Abstract a Comparative Study of 2024-T3 and 7075-T6

ABSTRACT

A COMPARATIVE STUDY OF 2024-T3 AND 7075-T6 ALUMINUM ALLOYS FRICTION STIR WELDED WITH BOBBIN AND CONVENTIONAL TOOLS

by Paul Aaron Goetze

Aluminum alloys 2024-T3 and 7075-T6 were joined by friction stir welding using bobbin and conventional tools types. The effectiveness of placing each alloy on the advancing and retreating side within the bobbin and conventional tool configurations was investigated, and comparisons were made between the welds of the two tool types. Microstructural imaging, micro-hardness mapping, tensile testing, and computational modeling were used to evaluate the weld quality of each tool and material configuration. Temperature data and simulation profiles indicated that the temperature distribution from both tools favored the advancing side with the bobbin tool reaching higher temperatures than the conventional tool. Higher mechanical properties were reported for conventional tool welds than those performed with the bobbin tool, and material placement affected the weld performance in the conventional tool configurations. All tensile specimens fractured on the 2024 side of the weld and specimens joined with the conventional tool were more consistent than those with the bobbin tool. Differential scanning calorimetry identified the precipitation behavior of the alloys which correlated well to the mechanical properties of the welds. Optical microscopy and EBSD analysis highlighted advanced stirring patterns through the weld thickness. Identical equiaxed grain structures observed in the stir zone of both alloys suggested complete recrystallization.

A COMPARATIVE STUDY OF 2024-T3 AND 7075-T6 ALUMINUM ALLOYS FRICTION STIR WELDED WITH BOBBIN AND CONVENTIONAL TOOLS

A Thesis

Submitted to the

Faculty of Miami University

in partial fulfillment of

the requirements for the degree of

Master of Science

Paul Aaron Goetze

Miami University

Oxford, Ohio

2019

Advisor: Dr. Carter Hamilton

Reader: Dr. Giancarlo Corti

Reader: Dr. Fazeel Khan

This Thesis titled

A COMPARATIVE STUDY OF 2024-T3 AND 7075-T6 ALUMINUM ALLOYS FRICTION STIR WELDED WITH BOBBIN AND CONVENTIONAL TOOLS

Paul Aaron Goetze

has been approved for publication by

The College of Engineering and Computing

and

Department of Mechanical and Manufacturing Engineering

______Dr. Carter Hamilton

______Dr. Giancarlo Corti

______Dr. Fazeel Khan

Table of Contents 1. INTRODUCTION ...... 1 2. LITERATURE SURVEY ...... 3

2.1 INFLUENCE OF TOOL GEOMETRY ...... 3 2.1.1 Conventional Tooling Geometries ...... 3 2.1.2 Bobbin Tooling Geometries ...... 4 2.2 INFLUENCE OF PROCESSING TEMPERATURE ...... 4 2.2.1 Precipitation Behavior of 2024 ...... 5 2.2.2 Precipitation Behavior of 7075 ...... 5 2.2.3 Heat Input, Process Temperature and Material Flow ...... 5 2.3 WELD CHARACTERIZATION ...... 8 2.3.1 Microscopy Methods ...... 8 2.3.2 Mechanical Methods ...... 9 3. EXPERIMENTAL PROCEDURE ...... 12 4. COMPUTATIONAL MODELING APPROACH ...... 14

4.1 GEOMETRY, MESHING, GENERAL PHