sample in a required manner at each point of the thermal history. DSC work of this research is treated in detail in Chapter 4.

DSC is used to measure and calculate the net heat energy used to heat the sample through small temperature increase. The net heat does not include heat losses. This net heat energy includes the specific heat necessary for the temperature increase and the energy used for any solid state reaction if there was one. If the specific heat required for the temperature increase is subtracted from the net heat measured, the resulting heat energy can be related to energy produced or absorbed due to a solid state reaction at that temperature range. The specific heat capacities of pure alloys at any temperature range can be obtained from the open literature. However, finding the specific heat capacity of

46 an alloy at a range of temperature is tricky due any solid state reaction that might occur.

However, it has been shown that the specific heat capacity of an alloy can be found by using the Neumann- Kopp rule (Illeková, Gachon and Kuntz 2002). This calculation is treated in detail in chapter 4. In other words, the variation of the difference in specific heat capacity at different temperatures of a heating cycle indicates a solid state reaction occurring at that temperature (Hirano 1974).

This technique has been used to study the solid state reactions of different aluminum alloys during different thermal cycles for various research works (Hirano

1974) (Rooyen, Maartensdijk and Mittemeijer 1988) (Zhen, et al. 1997). Hirano (1974) used DSC thermal analysis to study the precipitation process of aluminum-rich Al-Zn-Mg alloys. Even though, his work covers DSC studies of the ageing process of a previously supersaturated alloy, and this work focuses on the solution treatment cycle, this work is reviewed here to get a basic understanding of using DSC study for the analysis of thermal processing of aluminum alloys. In their study, the heat absorption calculated by DSC measurement during the heating of the aluminum alloy was attributed to the dissolution of second phase particles in the matrix. The heat evolution observed with DSC during the aging of solutionized and quenched aluminum alloy was attributed to the precipitation of second phase particles from the supersaturated matrix.

Rooyen, Maartensdijk and Mittemeijer (1988) used DSC analysis for the study of precipitation of Guinier-Preston (GP) zones in aluminum-magnesium alloys. Aluminum alloys with various weight percentages magnesium content were used in the study. These samples were annealed with heating cycles with constant heating rate and DSC analysis was performed on the cycles to estimate the heat of dissolution of the GP-zones and the

47 enthalpy of formation of GP-zones. Even though this work is not directly relevant to the current research work, it provided guidelines on analyzing the results of DSC measurements.

Unlike the above-mentioned two studies, where DSC study was applied on aged samples and the net heat absorption or evolution was related to the formation or dissolution of GP zones, there are many studies using DSC analysis of solutionized and quenched samples. These studies look into the thermograms obtained when already thermally processed aluminum alloy samples were being heated. The peaks in the thermograms were then related to dissolution of different precipitates. Even though this approach will not be used in the current work, these works are discussed to show the versatility of the DSC analysis.

Zhen, et al. (1997) investigated the precipitation behavior of Al-Mg-Si alloys with high silicon content. DSC study and transmission electron microscopy (TEM) were used for the study. First, baseline thermal energy data for a high purity aluminum sample for a similar thermal cycle was obtained. By subtracting this baseline data from the net thermal energy transfer, the thermal energy associated with solid state transformation during the thermal cycle was estimated. The resulting thermogram showed 6 peaks.

TEM was used to identify the precipitates related to these peaks. Individual samples were heated to each of the identified peak temperatures and held at those temperatures for

3 minutes before rapid quenching. These specimens were studied with TEM. Using this procedure, they studied the effect of the amount of cold work and the effect of the solution heat treatment time on the precipitation. Even though TEM is not used in the

48 current research, this paper indicated that DSC could be used to investigate the solid state transformation of aluminum alloys.

DSC and TEM techniques were also used by DeIasi and Adler. (DeIasi and Adler

1977) (Adler and Deiasi 1977). In the first study, these techniques were used to study the precipitation characterization of AA 7075. Highest strength (T651) temper and overaged

(T7351) temper AA 7075 were used in the study. Quantitative analysis was performed to understand the solid state reactions occurring when these temper alloys were heated to

500 ºC.

In the second paper, the authors applied the same technique to compare the precipitates in AA 7000 series alloys (Adler and Deiasi 1977). Their DSC results matched excellently with their TEM results.

Precipitation in AA 2219 was studied with DSC and TEM techniques by Papazian

(Papazian 1982) (Papazian 1981). The variation of differential heat capacity of individual AA 2219 samples after specific ageing cycles with increasing temperature was measured with DSC. The thermograms showed distinct characteristics for some of the ageing treatments. In another study, Papazian (1982) used DSC analysis to investigate the precipitation and dissolution kinetics in AA 2219 and AA 7075. In this study, different heating rates were used for heating the samples for dissolution and resulting thermograms were studied.

The next chapter looks into savings in energy and processing time associated with rapid heating method.

CHAPTER 3: SAVINGS IN ENERGY AND PROCESSING TIME

3.0 Introduction

The savings in energy and time as a result of infrared thermal processing are briefly presented in this chapter.

3.1 Modeling of Heat Transfer

In order to explore the savings in energy and time resulting from using infrared heating instead of conventional gas furnace heating, the heat transfer process for both these types of heating is modeled in this section.

In a conventional gas-fired furnace, heat transfer to the forged parts occurs mostly due to convective heat transfer and the medium is air. This heat in the hot air has to

transfer to the parts. Based on heat transfer theory, the heat transfer rate (Q& c ) due to

convection can be expressed as,

Q& c = hAΔT (1)

where h is the convective heat transfer coefficient, A is the surface area for heat transfer and ΔT is the temperature difference. For forced air, the convective heat transfer coefficient ranges between 10 and 100 W/m2K (Convective Heat Transfer 2005). The

flame temperature of propane flame with air as an oxidizer is around 2300 K (Convective

Heat Transfer 2005). Hence the convective heat transfer rate in a gas-fired furnace is

between 23,000 -230,000 W/m2s (Convective Heat Transfer 2005).

In an infrared heating furnace, heat transfer occurs by convection as well as by

radiation. As there are two types heat transfer involved, infrared heating is sometimes

referred as a hybrid heating method. The convective heat transfer portion of infrared

50 heating is similar to that of a conventional gas-fired furnace. There is however a difference in the source temperatures of a gas-fired furnace and an infrared furnace.

The radiation heat transfer from the infrared furnace to the load can be modeled

using the Stefan-Boltzmann law. The radiation heat transfer rate (Q& r ) can be expressed

as,

4 4 Q& r = σε(T f −Ts ) (2)

where σ is the Stefan-Boltzmann constant, ε is emissivity, Tf is the temperature of the filament of the infrared heating source and Ts is the surrounding temperature. The value

−8 -2 -4 of the Stefan-Bolzmann constant is 5.670400×10 W·m ·K . The filament temperature

of the infrared heat source is above 2400 K. This would result in a radiation heat transfer

rate of 1,600,000 W/m2s for an emissivity value of 0.9 and some infrared heaters can have an emissivity up to 0.9 (Electric heating Association Incoporated n.d.).

These formulae show that the heat transfer by convection is between 23 and 230

kW/m2s, while the heat transfer by radiation can be up to 1600 kW/m2s. It should be

noted that the infrared heating method uses both convective heat transfer and radiation heat transfer.

CHAPTER 4: DSC EXPERIMENTATION AND ANALYSIS

4.0 Introduction

Differential scanning calorimetry (DSC) studies on the effect of heating rate on solutionizing of aluminum alloys are presented in this chapter. First a brief introduction is given on DSC studies. The experimental procedure is then presented. This is followed by the results of the DSC studies. The thermal processing of aluminum forgings with rapid heating is explained in light of the results of the DSC experiments.

4.1 Differential Scanning Calorimeter

Differential scanning calorimetry measures the temperature change in a sample during a time interval and the net heat input to the sample during that time interval. The differential scanning calorimeter (DSC) can be programmed to heat the sample at a specified rate provided the heating rate is within the range possible with the DSC.

During the thermal cycle, the DSC records the temperature history of the sample and the heat input to the sample at each interval of time as selected. First a blank

(container) is heated in such a way that its temperature history would be similar to that of the sample to be analyzed. All the heat input during this heating cycle is equal to the heat used for heating the blank and environmental heat loss. When the sample is heated, it is enclosed in a blank with the same dimensions and of same material. By subtracting the heat input during the heating of the blank from heat input during the heating of the sample, the net heat energy used for the heating of the sample at each interval can be found.

The specific heat capacity of the sample can be found at different temperatures from these measurements and the data from heating a reference material in similar fashion.

4.2 Experimentation

A Mettler-Toledo differential scanning calorimeter was used in this work. This instrument is a DSC822e model. Liquid nitrogen was used as the coolant in the system.

A photograph of the DSC system used in this work is shown in Figure 4.

Figure 4: DSC with the liquid Nitrogen tank.

Since the aim of the investigation was to study the effect of heating rate on the

solutionizing of aluminum alloys, two heating rates, 2 ºC/min and 20 ºC/min were

selected for the experiments. However, at a heating rate of 2 ºC/min the thermal cycles

with the reference material (sapphire) did not give a consistent result. Hence it was

decided to use thermal cycles with heating rates 5 ºC/min and 20 ºC/min. The intention

53 of the study was to measure the heat transfer to the sample when heating at these rates.

Specific heat capacity (cp) values and the variation of specific heat capacity values at

different temperatures while heated at these two heating rates can be calculated from the

measured heat transfer data. Solid state reactions, such as precipitation, etc., can be

recognized from the peaks in the variation of specific heat capacity (Δcp).

To eliminate the effect of the variation of room temperature, each heating cycle was started with initially keeping the sample at 20 ºC for 20 minutes. The heating cycles started at 20 ºC and increased to 532 ºC for both alloys since the conventional solutionizing temperature of AA 6061 is 532 ºC and the conventional solutionizing temperature of AA 2618 is 530 ºC. In order to study the effect of heating rate on 2nd- phase dissolution at solutionizing temperature, the cycles included a dwell period at 532

ºC for 30 minutes. The heat transfer during this period was measured and this would be an indicator of any solid state reactions taking place during this time.

The samples used for the DSC studies were obtained from extruded bar stock of

AA 2618 and AA 6061 alloys. The samples were cylinders about 5 mm in diameter and

1.5 mm in height. The samples were placed inside aluminum blanks with internal diameters of 5 mm and 2mm in height. To facilitate proper heat transfer between the aluminum sample and the blank, the round faces of the cylinders were polished to get flat surfaces. Once the sample was placed inside the blank, it was sealed by placing the cap on it and pressing it with a small press. The mass of the sample was measured before running the test.

First an experimental thermal cycle was applied to a standard aluminum blank and relevant thermal data were recorded. The data included time, temperature of the blank

54 and net rate of heat transferred to the blank during each time interval. The time interval was set to 1 second. The data from the run (with blank only) was used to account for the heat losses to the environment and the heat used for the heating of the blank.

Secondly, a sapphire sample was heated with the same thermal cycle. Sapphire is commonly used as a standard material in the calculation of specific heat capacity (cp) of the material being analyzed using a DSC. The calculation of cp is based on Equation 1

(Hirano 1974).

(HeatFlow) c = kc (1) p p (HeatingRate) where kcp is specific heat of the reference material

Since, the thermal processing is related to the dissolution of intermetallics phases,

which depend on the alloying elements and their amount, the chemical composition of the

alloys and possible second phases in the alloys are presented in the next section.

4.3 Chemical Composition of the Alloy Samples

The extruded alloy samples were obtained from Kaiser Aluminum, Newark, Ohio,

and they were processed in the following manner. Ingot 381 mm (15 inches) in diameter

were direct-chill (DC) cast with the required chemical composition. Using a 58.7 MN

(6600 US ton) indirect press with a 5-hole die, the DC cast billets were extruded to 57

mm (2¼ inches) diameter rods. These rods were from the same batch of extruded rods

used for the production of prototype forgings discussed in Chapter 6. The chemical

compositions of AA 2618 and AA 6061 rods are presented in Table 4.1 and Table 4.2 as

determined by atomic emission spectroscopy. These are compared with the compositions

specified by the Aluminum Association (AA) for corresponding alloys (Cayless 1993).

Table 4.1: Composition of AA2618 test material in weight percentage, via atomic emission spectroscopy. The balance is aluminum. Elements by weight percent Alloy Cu Fe Si Mn Mg Zn Cr Ni Ti

AA 2618 1.97 1.07 0.24 0.02 1.52 0.01 ----- 1.17 0.07

AA 1.9- 0.9- 0.10- 0.05 1.3- 0.10 0.05 0.9- 0.04-

specification 2.7 1.3 0.25 max 1.8 max max 1.2 0.10

Table 4.2: Composition of AA6061 test material in weight percentage, via atomic emission spectroscopy. The balance is aluminum. Elements by weight percent Alloy Cu Fe Si Mn Mg Zn Cr Ni Ti

AA 6061 0.16 0.28 0.63 0.09 0.82 0.06 0.10 ----- 0.02

AA 0.15- 0.7 0.4- 0.15 0.8- 0.25 0.04- 0.15 ----- Specification 0.4 max 0.8 max 1.2 max 0.35 max

The solid state reactions during the thermal processing of the aluminum alloys

depend on the second phases (or intermetallics) present or created. Hence, second phases

that might be present in the two selected alloys are listed in Table 4.3 (Bray 2002)

(Hatch 1984). These include intermetallics that are soluble with temperature and

insoluble intermetallics.

Table 4.3: Conventional solution temperatures and precipitates present in the alloys under investigation. Solution

Alloy Temperature Primary Precipitates

(ºC)

2618-T61 530 CuAl2, CuMgAl2

(a) 6061-T6 532 Mg2Si, Mg5Si6

(a) Even though the primary precipitate in AA6061-T6 has been thought to be Mg2Si, recent works indicate that this is an equilibrium, non-hardening (incoherent) precipitate and the primary hardening precipitate is likely to be Mg5Si6 (Pangborn, et al. 2008).

4.4 Results and Analysis

As mentioned previously, the net energy transfer to the aluminum coupon at each of the time intervals in the heating cycles were measured. First, the effects of heating rates on the solutionizing of AA 2618 and AA 6061 were studied from these net energy measurements. Then the Δcp values (heat capacities) of the alloys at different temperatures were calculated and they were compared with the heating cycles of slower and rapid heating rates. Though the heating rates employed were 5 ºC /min and 20 ºC

/min and these two rates do not accurately represent the actual heating rates during conventional and rapid heating of aluminum forgings, it is reasonable to expect the results and analysis of the DSC measurements to be similar to that of conventional and rapid heat treating of aluminum forgings.

4.4.1 Net Energy Transfer at Different Temperatures

One AA 2618 coupon was kept at 20 ºC for 5 minutes and then heated to 532 ºC at a heating rate of 5 ºC/minute. Once the temperature reached 532 ºC, the coupon was kept at 532 ºC for 30 minutes. The DSC measured the net energy transfer from the coupon being studied. The net energy measurements were divided by the mass of the coupon to determine the net energy transfer per unit mass of AA 2618. The variation of the net heat transfer is shown in Figure 5. Also shown in this figure is the temperature of the coupon at the time of measurement.

Only one coupon was subjected to each thermal cycle for DSC analysis. There might be a spread in the results, had there been multiple coupons were used for DSC analysis. However, since our aim was to look into the qualitative effect due to variation in heating rates, results with one coupon for each cycle could serve the current purpose.

0 2000 4000 6000 8000

40 Heat transfer of AA 2618 at 5 °C/min 600 Temperature of the AA 2618 sample

20 500 C)

400 ° 0

300 -20

200 -40 Temperature (

Heat Transfer (mW/kg) Transfer Heat 100 -60 0 -80 0 2000 4000 6000 8000 Time (s)

Figure 5: Heat flow rate for unit mass of AA 2618 during the heating cycle with a heating rate of 5 ºC/min. The corresponding temperature history is also shown.

In the similar fashion, a DSC study was conducted on another AA 2618 coupon with the higher heating rate (20 ºC/ minute). The resulting net heat transfer and temperature increase are shown in Figure 6.

-500 0 500 1000 1500 2000 2500 3000 3500 700 100 Heat transfer of AA 2618 at 20 °C/min Temperature of AA 2618 600

0 500 C) ° -100 400

-200 300

200 -300 ( Temperature

Heat Transfer (mW/kg) 100 -400 0 -500 -500 0 500 1000 1500 2000 2500 3000 3500 Time (s)

Figure 6: Heat flow rate for unit mass of AA 2618 during the heating cycle with a heating rate of 20 ºC/min. The corresponding temperature history is also shown.

When both figures are compared, the following points can be noted. As expected,

the net heat transfer to the coupon at a higher heating rate was greater than that at lower

heating rate. This is logical since the higher heating rate requires a greater heat transfer

rate than the lower heating rate. However, it was decided to keep the AA 2618 coupon at

532 ºC for 30 minutes to check whether the DSC measurements indicate any solid state

reaction occurring at that temperature. Figure 5 shows that when the AA 2618 coupon

reached 532 ºC at a heating rate of 5 ºC/minute, the coupon kept losing heat at the constant temperature of 532 ºC. One or both of the following might be the reason for

this: Dissolution of second phases and grain coarsening. As the current work did not

involve measurement dissolution of second-phase particles in aluminum alloys with slow

heating method and measurement of grain size with solutionizing duration at slower

60 heating rate, these two factors could not be directly related to the above-mentioned heat loss at 532 ºC. On the other hand, when another AA 2618 coupon was heated to 532 ºC at a heating rate of 20 ºC/minute and held at 532 ºC, the net heat transfer to the coupon was nearly zero. This probably indicates that no solid state reaction was taking place as dissolution of soluble intermetallics phases had almost completed during the rapid heating. However, this will be discussed in Chapter 9.

Similar DSC studies were conducted with AA 6061 coupons at the two different heating rates. The net heat transfer to the AA 6061 coupon and corresponding temperature variations are shown in Figure 7 and Figure 8.

0 2000 4000 6000 8000

o 700 40 Heat transfer of AA 6061 at 5 C/min Temperature of AA 6061 600

20 C) 500 °

0 400

-20 300

200 -40 Temperature (

100 Heat Transfer (mW/kg) -60 0 -80 0 2000 4000 6000 8000 Time (s)

Figure 7: Heat flow rate for unit mass of AA 6061 during the heating cycle with a heating rate of 5 ºC/min. The corresponding temperature history is also shown.

-500 0 500 1000 1500 2000 2500 3000 3500 600 0 500 -100 C)

400 ° -200 300 -300 200 -400 100 ( Temperature o -500 Heat transfer of AA 6061 at 20 C/min Heat Transfer (mW/kg) Transfer Heat Temperature of AA6061 0 -600 -500 0 500 1000 1500 2000 2500 3000 3500 Time (s)

Figure 8: Heat flow rate for unit mass of AA 6061 during the heating cycle with a heating rate of 20 ºC/min. The corresponding temperature history is also shown.

As in the case with the results of the AA 2618, higher heating rates indicated heat transfer at higher rates. Secondly and more interestingly, when the AA 6061 coupon was held at 532 ºC after heating at 5 ºC/ minute, the coupon was releasing heat at a rate of 20 mW/kg while the AA 6061 coupon heated at rate of 20 ºC/minute did not release or absorb heat when held at 532 ºC. This also will be discussed in Chapter 9.

The direct comparisons of net heat transfer to the aluminum coupons during thermal cycles with different heating rates were reviewed above. This is followed by comparisons of changes in specific heat capacities during the thermal cycles with different heating rates.

4.4.2 Variation of Specific Heat Capacities

The specific heat capacities were calculated by the software using Equation 1.

The variation of the specific heat capacities with temperature increase was due to two factors. The first one is the inherent change of specific heat capacity due to temperature increase. The second factor was the solid state reaction of the alloys. The inherent changes of specific heat capacities of the selected alloys due to temperature increase are not available in the open literature. However, the inherent changes of specific heat capacities of pure elements constituting the alloys due to temperature increase were found from the literature. Using these data and the Neumann–Kopp rule, the inherent changes of specific heat capacity of the alloy due to temperature increase were calculated at temperatures between 20 ºC and 532 ºC (Illeková, Gachon and Kuntz 2002). By subtracting the calculated inherent change of specific heat capacity due to temperature increase from that obtained from the DSC study, the change in specific heat capacity due to solutionizing was obtained. Figure 9 shows two such plots of change in specific heat capacity of AA2618 during the two thermal cycles starting from 20 ºC to 532 ºC.

0.5 o Variation of Δc for AA 2618 at 5 C/min p o 0.4 Variation of Δc fop AA 2618 at 20 C/min p 0.3

0.2

0.1 p c

Δ 0.0

-0.1

-0.2

-0.3

-0.4

-0.5 0 100 200 300 400 500 Temperature (°C)

Figure 9: Variations of change of cp of AA 2618 at two heating rates during thermal cycles with the same maximum temperature.

The area under the line showing the variation of Δcp is equal to the specific energy of the solutionizing reaction corresponding to the temperature.

Similar plots for DSC studies with AA 6061 coupons are shown in Figure 10.

0.4

0 Variation of Δc for AA 6061 at 5 C/min p 0.3 o Variation of Δc for AA 6061 at 20 C/min p

0.2

0.1 p c

Δ 0.0

-0.1

-0.2

-0.3 0 100 200 300 400 500 Temperature (°C)

Figure 10: Variations of change of Cp of AA 6061 at two heating rates during thermal cycles with the same maximum temperature.

4.5 Conclusions

The region of thermal history being analyzed using DSC corresponds to the

heating of the sample for solution treatment. It has been established that the solid state

reaction occurring in this region is the dissolution of the second phases (ASM Handbook

2002).

Based on the results of the DSC analyses with AA 2618 and AA 6061, the following can be concluded.

1. As expected, high heating rate required higher flow of heat to the sample. In the

thermal cycle with higher heating rate (20 ºC/min), the net heat flow rate after

reaching 532 ºC was almost zero. Faster heating might have accelerated the

diffusion process and hence once reaching 532 ºC, there was not much significant

diffusion to be completed. On the other hand the net heat flow rate at 532 ºC

during the thermal cycle with the slower heating rate is significant. This might

indicate the occurrence of a solid state reaction in the sample at 532 ºC during the

thermal cycle with the slow heating rate. This might indicate that solid state

reaction is continuing after the temperature reached 532 ºC with slow heating.

This will be discussed in Chapter 9 in light of the results from other experiments

in this research work.

2. The area under the Δcp vs. temperature curve corresponds to the energy used for

any solid state reaction. Comparisons of the time vs. Δcp curves for thermal

cycles with different heating rates seem to show that at a higher heating rate, more

net energy is used for the solid state reaction. The only solid state reaction that

might occur at this stage of a thermal cycle is the dissolution of the second phases.

Hence, it seems that at a higher heating rate, more dissolution occurs in AA 2618

and AA 6061.

Physical simulation of the solution treatment of AA 2618 and AA 6061 was conducted in a laboratory furnace to determine the “optimum” solution treatment cycle with rapid heating. This investigation is presented in the next chapter.

CHAPTER 5: PHYSICAL SIMULATION OF SOLUTION TREATMENT WITH

RAPID HEATING

5.0 Introduction

Previous studies have shown that solution treatment with rapid heating can reduce both energy consumption and lead times while improving mechanical properties, or at least producing parts that are on par with those produced with conventional heating methods (Lu, et al. 2004) (Blue, et al. n.d.). However optimum solutionizing parameters with rapid heating methods have thus far not been established for AA 2618 and AA 6061 forgings, in essence, wrought aluminum alloys.

Physical simulation of solution treatment with rapid heating was performed using a laboratory furnace. Rapid solution cycles with different solution treatments were employed and the resulting degrees of solutionizing were measured. Based on the results of this study and the practical limitations of rapid infrared solution treatment at the shop floor, the optimum solution treatment cycles were determined for AA 2618 and AA 6061 forgings.

5.1 Materials

The aluminum alloy materials used for the physical simulation of the solution treatment were of the same batch used for the DSC study. The elemental composition and information on the production process of the alloys was presented in Chapter 4.

5.2 Experimental Procedure

An infrared furnace can heat aluminum alloy forgings at a very rapid rate. It is reported that the heating rate on medium sized parts is about 20 ºC/minute (Kervick, et al.

2006). In order to obtain the rapid heating rate using a laboratory furnace, small aluminum alloy coupons were used for the trial solutionizing experiments. The idea was to individually apply several different rapid solution treatment cycles to different aluminum alloy coupons and to evaluate the resulting degree of solutionizing. Since the round parts from which the coupons were made were extruded and in wrought form, it was reasonable to expect that their solutionizing behavior would be similar to that of the forgings.

Rods of 19 mm diameter were machined from the extruded bars by wire electro discharge machining (EDM). These 19 mm diameter rods were sliced into disks of

4.8 mm thickness. These disks were the coupons used in the solution treatments. Figure

11 shows an extruded bar and the coupons obtained from it.

Figure 11: Extruded bars and coupons obtained from them for solutionizing trails. The coupons were 19mm in diameter and 4.8mm thick.

The degree of solutionization was measured by the increase in microhardness and

change in electrical conductivity. To facilitate these measurements, one surface of each

coupon was mechanically polished with 6μm diamond media before the solution

treatment. To accurately measure and control the temperature of the coupon, it was

decided to measure the temperature directly from the coupon itself. This was accomplished by fine (30 AWG) K-type thermocouple wire welded to the unpolished side of the coupon.

The prepared coupon was individually placed in a small natural-convection lab furnace (Thermalyne 1300). A data acquisition module (DAQ) attached to a computer recorded the temperature of the coupon when it was being solutionized and quenched.

Once the temperature of the coupon was close to the pre-decided solution temperature, the temperature of the furnace was manually adjusted to keep the coupon at the desired

69 solution temperature. After the temperature of the sample reached the desired solutionizing temperature, the coupon was held at that temperature for the desired amount of time and immediately quenched in water. The temperatures of the quenching water were 65 ºC and 100 ºC for AA 6061 and AA 2618 coupons respectively. As an example, the temperature cycle of a coupon during a rapid solution treatment at 532 ºC for 1200 seconds is shown in Figure 12.

520 480 440 400

C) 360 ° 320 280

240 200 Temperature( 160 120 80 40 0 0 200 400 600 800 1000 1200 1400 1600 1800 2000 2200 2400 Time (s)

The time delay between the removal of the sample from the furnace and the time

it was put into the quenching media is very important. The Aerospace Metals

Specifications (AMS) emphasizes this and specifies 15 seconds as the maximum time

delay (AMS 2770E Heat Treatment of Wrought Aluminum Alloy Parts 1989). In all the

solutionizing cycles performed, this time delay was kept well below this maximum

70 duration. The typical temperature variation during time of quenching is shown in Figure

13.

Time (sec) 936.0 937.2 938.4 939.6 940.8 942.0 943.2 560 520 480 440 Enter quench Commence specimen 400 removal from furnace

C) 360 ° 320 280

240 200 Temperature ( 160 120 80 40

15.60 15.62 15.64 15.66 15.68 15.70 15.72 Time (min)

Figure 13: Typical quenching cycle for samples heated to 532 ºC (Gowreesan, et al. 2006).

The temperature loss from the removal of the coupon from the furnace to entering

the quenching media is about 10-15 ºC. The coupon was kept in the quenching media for about 5 minutes. Then the coupon was dried and the electrical conductivity and hardness

were measured on the previously polished side of the coupon. The electrical conductivity

and hardness measurements are effective methods of measuring the degree of solution

treatment (Andreatta, Terryn and de Wit 2003). A Foerster Sigmatest® 2.068 eddy current tester was used for the electrical conductivity measurements. The instrument had

71 a standard 14 mm probe. The measurements were taken at a frequency of 120 kHz.

Before electrical conductivity was measured on each coupon, the instrument was calibrated using standard calibration disks provided by the manufacturer of the equipment. The coupon was immediately placed in a freezer at -18ºC to retard the natural ageing process.

Once all the individual coupons were solution treated, quenched and frozen, they were precipitation hardened according to the specified ageing process relevant to the alloy. The ageing cycle used for all AA 6061coupons was ageing at 175 ºC for 8 hours, and this cycle creates the T6 condition for AA 6061. The solutionized AA 2618 coupons were aged to the T61 condition by ageing at 199 ºC for 20 hours. Electrical conductivity and hardness were measured after the ageing process also.

Metallurgical analyses were performed on selected solutionized coupons. The coupons for metallurgical analysis were selected based on the results of the electrical conductivity and hardness measurements. For example, these measurements did not show much variation above solutionizing times of 900 seconds. Hence, coupons solutionized for 20 seconds and coupons solutionized for 900 seconds were selected for metallography. In addition, as-extruded coupons were also metallographically analyzed.

Pieces from the selected coupons were mounted in epoxy-resin mounting media.

After the mounting media was cured, one surface of the sample was mechanically polished and electrochemically etched to reveal the grain structures. A solution of 5ml fluoroboric acid (HBF4) and 200 ml distilled water was used as the media for etching.

The surface of the sample to be etched was immersed in the etching solution and it acted

as the anode while a strip of stainless steel immersed in the etchant was used as the

72 cathode. The electrodes were connected to a 22V DC power source. A magnetic stirrer was used to stir the etchant continuously to remove the build-up of bubbles on the electrodes during the etching. An etching time of about 3 ¾ minutes was found to be the right etching time to reveal the grain boundaries in the aluminum alloy coupons. The etched coupons were rinsed with water and immediately dried with compressed air.

Since the coupons were from extruded bars, it was decided to perform metallographic analysis on planes perpendicular to the direction of extrusion and on planes parallel to the direction of extrusion. A Nikon optical microscope was used to view the grain structure at higher magnifications. Grain size measurements were performed on the digital pictures of the grain structures using ASTM line intercept method ( Standard Test Methods for Determining Average Grain Size 2004).

The results of the physical simulation of solution treatment of AA 2618 and AA

6061 with rapid heating are as follows. The results include electrical conductivity measurements, hardness measurements and optical microscopy of the aluminum alloy coupons under the various solution treatments.

5.3 Results

Results of testing with AA 6061 coupons are presented first. This is followed by the results of testing with AA 2618 coupons.

5.3.1 Results of Solutionizing Tests of AA 6061

As previously mentioned, first the coupons individually underwent different solution cycles with rapid heating and the resulting electrical conductivity and hardness

73 were measured before the ageing process. These measurements are presented in Figure

14 and Figure 15 with respect to their corresponding solution parameters.

25.6

25.4

25.2

o 25.0 532 C exponential fit 24.8

24.6

24.4

o 24.2 552 C exponential fit

Conductivity (MS/m) 24.0

23.8 572 oC 23.6 exponential fit

0 500 1000 1500 2000 2500 3000 3500 28800 Solutionizing time (s)

Figure 14: Average electrical conductivity of AA6061 as a function of solutionizing time and solution test temperature (Gowreesan, et al. 2006). Error bars represent ±1 standard deviation. Each data point is an average of 5 measurements on the sample. This graph only includes measurements on “solutionized” samples. The conductivity of as-extruded/un-solutionized material was 29.8 MS/m.

o 60 572 C exponential fit 58

56 o 552 C exponential fit 54

1kgf HV 50 532 oC exponential fit 48

44 as-extruded/un-solutionized

42 0 500 1000 1500 2000 2500 3000 3500 28800 Solutionizing time (s)

The solutionized AA 6061 coupons were artificially aged to T6 condition and the electrical conductivity and hardness values were measured again. The results of these measurements are presented in Figure 16 and Figure 17.

26.4 532 oC 26.2 552 oC o 26.0 572 C

25.8

25.6

25.4

25.2

Conductivity (MS/m) Conductivity 25.0

24.8

24.6

0 500 1000 1500 2000 2500 3000 3500 28800 Solutionizing time (s)

130 128 126 124 122 120 118 1kgf 116

HV 114 112 110 532 oC 108 552 oC 106 572 oC 104

0 500 1000 1500 2000 2500 3000 3500 28800 Solutionizing time (s)

The results indicate that an increase in solutionizing temperature results in increase in hardness and decrease in electrical conductivity in the solutionized/quenched condition and aged T6 condition. The change in solutionizing temperature results in an average 6% variation in electrical conductivity. Similarly a change in solutionizing temperature results in an average increase of 6 HV in hardness. Comparison of these properties resulting from rapid heating solution treatment with those of conventionally processed samples will be discussed in Chapter 8 in detail. These property changes are usually attributed to increase in coherent precipitates in the aged condition.

5.3.2 Results of Solutionizing Tests of AA 2618

Similar experimentation was conducted with AA 2618 coupons and the results are presented in Figure 18 and Figure 19. The effect of solutionizing parameters on electrical conductivity and hardness values are shown in these figures. The measurements taken before and after the ageing of the solutionized AA 2618 coupons are shown in both figures. Only one coupon was treated for 3600 seconds and the corresponding solutionizing temperature is 530º C.

19.3

19.2 530 oC Aged Condition 535 oC Aged Condition 19.1 540 oC Aged Condition 19.0 545 oC Aged Condition

18.9

18.8

18.7

18.6 Conductivity (MS/m) 18.5

18.4

18.3 0 500 1000 1500 2000 2500 3000 3500 4000 Solutionizing Time (s)

150

140

130

120 530 oC Aged Condition 535 oC Aged Condition

1kgf 110 540 oC Aged Condition

HV o 100 545 C Aged Condition

80 as-extruded/F-temper 70

0 500 1000 1500 2000 2500 3000 3500 28800 Solutionizing Time (s)

The results of solutionizing trials with AA 2618 do not exhibit the same clear trend of increased solutionizing with temperature. The change in solutionizing temperature has shown a very slight change of 2% in electrical conductivity and a 15 HV change in hardness. However, it should be noted that the safer solutionizing temperature ranges are

530 ºC -549 ºC and 532 ºC - 580 ºC for AA 2618 and AA 6061 respectively. In other words, the solution treatment temperature range for AA 6061 is considerably higher than that of AA 2618, and as a result, the results of solution treatment tests of AA 6061 indicated clear dependence of solutionizing on the solutionizing temperature.

On the other hand, the results of AA 2618 solution treatment trials clearly seem to indicate that the solutionizing occurred quickly with rapid heating.

5.4 Metallographic Evolution of AA 6061 with Solution Treatment

The selected solution temperatures for AA 6061 were 532 ºC, 552 ºC and 572 ºC.

The results of electrical conductivity and hardness measurements indicated that most of the solutionizing was completing within 900 seconds. Therefore it was decided to perform metallographical analysis on AA 6061 coupons that underwent solution cycles at these temperatures for 20 seconds and 900 seconds. For comparison purpose, metallographical analyses were also performed on as-extruded AA 6061 coupons. In the following pages, photomicrographs of AA 6061 coupons after solution treatment with different solution treatment parameters are presented.

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 20: Micrograph of as-extruded AA 6061 Figure 21: Micrograph of as-extruded AA 6061 coupon on plane parallel to the extrusion coupon on plane perpendicular to the extrusion direction (Mayer, et al. 2007) direction (Mayer, et al. 2007)

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 22: Micrograph of AA6061 coupon after Figure 23: Micrograph of AA6061 coupon solutionizing at 532 ºC for 20 seconds. Taken in after solutionizing at 532 ºC for 20 seconds. plane parallel to the extrusion direction (Mayer, Taken in plane perpendicular to the extrusion et al. 2007) direction (Mayer, et al. 2007)

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 24: Micrograph of AA6061 coupon after Figure 25: Micrograph of AA6061 coupon solutionizing at 532 ºC for 900 seconds. Taken after solutionizing at 532 ºC for 900 seconds. in plane parallel to the extrusion direction Taken in plane perpendicular to the extrusion (Mayer, et al. 2007) direction (Mayer, et al. 2007)

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 26: Micrograph of AA6061 coupon after Figure 27: Micrograph of AA6061 coupon after solutionizing at 552 ºC for 20 seconds. Taken in solutionizing at 552 ºC for 20 seconds. Taken in plane parallel to the extrusion direction. plane perpendicular to the extrusion direction.

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 28: Micrograph of AA6061 coupon after Figure 29: Micrograph of AA6061 coupon after solutionizing at 552 ºC for 900 seconds. Taken solutionizing at 552 ºC for 900 seconds. Taken in in plane parallel to the extrusion direction. plane perpendicular to the extrusion direction.

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 30: Micrograph of AA6061 coupon after Figure 31: Micrograph of AA6061 coupon after solutionizing at 572 ºC for 20 seconds. Taken in solutionizing at 572 ºC for 20 seconds. Taken in plane parallel to the extrusion direction (Mayer, plane perpendicular to the extrusion direction et al. 2007) (Mayer, et al. 2007)

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 32: Micrograph of AA6061 coupon after Figure 33: Micrograph of AA6061 coupon after solutionizing at 572 ºC for 900 seconds. Taken in solutionizing at 572 ºC for 900 seconds. Taken plane parallel to the extrusion direction (Mayer, in plane perpendicular to the extrusion et al. 2007) direction (Mayer, et al. 2007)

The as-extruded grain structure exhibits the presence of longitudinally distorted or fibrous, elongated grains in the direction of extrusion. The solution treatment does not

recrystallize the elongated grains. Instead grain coarsening seems to have occurred.

Grain size measurements were performed on different photomicrographs. The

variation of grain size with the solution treatment parameters is shown in Figure 34. The

83 grain size measurements were taken from photomicrographs taken on planes perpendicular to the direction of extrusion.

The solutionizing times in the plot corresponds to solutionizing duration in addition to the time taken for the coupon to reach the solution treatment duration. For example, the data point corresponding to 20 seconds at 532 ºC indicates a solution cycle including a heating step up to 532 ºC and therefore it would be different from as-extruded coupon.

0.23

0.22

0.21

0.20

0.19

0.18 Solutionizing at 532°C 0.17 Solutionizing at 552°C Solutionizing at 572°C Grain Size(mm) 0.16 As-extruded sample

0.15 As-extruded sample 0.14 0 200 400 600 800 1000 Solutionizing Time (s)

Figure 34: Variation of grain size with time at different solutionizing temperatures. Grain sizes were measured in planes perpendicular to extrusion direction.

The variation of grain fiber thickness with solution treatment parameters is shown in

Figure 35.

0.22 0.21 0.20 0.19 0.18 0.17 0.16 0.15 0.14 0.13 Samples solutionized at 532°C

Intercept (mm) 0.12 Samples solutionized at 552°C 0.11 Samples solutionized at 572°C 0.10 As-extruded sample 0.09 As-extruded Sample 0.08 0 200 400 600 800 1000 Solutionizing Time(s)

Figure 35: Variation of intercept or thickness of fibrous grains with solutionizing time. The grain thickness was measured from micrographs from planes parallel to the direction of extrusion.

It should be noted that the grain size measurements were made on one coupon per processing condition and the area available for the grain size measurement was very small. All the measured grain sizes except that corresponding to solutionizing at 532º C

do show grain coarsening with increase in solution treatment duration.

5.5 Results

The metallographic analysis of solutionized AA 6061 coupons seems to indicate

that the thickness of fibrous grain increases with solution treatment time.

The measurements of grain sizes after heating up to solution temperatures but without holding the samples at those temperatures are points corresponding to ‘0’ seconds solutionizing time in Figure 34 and Figure 35.

During the holding time, the thickness slightly increases for samples at 552 ºC and 572 ºC. Increase of holding time for solution treatment at 532 ºC does increase the grain size significantly. These results indicate that solution treatment for longer duration at the higher end of the solution treatment temperature result in grain coarsening and hence shorter solution treatment duration is desired at these temperatures.

5.6 Metallographic Evolution of AA 2618 with Solution Treatment

Similar metallographic analyses were conducted on solutionized AA 2618 coupons. The test solution treatment temperatures were 530 ºC, 535 ºC, 540 ºC and

545 ºC. Since the variations of electrical conductivity and hardness of solutionized AA

2618 did not vary much after 900 seconds of solution treatment, it was decided to conduct metallographic analysis on samples solutionized for 20 seconds and 900 seconds only. These micrographs are shown in following pages.

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 36: Micrograph of AA2618 coupon after Figure 37: Micrograph of AA2618 coupon after solutionizing at 530 ºC for 20 seconds (Mayer, et solutionizing at 530 ºC for 20 seconds (Mayer, et al. 2007). Taken in plane parallel to the al. 2007). Taken in plane perpendicular to the extrusion direction. extrusion direction

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 38: Micrograph of AA2618 coupon after Figure 39: Micrograph of AA2618 coupon after solutionizing at 530 ºC for 900 seconds (Mayer, solutionizing at 530 ºC for 900 seconds (Mayer, et al. 2007). Taken in plane parallel to the et al. 2007). Taken in plane perpendicular to the extrusion direction. extrusion direction.

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 40: Micrograph of AA2618 coupon after Figure 41: Micrograph of AA2618 coupon after solutionizing at 535 ºC for 20 seconds. Taken in solutionizing at 535 ºC for 20 seconds. Taken in plane parallel to the extrusion direction. plane perpendicular to the extrusion direction.

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 42: Micrograph of AA2618 coupon after Figure 43: Micrograph of AA2618 coupon after solutionizing at 535 ºC for 20 seconds. Taken in solutionizing at 535 ºC for 900 seconds. Taken plane perpendicular to the extrusion direction. in plane perpendicular to the extrusion direction.

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 44: Micrograph of AA2618 coupon after Figure 45: Micrograph of AA2618 coupon solutionizing at 540 ºC for 20 seconds (Mayer, et after solutionizing at 540 ºC for 20 seconds al. 2007). Taken in plane parallel to the (Mayer, et al. 2007). Taken in plane extrusion direction. perpendicular to the extrusion direction.

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 46: Micrograph of AA2618 coupon after Figure 47: Micrograph of AA2618 coupon solutionizing at 540 ºC for 900 seconds (Mayer, after solutionizing at 540 ºC for 900 seconds et al. 2007). Taken in plane parallel to the (Mayer, et al. 2007). Taken in plane extrusion direction. perpendicular to the extrusion direction.

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 48: Micrograph of AA2618 coupon after Figure 49: Micrograph of AA2618 coupon after solutionizing at 545 ºC for 20 seconds (Mayer, et solutionizing at 545 ºC for 20 seconds (Mayer, et al. 2007). Taken in plane parallel to the al. 2007). Taken in plane perpendicular to the extrusion direction. extrusion direction.

2.5%HBF4 Electrolytic etchant 2.5%HBF4 Electrolytic etchant

Figure 50: Micrograph of AA2618 coupon after Figure 51: Micrograph of AA2618 coupon after solutionizing at 545 ºC for 900 seconds (Mayer, solutionizing at 545 ºC for 900 seconds (Mayer, et et al. 2007). Taken in plane parallel to the al. 2007). Taken in plane perpendicular to the extrusion direction. extrusion direction.

The photomicrographs indicate the existence of nearly equiaxed grains in the solutionized

AA 2618. For a better understanding of the changes in grain sizes with solution treatment parameters, grain size measurements were performed.

Even though the grains appeared to be equiaxed in the plane parallel to the extrusion direction, the grain size measurements did not confirm this. But the grain size measurements from photomicrographs from planes perpendicular to the extrusion direction confirmed that the grains were equiaxial in that plane. Hence the variations of grain size with solution treatment parameters are treated separately, based on the orientation of the planes analyzed.

Figure 52 shows the variation of grain size on planes perpendicular to the extrusion direction. As mentioned in the experimental section of this chapter, the grain size measurements are averages from two perpendicular directions in the same plane.

38 Solutionized at 530°C 36 Solutionized at 535°C

m) ° μ 34 Solutionized at 540 C Solutionized at 545°C 32

24 Average Average Grain Size ( 22 0 200 400 600 800 1000 Solutionizing Time (s)

Figure 52: Variation of grain size on planes perpendicular to the extrusion direction

On the other hand, grain size measurements from planes parallel to extrusion seemed to be slightly oblong. Hence the intercepts measurements are plotted separately.

Figure 53 and Figure 54 show the variations of the intercepts with the solution treatment parameters.

36 Solutionized at 530°C Solutionized at 535°C 34 Solutionized at 540°C m) μ 32 Solutionized at 545°C

Vertical intercept ( 22

20 0 200 400 600 800 1000 Solutionizing Time(s)

Figure 53: Variation of vertical intercept with solutionizing time. The grain intercept was measured from micrographs from planes parallel to the direction of extrusion.

48 Solutionized at 530°C Solutionized at 535°C m) μ 44 Solutionized at 540°C Solutionized at 545°C 40

Horizontal intercept ( 28

0 200 400 600 800 1000 Solutionizing Time (s)

Figure 54: Variation of horizontal intercept with solutionizing time. The grain intercept was measured from micrographs from planes parallel to the direction of extrusion.

Interestingly, the aspect ratio of the vertical intercept and horizontal intercept measured from micrographs from planes parallel to the direction of extrusion did not seem to change with holding time for most of the cases. For example, the aspect ratio of the grains in the plane parallel to the extrusion direction after holding at 545ºC for 20 seconds is 1.46 whereas it is 1.43 after holding for 900 seconds.

5.7 Conclusions

Based on the physical simulation of solution treatment of AA 2618 and AA 6061 with rapid heating, the following conclusions have been reached. For both AA 2618 and

AA 6061, solution treatment with rapid heating accelerates the solutionizing.

Solutionizing seems to complete within a short period and these times depend on solution temperature. The solution treatment trials on an infrared furnace showed that it took about 20-30 minutes for the samples to reach the solution treatment temperature

(Mayer, et al. 2007). Based on the results of physical simulation and this inherent limitation of the infrared furnace, the following solution treatment cycles were selected as trial cycles for AA 6061 prototype forgings: 30 minutes at 532 ºC, 20 minutes at 552 ºC and 10 minutes at 572 ºC.

Similarly 40 minutes at 530 ºC and 20 minutes at 539 ºC were selected as the rapid solution treatments for AA 2618 forgings prototypes.

In the next chapter, production of prototype forgings and extraction of test specimens for mechanical testing are discussed in detail.

CHAPTER 6: PROTOTYPE FORGING PRODUCTION AND SAMPLE

EXTRACTION

6.0 Introduction

Based on the results from physical simulations of solutionizing with rapid heating and in light of some practical limitations of the infrared furnace at the shopfloor, solutionizing parameters for thermal processing were selected for AA 2618 and AA 6061 forgings. A prototype infrared furnace was installed at the Queen City Forging

Company, Cincinnati, OH for this effort. AA 2618 and AA 6061 prototype forgings were produced and processed with the rapid infrared heating method. The predetermined solutionzing parameters were used for the thermal processing with rapid infrared heating.

For comparison purposes, conventionally processed AA 2618 and AA 6061 forging prototypes were also produced. Prototype forging production with rapid infrared processing and extraction of samples for mechanical testing are covered in this chapter.

Prototype forging production includes the upsetting process and thermal processing with the infrared furnace.

6.1 Aluminum Alloy Materials

AA 2618 and AA 6061 were the alloys used for this work. Kaiser Aluminum,

Newark, Ohio produced the raw material by hot extrusion of DC (Direct Chill) cast billet.

A 5-hole die with an extrusion ratio of 9.2 was used in an indirect press of 5.8MN (6600

US ton) capacity.

The chemical composition of the alloy was determined via atomic emission spectroscopy by the supplier. Table 1 of Chapter 3 compares the chemical compositions

95 of the alloys with corresponding compositions specified by Aluminum Association standards.

Macro analysis of the extruded AA 2618 bars was performed at Oak Ridge

National Laboratory (ORNL) and the resulting macrostructure is shown in Figure 55.

The plane shown is parallel to the direction of the extrusion in the horizontal direction.

The macrostructure shows evidence of the existence of equiaxed fine grains.

Hence, recrystallization might have occurred either during or immediately after the

extrusion process.

However, metallographic analysis of AA 6061 extruded bars showed a different

structure. Figure 56 shows the microstructure of extruded AA 6061 bars.

(a) (b)

Figure 56: Longitudinal and cross sectional micro-structure of AA6061 forging billet in the as- extruded condition (Mayer, et al. 2007). The photo shows a fine, elongated grain structure in the direction of extrusion. Extrusion direction is horizontal to page.

Figure 56a shows the microstructure of the bars in the plane parallel to the

direction of the extrusion. The existence of a distorted or fibrous structure due to extrusion is clearly evident from this microstructure. Microstructure from a plane perpendicular to the direction of extrusion is shown in Figure 56b. These photos indicate that recrystallization did not occur either during or after the extrusion process.

6.2 Forging

Forging and subsequent thermal processing were performed by Queen City

Forging, Cincinnati, OH. Forging billets 152 mm (6 inch) long were obtained from the

extruded bars. Since the aim was to compare the mechanical properties resulting from

conventional processing to that of rapid infrared processing, two types of heating

methods, a conventional gas-fired furnace and flatbed infrared furnace, were used for

preheating of the billets as well as for the thermal processing, following the forging

operation. For some samples (control samples), preheating and solutionizing were

97 accomplished with a conventional gas-fired furnace while for the remainder of the samples, preheating and solutionizing were performed with the infrared furnace.

With conventional heating, it took about 2 hours for the billet to reach the preheating temperature and it was subsequently held at that temperature for an additional hour. With the rapid infrared furnace, it took about 20 minutes to reach the preheating temperature and the heated billet was forged immediately without holding the billet at the forging temperature for any appreciable time. The preheating temperature of AA 2618 was 427 ºC (800 ºF) and the preheating temperature for AA 6061 was 430 ºC- 480 ºC

(806 ºF-896 ºF).

A flatbed type continuous infrared furnace with billets to be forged and forged billets in the foreground is shown in Figure 57.

The forging process was a simple upsetting operation. The selected 152 mm (6 inch) long billets were upset into 76 mm (3 inch) long blocks with a single stroke of a mechanical press. Figure 58 shows the macrostructure of one of the conventionally forged AA 2618 billets. The macro analysis was performed at ORNL.

Figure 58: Longitudinal macro-structure of AA2618 billet after upsetting 50% at 427 ºC (800 ºF) (Mayer, et al. 2007). The grain structure remained similar to that of the as-extruded condition. Pre- heating of this sample was via conventional methods. Extrusion and forging directions are vertical to the page. Photo was taken at ORNL.

The macrostructure indicates the existence of equiaxed fine grains similar to that present

in the extruded AA 2618 bars. Similarly Figure 59 shows the macrostructure of one of

the conventionally processed AA 6061 forging.

Figure 59: Longitudinal macro-structure of AA6061 billet after upsetting 50% between 430 ºC-480 ºC (806 ºF-896 ºF) (Mayer, et al. 2007). The grain structure remained similar to that of the as- extruded condition. Pre-heating of this sample was via conventional methods. Extrusion and forging directions are vertical to the page. Photo was taken at ORNL.

6.3 Infrared Furnace

The newly installed infrared furnace system was fine-tuned to generate the

required heating cycles for the forgings. For this purpose, several trial thermal cycles

were performed. The trial samples were attached with thermocouples in their centers.

The temperatures of the trial samples were monitored and recorded, and the system was

modified accordingly to obtain the required control over temperature.

6.4 Solution Treatment

As mentioned previously, two types of thermal processing were used in the

production of prototype forgings. The conventional thermal processing for AA 2618

consists of solution treatment at 530 ºC (986 ºF) for 8 hours in a conventional furnace.

The corresponding rapid infrared solution treatments determined from the physical

simulation of solution treatment with rapid heating were a solution treatment at 530 ºC

(986 ºF) for 40 minutes and a solution treatment at 539 ºC (1002 ºF) for 30 minutes for

100

AA 2618 forgings. These cycles include a heat up time (time to reach the solutionizing temperature) of 30 minutes with the infrared furnace.

Similarly, the conventional processing cycle includes solution treatment at 529 ºC

(984 ºF) for 8 hours for AA 6061 alloy forgings. As presented in Chapter 5 of this dissertation, the following solution cycles were found to be the ‘optimum solution treatments’ for the rapid infrared processing of AA 6061 forgings: solutionizing at 532 ºC for 30 minutes, solutionizing at 552 ºC for 20 minutes and solutionizing at 572 ºC for 10 minutes.

Table 6.1 summarizes the experimental infrared solution test cycles used for AA

2618 and AA 6061 forging prototypes.

Table 6.1: Experimental Infrared Solution Test Cycles for AA 2618 and AA 6061 Alloy Forging Prototypes (Mayer, et al. 2007). Alloy Solution Temperature Solution Time

532 ºC (990 ºF) 30 min.

6061 552 ºC (1026 ºF) 20 min.

572 ºC (1062 ºF) 10 min.

530 ºC (986 ºF) 40 min. 2618 539 ºC (1002 ºF) 20 min.

The solutionized forgings were quenched in a water bath at 60-70 ºC (140-160 ºF)

as specified by the Aerospace Material Specification, AMS 2770E, Heat treatment of

101 wrought aluminum alloy parts (AMS 2770E Heat Treatment of Wrought Aluminum

Alloy Parts 1989). The quench delay was kept within 15 seconds which is maximum allowable delay specified by the same reference.

Solutionized AA 2618 forgings were artificially aged to T61 temper by ageing at

199 ºC (360 ºF) for 20 hours. Solutionized AA 6061 forgings were artificially aged to T6 temper by ageing at 175 ºC (347 ºF) for 8 hours.

6.5 Sample Extraction

The test specimen geometry for tensile and fatigue tests were designed based on

ASTM E08-03 (Standard Test Methods for Tension Testing of Metallic Materials 2004) and ISO 1143 (International Standard 1143 1975), respectively. Based on the required number of tests, different numbers of samples were extracted from the forged parts. For example, Figure 60 and Figure 61 show schematic diagrams illustrating the extracted samples from conventionally processed AA 2618 forgings.

102

Figure 62 shows a photograph of extracted samples and the remaining forging prototype. The rod samples were extracted by wire EDM.

103

Final test sample geometry was obtained by machining in a computer numerically controlled (CNC) lathe. In the next chapter, the experimental methodology is covered in detail.

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CHAPTER 7: EXPERIMENTAL METHODOLOGY

7.0 Introduction

The mechanical testing methods used in this work are discussed in this chapter.

They include tensile testing, hardness testing, high cycle fatigue and strain-controlled low cycle fatigue testing.

7.1 Hardness Testing

The hardness testing was performed based on ASTM E18-03, standard test methods for Rockwell hardness and Rockwell superficial hardness of metallic materials.

A Wilson hardness tester was used for the testing. The load used for the testing was

100kg and the indenter used was 1/16 inch ball type (Standard Test Methods for

Rockwell Hardness and Rockwell Superficial hardness of Metallic Materials, 2004). The

Rockwell B scale was used for the testing. The surfaces of the forged and thermally processed blocks were polished with 180 grit and 320 grit silicon carbide paper on a Leco

System 2000 polishing wheel before conducting the hardness measurements. Figure 63 shows the polished flat surface of the forged block to be tested in the hardness testing apparatus.

105

Figure 63: Polished flat surface of the forging being tested for hardness.

Five hardness readings were taken at different locations of the polished flat surface for each of the forgings.

7.2 Tensile testing

Tensile test was performed according to ASTM E8-04, standard test methods for tension testing of metallic materials (Standard Test Methods for Tension Testing of

Metallic Materials 2004). Figure 64 shows the geometry of a standard tensile test sample obtained from the rod samples of the processed forgings. The sample has a section with a uniform cross sectional area in the middle and two threaded sections at both ends. The threaded sections were inserted into specially made grips that were being held with grips of the servo-hydraulic MTS 810 system used for the tests.

106

6mm

25mm

Figure 64: Tensile sample extracted and machined from the forging prototype. It has treaded sections at both ends to facilitate attaching the sample to the grips for loading.

It was decided to test one tensile test specimen from each of the processed

forgings. More than one forging was made for some of the thermal processing

conditions. Correspondingly more than one tensile test specimen was obtained for these

processing conditions and tensile properties were measured. The MTS machine used for the testing was controlled by a computer with Teststar IIs and Multipurpose Testware software. Programs can be written using sequence of commands in Multipurpose

Testware. Commands are used for applying different types of loading with the MTS

machine, for recording data dynamically, and even for controlling external devices such

as a furnace connected to the system. Once the tensile sample was inserted into the grips

that had already been attached to the MTS 810 system, a 25.4 mm (1 inch) extensometer

with the pre-aligned safety pin was attached to the sample. The extensometer was

107 attached to the uniform section of the sample with small elastic bands. Then, the safety pin was carefully removed.

A test program was created for the tensile test. The first command in the program was to start real-time recording of strain, loading force, displacement and time. The second command was for pulling the sample such that it would elongate 0.254 mm (0.01 inch) in 10 seconds in the gauge section. This would result in a strain rate of 0.001/s.

Once the sample was elongated by 0.254 mm, the program paused the test for 40 seconds to facilitate the removal of extensometer. The extensometer was used in the initial step of the loading to record the strain during the initial loading and the strain data was later used for the calculation of the yield strength of the sample tested. Then the extensometer was removed so that the specimen could be pulled to fracture without damaging this instrument.

The ultimate tensile strength, 0.2 % yield stress and elongation at failure were obtained from the tensile test.

7.3 High Cycle Fatigue Testing

High cycle fatigue testing was performed per ISO standard 1143, Metals-

Rotating bar bending fatigue testing (International Standard 1143 1975). An RR Moore rotating beam instrument was used for this testing.

7.3.1 R.R. Moore Testing Machine

An Instron R.R. Moore fatigue machine was used for the tests. Figure 65 shows the testing machine.

108

Figure 65: Photograph of Instron, R.R. Moore fatigue testing machine.

In this system, the test sample is a “rotating beam” under load. In the test section

(at the section where the diameter is the minimum) the stress is the maximum. The lower

portion of this section is under tensile stress while the upper portion is under compressive stress. As the sample rotates, a maximum cyclic stress state is applied to the sample

section at the minimum cross sectional area. The speed of the rotation can be adjusted

from 1 to 10000 rev/min. However the suggested rotation speed for the test was 9000

rev/min, and thus the samples in this test were rotated at 9000 rev/min.

A battery powered meter in the system monitors and displays the number of

rotations of the sample. This is shown in Figure 66.

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Figure 66: The rotation counter in the R.R. Moore fatigue tester. The black knob at the lower left is for changing the speed of revolution.

The test system has a trigger to detect the failure of the sample and to immediately stop the rotation of the sample.

7.3.2 Sample Preparation

The geometry of the high cycle fatigue test sample is shown in Figure 67.

110

Figure 67: Geometry of the high cycle fatigue samples.

The samples were machined from the bars extracted from the forging prototypes.

A CNC lathe was used for the machining. Specific machining guidelines set forth in ISO

1143 were used, specifically depth of cut for finishing (International Standard 1143

1975). The machined samples were then polished with 600 grit and 900 grit silicon carbide paper on a table top lathe.

The sample achieves maximum stress and failure at its mid section, the location of minimum cross sectional area. The diameter was measured with a custom-made device.

Figure 68 shows this device with a sample being measured. The anvils of the device were changed to tips made of nylon to prevent them from scratching the polished surface of the test sample.

111

Figure 68: Measuring device to check the minimum diameter of the sample without damaging the surface. Nylon anvils attached to end of edges prevent damage of the aluminum sample surface.

7.3.3 Theory

The design of the loading system of the rotating beam is that the maximum stress in the test sample is independent of the distance between the loading points (or the length of the sample). The free body diagram of the sample held by two sample grips is shown in Figure 69.

112

W is the load applied to the sample grips. L is the distance between the applied

loads. R is the reaction force on the sample grip at the support point and D is the distance

between the support point in the grip and the applied load in the grip, and this is a fixed

length for the grip.

For equilibrium for the system in the vertical direction

2 2

(1)

Consider the portion of the beam to the left of XX, a section X distance from the left

support.

Equating momentum at that point provides

(2)

Substituting for R from Equation (1),

113

(3)

Therefore, at any point between the loads (W), the moment is WD, independent of the distance between the loading points, or the length of sample to be tested.

In order to derive maximum stress concentration in the sample, the gauge section of the sample at the minimum cross sectional area is considered.

The bending moment formula is,

σ M = (4) y I x

where σ is the bending stress at radius of y and M is the moment at the neutral axis. Ix is the area moment of inertia about the neutral axis. The bending stress is the maximum at the edge where the radius y is the maximum.

The area moment of inertia for a circular cross section of the fatigue test specimen is,

π I = d 4 (5) x 64

Where d is the diameter of the circular cross section.

Substituting, Equations (3) and (5) in (4),

σ WD = 64 (6) y πd 4

and

WD σ = 64 y (7) πd 4

Considering the stress at the lower edge of the cross section where y = d/2, the stress can

be expressed as,

WD σ = 32 (8) πd 3

114

Similarly, the stress at the upper edge of the cross section can derived by substituting for y= -d/2.

WD σ = −32 (9) πd 3

As the fatigue sample rotates, the stress state at any point in the outer radius of cross

section varies between the stresses calculated using Equations (8) and (9).

Hence the stress ratio is -1. The stress varies between a tensile state and a compressive

state.

In high cycle fatigue testing, the load applied is changed to conduct high cycle

fatigue tests at different stress levels. Due to the geometric symmetry of the loading set

up, W in Equation 8 is half of the loading weight (Wl).

The load to be applied for a desired maximum stress was found using Equation

(8) (Operating Instructions, Instron Model R.R.Moore High Speed Rotating Beam

Fatigue Testing Machine n.d.).

πd 3σ Wl = 16D (10)

D is 101.6 mm (4 inches), a fixed length for the RR Moore fatigue testing machine used.

High cycle fatigue tests at approximately 50%, 60%, 70% and 80% of the 0.2%

yield stress of the specimen material were conducted for this work. The sequence of the

tests was designed based on random order.

7.4 Low Cycle Fatigue Testing

The same servo hydraulic MTS system used for the tensile testing was used for

the low cycle fatigue testing. A loading cycle with 1 Hz frequency was used for the

115 testing. The test was axial strain controlled. The MTS system applied a sinusoidal strain loading cycle with specified maximum and minimum strains. The MTS system measured and recorded the strain via a 25.4 mm (1 inch) extensometer attached to the sample being tested. Figure 70 shows the extensometer attached to a sample.

Figure 70: Low cycle fatigue sample fixed to the grips in the MTS machine. An extensometer is attached to the gauge section of the sample to measure and feedback the strain in the sample.

The low cycle fatigue tests were performed according to ASTM E606-92,

standard practice for strain-controlled fatigue testing (Standard Practice for Strain-

Controlled Fatigue Testing 2004). The test sample had a uniform cross section and two

threaded ends. The sample is similar to that shown in Figure 64. The geometry was

designed based on the guidelines provided by ASTM E606-92 and the size of the

available rods extracted from the forging prototype. The machining was performed in a

CNC lathe. The machining sequence and parameters were based on the ASTM

116 guidelines for axial strain controlled fatigue testing. After machining, the surface of the uniform cross section was polished with 600 and 900 grit silicon carbide paper in a table- top lathe.

7.4.1 Loading Cycles

All the loading cycles applied were with a strain ratio (R) value of -1. Loading cycle with maximum strains between 0.2 and 1.2 % were used. These maximum strains included 0.2%, 0.3%, 0.56%, 0.8%, 1.0% and 1.2%. A sample dynamic plot of the loading is shown in Figure 71.

Figure 71: A computer screen example of the dynamic variation of applied force and measured strain during low cycle fatigue testing.

117

As in the case with the high cycle fatigue tests, random order was used for low cycle fatigue testing sequence.

The results of the tests are presented in the next chapter. This also includes the results of mechanical testing of samples extracted from AA 2618 and AA 6061 forging prototypes. These results are compared with corresponding mechanical properties provided in the ASM handbook.

118

CHAPTER 8: RESULTS

8.0 Introduction

In this chapter, the results of mechanical testing of samples from AA 2618 and

AA 6061 forging prototypes are presented. These include results of tensile testing, hardness testing, grain size measurements, and high cycle and low cycle fatigue testing.

These results are compared with specifications and data from the Aerospace Structural

Materials Handbook and ASM standards.

The results of testing of the two alloys are treated separately for simplicity. In addition, grain size and hardness measurements on prototype AA 2618 forgings of different geometry are presented. These measurements were performed by Queen City

Forging, Cincinnati, OH, a project partner.

8.1 Results of Testing of AA 2618 Forgings

The previous chapters presented the mechanical and thermal processing involved in the preparation of the AA 2618 prototype forgings. The chemical composition of the alloy is also presented in Chapter 4. Chapter 7 provided a detailed account on the mechanical testing procedures used for this research.

Uniaxial tensile testing, hardness testing and fatigue testing were performed on samples obtained from AA 2618 forging prototypes. The results follow in the next section.

119

8. 1. 1 Mechanical Property Comparisons

Table 8.1 lists mechanical properties of conventionally processed AA 2618 forgings with those treated using rapid infrared heating. Two different rapid infrared thermal solution cycles were used. The first infrared solution cycle consisted of an infrared solution treatment at 539 ºC for 20 minutes. Three prototype forgings underwent this solution treatment. The second infrared solution treatment cycle consisted of an infrared solution treatment at 530 ºC for 40 minutes. Another three prototype forgings underwent this cycle. Three more prototype forgings were conventionally solutionized at

530 ºC for 8 hours. All the prototypes were subsequently aged to T61 condition. The measured mechanical properties are compared with the minimum specified values for this alloy forging conventionally processed to T61 condition.

Table 8.1: Tensile and hardness test values for infrared solution treated AA 2618 samples and conventionally solution treated AA 2618 samples (Mayer, et al. 2007). All samples underwent a standard aging cycle also. Each sample was taken from a different forging upset of that lot. The bottom entries in the table are handbook specifications for comparison (Brown and Setlak, Code 3213, page 2, 2003). 0.2% yield Tensile Elongation Hardness Condition Sample stress stress (%) (HRB) ksi (MPa) ksi (MPa) IR solution- 1 51.5 (355) 62.9 (434) 12.0 75.8

treated 2 49.0 (338) 64.4 (444) 9.9 75.6

20 min @ 539ºC 3 49.0 (338) 61.7 (425) 10.0 75.5

IR solution- 1 50.0 (354) 62.4 (430) 10.1 75.5

treated 2 50.0 (345) 62.3 (430) 10.0 76.0

40 min @ 530ºC 3 49.5 (341) 61.7 (425) 10.1 75.8

Conventionally 1 49.0 (338) 60.4 (416) 11.1 74.0

120 solution-treated 2 51.2 (353) 62.6 (432) 11.4 75.6

8 hrs @ 530ºC 3 50.2 (344) 62.7 (432) 12.3 74.2

Minimum N/A 48.0 (331) 58.0 (400) 6.0 72.0 Specification

The results indicate that the static mechanical properties of AA 2618 forgings

treated with rapid infrared heating are equal to or better than that of conventionally

processed AA 2618 forgings. All the measured mechanical properties of infrared treated

and conventionally processed forgings were above the minimum values specified for the

corresponding alloy forging.

8.1.2 High Cycle Fatigue Testing

Fatigue testing included high cycle fatigue testing and low cycle fatigue testing.

Figure 72 compares the results of high cycle fatigue testing of AA 2618 prototype

forgings with corresponding data for AA 2618 forgings from Aerospace Structural

Metals Handbook (Brown and Setlak, Code 3213, page 16, 2003).

121

440 63.8

Data from Aerospace 400 Structural Metals Handbook 58.0 ASM Handbook Endurance Limit Conventional solution treatment and age 360 IR solution treated for 52.2 20 min. at 539 oC, conventional age 320 IR solution treated 46.4 for 40 min. at 530 oC, conventional age 280 40.6

240 34.8

200 29.0 Maximum Stress (ksi) Stress Maximum

Maximum Stress (MPa) Stress Maximum 160 23.2

120 17.4

103 104 105 106 107 108 Cycles to Failure

In addition, the endurance limit for AA 2618 forging from ASM handbook is included in the plot. The infrared treated AA 2618 seems to have better fatigue properties than published data for the conventionally treated AA 2618 forgings. This enhancement in fatigue behavior appears to be more pronounced in the lower stress region of the fatigue data.

8.1.3 Low Cycle Fatigue

Figure 73 shows the comparison of the results of the low cycle fatigue testing of conventionally treated and infrared treated AA 2618 forgings.

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1.0 IR solutionized for o o 0.9 20 min @ 539 C (1002 F) IR solutionized for 0.8 40 min @ 530 oC (985 oF) conventionally solutionized for 0.7 8 hrs @ 530 oC (985 oF) 0.6

0.5 Strain (%) 0.4

0.3

0.2

0.1 100 1000 10000 100000 1000000 Cycles to Failure

Figure 73: Low cycle fatigue data for conventionally heat-treated and rapid infrared treated AA 2618 samples (Mayer, et al. 2007).

Both conventionally treated AA 2618 forgings and rapid IR treated AA 2618

forgings seem to exhibit similar fatigue behavior. These results show that the low cycle

fatigue properties of infrared treated AA 2618 forgings are on par with those of

conventionally processed AA 2618 forgings.

Variation of peak nominal stress with peak strain measured during the half life loading cycles for the low cycle fatigue tests are shown in Figure 74. The Young’s modulus of AA2618 is also shown in the plot.

123

450 65

400 58

350 51

300 44

250 36 ksi 200 29

150 22

100 Conventionally Processed 15 Maximum nominal stress in the loadingcycle (MPa) Rapid IR Processed 50 Elastic Modulus of AA2618 7 0 0 0.0 0.2 0.4 0.6 0.8 1.0 Maximum strain in the loading cycle %

Figure 74: Cyclic stress-strain for AA 2618 T61. The stress and strain measurements were taken at half life loading cycle.

For loading cycles with maximum strain values below 0.3%, the stress-strain behavior did not deviate from the Young’s modulus. The recorded data corresponding to the loading cycles were analyzed. The results showed that the stress ratio for the all of the tests did not have a stress ration of -1 even though the all the tests were programmed to have -1 as the stress ratio. The plot shows that forging prototypes processed with infrared heating and conventionally processed forging prototypes responds in similar manner for similar strain cycles.

8.2 Results of Testing of AA 6061 Forgings

The results of mechanical testing of conventionally treated and rapid infrared treated AA 6061 forgings are presented in this section. Corresponding mechanical properties of conventionally treated AA 6061 as specified in the standards (Brown and

Setlak, Code 3213, page 2, 2003) (Brown and Setlak, Code 3213, page 16, 2003) (Bray

124

2002) (Brown and Setlak, Code 3206, page 5, 2003) (Brown and Setlak, Code 3206, page

15, 2003) are also presented where applicable.

The mechanical properties include tensile stress, 0.2% yield stress, elongation, average hardness, high cycle fatigue and low cycle fatigue. These are presented in detail in the following sub-sections.

8.2.1 Mechanical Property Comparisons

As previously mentioned, AA 6061 forging prototypes were made from two batches of billets and they were thermally processed with different solution treatment parameters. Two forgings from lot 1 were conventionally processed at 532 ºC for 480 minutes. That is, using the conventional furnace it took 4 hours for the forgings to reach the solution temperature of 532 ºC and then they were held at that temperature for another

4 hours. Another three AA 6061 forgings were produced from lot 1 and they were thermally processed using the infrared furnace. These three forging prototypes were solution treated at 552 ºC for 20 minutes. The heat-up times for these samples were 20-

25 minutes.

The rest of the forging prototypes were made with billets from lot 2. All these forgings were thermally processed with the rapid infrared heating method. Two of these forgings were infrared solution treated at 552 ºC for 20 minutes. Three of these forgings were solution treated with infrared heating at 572 ºC for 10 minutes and the remaining three forgings were infrared solution treated at 532 ºC for 30 minutes.

One tensile sample was obtained from each of the forging prototypes and tested for tensile properties. The hardness was measured on the polished flat surfaces of the

125 forgings. Table 8.2 presents the measured mechanical properties of infrared treated and conventionally treated AA 6061 forgings. For comparison, corresponding data for conventionally processed AA 6061 forgings from ASM handbook (Brown and Setlak,

Code 3206, page 5, 2003) is presented in Table 8.2.

Table 8.2: Tensile and hardness test values for infrared solution treated samples and conventionally solution treated AA 6061 samples (Mayer, et al. 2007). All samples underwent a standard aging cycle. Each sample was taken from a different forging upset of that lot.

Solution Solution Ave. 0.20% Tensile % Lot Temp. time Hardness yield stress Stress Ave. Ave Elong (ºC) (min) HRB Ave. Ave. ksi MPa MPa ksi MPa MPa 532 480 58.7 44.5 307 49.5 341 4.4 1 59.0 306 338 7 532 480 59.2 44.1 304 48.4 334 8.7 552 20 64.4 47.5 328 51.2 353 6.4 1 552 20 66.6 66.0 44.0 303 314 51.0 352 351 12 8 552 20 67.0 45.0 310 50.3 347 5.5 552 20 65.6 43.5 300 49.6 342 12 65.7 298 341 12 2 552 20 65.8 43.0 297 49.3 340 11.1 572 10 67.2 41.2 284 47.2 326 10.9

572 10 66.2 66.2 43.2 298 293 49.1 339 334 12.3 12 2 572 10 65.2 43.0 297 49.1 339 12.9 532 30 66.0 43.8 302 50.0 345 10.3

532 30 65.4 65.4 43.0 297 298 48.9 337 340 12.8 11 2 532 30 64.8 43.0 297 49.2 339 10.8 Minimum Specification for 35.0 241 38.0 262 10 AA 6061-T6 Forgings (Bray 2002) (a) From Table 3.013 of (Brown and Setlak, Code 3206, page 5, 2003), data corresponding to forging, specimens machined from separately forged coupons or from stock representing the forging and in either case heat treated with the forging or machined from prolongations on the heat treated forging and also at T6 condition.

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The hardness of the infrared processed AA 6061 forgings seems to be slightly higher than that of conventionally processed AA 6061 forgings. The mechanical properties of forgings from lot 1 slightly differ from that of lot 2. In general, it seems that samples from lot 1 exhibit somewhat better strength but slightly lower elongation, regardless of the type of heating used for thermal processing.

The yield strength and tensile strength of infrared heat treated forgings seems to be equivalent to those of conventionally treated forgings. However the hardness of infrared treated AA 6061 forgings is about 6HRB units greater than the hardness of conventionally treated AA 6061 forgings. The reason for this improvement will be discussed in the discussion section of the next chapter.

All the measured properties of AA 6061 forgings seem to be greater than the minimum specified properties for AA 6061 forgings. However, measured elongation values are less than the minimum specified percentage elongation for AA 6061 forgings.

The macrographs of AA 6061 prototype forgings shown in Figure 59 in Chapter 6 shows existence of elongated grains and the axial direction of the tensile samples was parallel to the elongated grains. Hence it is speculated that the elongated grains had caused the reduced percentage elongation in the tensile tests.

8.2.2 High Cycle Fatigue

High cycle fatigue testing was performed on samples obtained from the AA 6061 prototype forgings. The results of these tests are plotted in Figure 75 along with corresponding data for conventionally processed AA 6061 from the Aerospace Metals

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Handbook (Brown and Setlak, Code 3206, page 15, 2003). Different symbols in the plot correspond to samples from forgings with different thermal history.

52 359

532 oC for 8 hours - conventional (lot 1) 48 331 552 oC for 20 min - IR heating (lot 1) o 44 552 C for 20 min - IR heating (lot 2) 303 572 oC for 10 min - IR heating (lot 2) o 40 532 C for 30 mins - IR heating (lot 2) 276 Aerospace Structural Metals Handbook 36 248

32 221

28 193 Maximum Stress (MPa) Maximum

Maximum Stress (ksi) Stress Maximum 24 166

20 138

16 110 1000 10000 100000 1000000 1E7 1E8 Cycles to Failure

The high cycle fatigue behavior of infrared processed AA6061 forging samples

seems to be equivalent to that of conventionally processed AA 6061 samples. Within the

selected range, the infrared solution treatment parameters appear to have no effect on the

resulting high cycle fatigue behavior. The high cycle fatigue behavior of the tested

forgings is on par with that of conventionally processed AA 6061 forgings as specified in

the Aerospace Structural Handbook (Brown and Setlak, Code 3206, page 15, 2003).

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Hence, the results indicate that the infrared thermal processing of AA 6061 forgings results in high cycle properties comparable to that obtained with conventional thermal processing.

8.2.3 Low Cycle Fatigue

Similarly, low cycle fatigue data from specimens from AA 6061 forging prototypes was obtained. Figure 76 shows the results of the testing.

1.2 552 oC for 20 min - lot 1 1.1 o 552 C for 20 min - lot 2 1.0 572 oC for 10 min - lot 2 532 oC for 30 min - lot 2 0.9 Conventionally processed- lot 1 0.8 0.7

0.6 Strain % Strain 0.5 0.4 0.3 0.2 10 100 1000 10000 Cycles to failure

Figure 76: Low cycle fatigue data for conventionally heat-treated and rapid infrared treated AA 6061 samples (Mayer, et al. 2007).

The low cycle fatigue testing of AA 6061 specimens was conducted at strain

levels of 0.3%, 0.56%, 0.8% and 1.0%. However, fatigue specimens from conventionally

129 processed AA 6061 deformed in the first loading cycle of the 1% strain cycle and hence the test was discontinued for those samples. The results show that the low cycle fatigue behavior of infrared processed AA 6061-T6 samples is comparable to that of conventionally processed AA 6061.

Peak nominal stress values and peak strain values were measured from the half life cycles of the low cycle fatigue tests of samples extracted from AA 6061 forging prototypes. The recorded data corresponding to the loading cycles were analyzed. As in the case with tests with AA 2818 specimens, the results showed that the stress ratio for the all of the tests did not have a stress ration of -1 even though the all the tests were programmed to have -1 as the stress ratio. Figure 77 shows the variation of peak nominal stress with corresponding peak strain of the half life loading cycles. The elastic modulus of AA6061 is plotted in the graph for comparison.

400 58

350 51

300 44

250 36

200 29

ksi 150 22

100 Conventionally Processed 15 Rapid IR processed Maximum nominal stress in the loading cycle (MPa) 50 Elastic Modulus of AA6061 7

0 0 0.0 0.2 0.4 0.6 0.8 1.0 Maximum strain in the loading cycle (%)

Figure 77: Cyclic stress- strain for AA 6061 T6. The stress and strain measurements were taken at half life loading cycle.

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The response to loading cycles with peaks strain values below 0.3% was similar

to the response of linear loading. AA 6061 forging prototypes processed with infrared

heating and conventionally processed AA 6061 forging prototypes responds in a similar

manner for similar strain cycles.

8.3 Grain Size Measurement of Conventionally Treated and IR Solutionized AA

2618 Forging Prototypes

Sections of failed tensile test specimens were mounted on mounting medium for

optical microscopy for grain size measurement. These sections were extracted from

regions away from the gauge section. ASTM line intercept method was used for the

grain size measurement ( Standard Test Methods for Determining Average Grain Size

2004). Figure 78 shows two photomicrographs obtained for this purpose.

(a) (b)

131 et al. 2007). The section shown in (a) is parallel to the forging direction and the section in (b) is transverse to the forging and extrusion directions.

Table 8.3 presents measured grain sizes and the corresponding thermal processing history. All these forgings had been aged to 2618-T61 condition following the solution treatment.

Table 8.3: Average grain size of AA 2618 forgings after different solution treatments (Mayer, et al. 2007).

Condition Grain size (μm)

IR solution-treated 20 min @ 539 ºC 34

IR solution-treated 40 min @ 530 ºC 34

Conventionally solution-treated 8 hrs @ 530 ºC 30

These measurements contradict previously reported (Kervick, et al. 2006) (Lu, et al. 2004) reduction of grain size in AA 2618 forgings with rapid infrared heating. Similar measurements on AA 6061 samples were not successful since they had elongated grains and the available facilities were not enough to measure those types of grain sizes.

8.4 Hardness and Grain Size Measurements of AA 2618 Prototype Part Production

The Queen City Forgings, a project partner, had conducted hardness and grain size measurements at different sections of a prototype part forging made of AA 2618

132 alloy. The shape of the forging was complex. Further details of the forgings could not be discussed due to the proprietary issues.

All these forgings were infrared solution treated at 530 ºC. The soak time at solutionizing temperature ranged from 10 minutes to 2 hours. Hardness measurements were performed on these forgings and Figure 79 shows the variation of measured hardness with soak time.

017 33 50 67 83 100 117 133 76 74 72 70 68 66

64 62

Hardness (HRB) Hardness 60 58

56 HRB of Prototypes solutionized at 530 oC

0 1000 2000 3000 4000 5000 6000 7000 8000 Solution Time (s)

The hardness variation shows that forgings with soak times lesser than 40 minutes

results in considerable variation in hardness values. Hence, a minimum soak time of 40

minutes is required to minimize the hardness variation.

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Similarly the grain size measurements were performed at flashpoints and at part centers on these forgings. These two locations correspond to high deformation and low deformation region respectively. The variation of grain size at these regions with the soak time is shown in Figure 80.

Minutes 0 1733506783100117133

100 At Flashline: High Deformation Region 90 At Part Center: Low Deformation Region

m) 70 μ 60

Grain Size ( 40

0 1000 2000 3000 4000 5000 6000 7000 8000 Solution time (s)

Figure 80: Variation of grain size in forged and IR solution treated prototype parts (Mayer, et al. 2007). Error bars represent ±1 standard deviation. (Data provided by QCF)

The plot clearly shows that at high deformation region, the variation in grain size is relatively very less and the soak time at solutionizing temperature does not affect the variation of grain size. On the other hand, the grain size varied significantly at regions with low deformation and the variation seems to increase after 40 minutes of soaking at

530 ºC.

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The hardness measurements and grain size measurements of the part prototype forgings confirms that rapid infrared solution treatment for 40 minutes would produce optimal properties for this part.

8.5 Conclusions

Based on the comparisons of the results of the mechanical testing of conventionally and infrared solution treated samples, the following conclusions can be made.

The static mechanical properties of AA 2618 forgings processed with infrared heating are similar to that of AA 2618 forgings processed with conventional heating.

These mechanical properties include tensile strength, yield strength, elongation before failure and hardness.

The results of high cycle fatigue testing of AA 2618 forgings seems to indicate that thermal processing with infrared heating slightly improves the high cycle fatigue response. The improvement is more evident in the test region with lower stresses.

However, the results of low cycle fatigue tests indicate that the low cycle fatigue properties of AA 2618 forgings processed with infrared heating and conventionally processed AA 2618 forgings are equivalent.

Prototype forgings made from aluminum alloy stocks from different lots showed slight variations in mechanical properties. However, within the same AA 6061 lot, the IR processing seems to produce forgings with slightly better static mechanical properties.

Both conventional thermal processing and IR thermal processing resulted in similar high-cycle fatigue and low cycle fatigue properties in AA 6061 forgings.

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In the next chapter, the conclusions derived from this research and recommendations for future works are discussed.

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CHAPTER 9: DISCUSSION

Several aspects of rapid heating of aluminum alloys have been investigated for this work and this section summarizes the work and interrelates the results (Mayer, et al.

2007). First, all the research work performed for this project is summarized. Then the results are discussed with respect to the following four aspects: time and energy savings due to rapid heating, accelerated solutionizing with rapid heating, control of grain size and the resulting mechanical properties of aluminum alloy forgings processed with the rapid heating method.

9.1 Summary of the Work

Baseline data from conventional thermal processing of the forgings of the selected aluminum alloys were obtained. Available literature in this area was reviewed to obtain information in determining an optimal solution treatment and application of other rapid heating for processing aluminum alloys. In addition, works investigating the solid state reaction during the thermal processing were also reviewed.

A series of experiments were conducted to physically simulate the thermal processing of aluminum forgings with rapid heating. Using the results of these experiments, the degree of dissolution of second phases was measured. The evolution of the grain structure during these thermal cycles with rapid heating was also studied.

Based on the results of the physical simulation of thermal processing aluminum alloys and the limitations and capacities of the infrared furnace, ‘optimum’ thermal cycles for solution treatment of AA 2618 and AA 6061 alloy forgings were selected.

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Prototype forgings were produced and individually processed with the selected thermal cycles with rapid infrared heating. In addition, conventionally processed prototype forgings were also produced. Mechanical tests were conducted on specimens obtained from these forgings.

9.2 Discussion

The results of the work are discussed under four aspects. They are the savings in time and energy due to rapid heating, accelerated solutionizing with rapid heating, control of grain size and the resulting mechanical properties of aluminum alloy forging processed with the rapid heating method.

9.2.1 Direct Energy Savings

The savings resulting from rapid heating of aluminum alloys are shown to be from the efficiency in energy transfer from the energy source to the part being heat treated, and the reduction in solution treatment time. Physical simulation of short solution treatment with rapid heating and comparison of mechanical properties of forging prototypes processed with rapid infrared solution treatment and conventionally processed forging prototypes confirmed that short solution treatment with this rapid heating method can produce forgings that are on par with conventionally processed forgings. In some cases, the rapid heating method produced better mechanical properties than conventional processing.

As the second phase precipitates and the resulting grain sizes of the aluminum alloy forgings are interrelated, they would be discussed together in the next section.

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9.2.2 Second Phase Particles and Control over Grain Size

Physical simulation of solution treatment of the selected aluminum alloys with rapid heating confirmed that the solution treatment completes faster when a rapid heating method is used. One reason for faster solutionizing can be the thermal stresses due to the rapid heating rate and the difference in thermal expansion coefficients of the second phase particles and the aluminum matrix. Rapid heating increases the temperature difference between the temperatures of the aluminum matrix and second phase particles.

This temperature difference along with the difference in thermal expansion coefficients creates thermal stresses. These thermal stresses accelerate the diffusion process and thereby improve the dissolution of second phases during rapid heating. On the other hand, the temperature difference between the second phase particles and aluminum matrix during slow heating is less and hence the thermal stresses during slow heating is much lower. A work reviewed in the open literature observed similar behavior in cast aluminum alloy A356 treated with rapid heating (Chaudhary, et al. 2003).

This observation is in agreement with reference (van de Langkruis, et al. 2000) even though this reference did not explicitly deal with the thermal stresses associated with different heating rates. In this published work, dissolution of Mg2Si was

theoretically modeled and analyzed in terms of initial Mg2Si particle size, heating rate

and holding time. Based on the results, dissolution diagrams were constructed.

Solution treatment closer to the eutectic temperature accelerates the dissolution of the

second-phase particles, and the better temperature control of an infrared furnace can be

used to take advantage of this phenomenon.

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The measurement of grain size at different locations of the prototype forgings has shown that grain size differed based on the amount of deformation. As Figure 80 in

Chapter 8 shows, the grain size at the high deformation zone of AA 2618 forgings is comparatively less than that at the low deformation zone of the forgings. This is consistent with the fact that a high amount of deformation refines grains. Another important thing to note from Figure 80 is the variation of grain size with solution treatment time. The grain size at the high deformation zone stayed between 20-30µm even with longer solution treatment durations, such as 7000 seconds. On the other hand, grain size changed from 25-45µm to 25-100µm when the solution treatment time was increased to above 7000 seconds. The variation in the initial grain size can be attributed to the amount of deformation in the forgings. However, the grain coarsening occurred in the low deformation region with increased solution treatment time, while that was not the case in the high deformation region, and the variation of the amount of deformation cannot be directly considered as responsible for this.

The retardation of grain growth can be attributed to the precipitates present in the forgings (Lu, et al. 2004) (Kervick, et al. 2006). Reference (Kervick, et al. 2006) attributes the grain refinement to the pinning effect of aluminum-iron-nickel precipitates

(Al9FeNi). Hence the precipitates present in the highly deformed regions of the forgings had been favorable for retardation of grain growth than those at regions with low amount of deformation. Several researchers have explored the effects of amount of deformation into the precipitation kinetics (Liang, et al. 2007) (Quainoo and Yannacopoulos 2004).

Reference (Liang, et al. 2007) reports that, the right amount of deformation is favorable for desirable precipitation in terms of uniform distribution of precipitates and resulting

140 strength of the Al-Cu-Li-Zr alloy containing Sc. It is suggested that high deformation region of the aluminum alloy prototype forgings created uniform distribution of precipitates and thus exhibited retardation of grain growth. Conversely, the low deformation regions of the prototype forgings ended up with heterogeneous precipitates which in turn did not retard grain growth uniformly in the low deformation regions. As a result, the grain size ranged between 25-100µm. That is, the precipitates had retarded the coarsening of grains in the high deformation regions while the precipitates in the low deformation regions did not retard grain coarsening.

It should be noted that a previous work in the application of infrared heat treatment of aluminum alloy forgings had reported a grain size reduction from 40µm

(conventionally processed forging) to 25 µm due to infrared processing of AA 2618 forgings (Kervick, et al. 2006). At the same time, forging preheated with infrared heating and solution treated with a conventional heating method also resulted in a grain size of

27µm in the current work. This seems to indicate that infrared solution treatment after forging did not have much effect on the grain size. The reported grain size 25 µm of the infrared treated AA2618 forging is consistent with that measured at the high deformation region of the AA2618 prototype forgings in this project. On the other hand, grain size measurements conducted at Ohio University differ slightly from these. Table 8.3 shows that the grain size of conventionally processed AA 2618 prototype forging had a grain size about 30 µm while the infrared treated AA 2618 prototype forging had a grain size about 34 µm. Comparing these two grain sizes with ranges of grain sizes measured by

Queen City Forgings i.e., from 25-45µm to 25-100µm, it can be said that a significant conclusion cannot be made from the grain size measurements of the prototype forgings.

141

Besides, these grain size measurements at Ohio University were made only on two forgings.

The physical simulation of infrared processing of aluminum alloy forging was conducted with coupons with almost same amount of deformation since the coupons were obtained from extruded bars which did not have variation in deformation in the longitudinal region. Hence, the effect of changes in thermal processing could be expected to be free from the effect of degree of residual stresses. This argument agrees well with the fact that well defined trends could be seen in the measured degree of solutionizing in the laboratory physical simulation of the solution treatment with rapid heating.

9.2.3 Improvement of Mechanical Properties

These simulations showed that it is possible to obtain almost full solutionizing in these alloys with rapid heating with shorter solution treatment times. This observation is consistent with typical solutionizing of second phases. Accordingly, the results showed that the mechanical properties of the prototype forgings processed with rapid infrared heating were equal to or better than that of conventionally processed forgings made of the corresponding aluminum alloys. The rapid infrared thermal processing improved the hardness of AA 6061 forgings compared to the conventional processing method. The macrograph of the AA 6061 preforms showed existence of elongated grains resulted from extrusion. This means AA 6061 forging prototypes had residual strain in them since there had not been any recrystallization. It is speculated that the residual stress in combination with the rapid heating improved the solution treatment in AA 6061. On the

142 other hand, AA 2618 had recrystallized and hence the amount of residual work present during solution treatment was less. Therefore the rapid heating process could not create significant improvement in hardness over the conventional heating method.

These results match the findings reported from the work of (Kervick, et al. 2006).

DSC studies also indicate that at a higher heating rate (20 ºC/min), the dissolution of second phases is faster than that at lower heating rate (5 ºC/min) for the AA 2618 and AA

6061.

The high cycle fatigue behavior of AA 2618 prototype forgings seemed to be slightly better than that of same-alloy prototype forgings processed with conventional heating methods. Better distribution of intermetallics particles may be the reason for the slight enhancement in the fatigue properties.

9.3 Conclusions

Based on this research and discussion above, the following conclusions were obtained.

The static mechanical properties of AA 2618 forgings processed with conventional heating are similar to that of AA 2618 forgings processed with infrared heating. These mechanical properties include tensile strength, yield strength, elongation before failure and hardness.

The results of high cycle fatigue testing of AA 2618 forgings seems to indicate that thermal processing with infrared heating slightly improves the high cycle fatigue strength.

The improvement is more evident in test regions with lower stresses. However, the results of low cycle fatigue tests indicate that the low cycle fatigue properties of AA

143

2618 forging processed with infrared heating and conventionally processed AA 2618 forgings are equivalent.

The tensile properties of AA 6061 forging prototypes processed with rapid infrared heating seemed to be equivalent to that of conventionally processed AA 6061 forging prototypes. All the measured mechanical properties were greater than the minimum standard values specified for the forgings. Only the percentage elongation values were slightly lower than the minimum value specified for AA 6061 forgings and this is attributed to the existence of elongated grains. The hardness of infrared heat treated AA 6061 forgings is about 6HRB greater than that processed with conventionally processed AA 6061 forgings.

It has been found that thermal stresses arising due to high heating rates improves the solutionizing process. In addition, it is thought that existence of right amount of residual work is necessary for the rapid heating to have a better effect on solution hardening effect on the aluminum alloy forgings.

The research has raised some interesting questions to explore. Based on these, recommendations for future work are presented in the next section.

9.4 Recommendation for Future Work

The results of the current work indicate that the application of rapid heating for the processing of AA 2618 and AA 6061 forgings can produce forgings with mechanical properties equal to or higher than that of conventionally processed forgings and at reduced energy consumption. While the results of the current work provided positive evidence to that, it raises the following issues to be investigated.

144

In-situ measurement of kinetics of the second phase during rapid infrared thermal processing can be measured. The amount of second phase particles affects the conductivity of the alloy being considered. While the alloy coupon is being solution treated at different heating rates, the resistivity of the coupon can be dynamically measured and the measurement can be converted to conductivity. The comparison of variation of the conductivity of the coupons at different heating rates can be related to the solutionization of the second phases in the alloy. This can be an inexpensive and fast test to understand the solution treatment behavior of an alloy and the data can be used for determining optimum solution treatment parameters for thermal processing with rapid heating rate for the alloy.

Another method for in-situ observation of second phase dissolution in the alloys is

X-Ray diffraction (Elmer, Palmer and Specht 2006). While a very small coupon is being heated in vacuum, X-ray diffraction method can be used for observing the second phase dissolution. This method can also be used to gain more data for future applications.

Application of the infrared heating method for other aluminum alloy forgings should be fully investigated. In addition, once the above-mentioned experimental procedure is perfected, the possibility of the application rapid heating methods for the thermal processing of superalloys can also be explored. The above-mentioned method does not require large amount of alloy sample and hence this method can be applied to expensive superalloys.

145

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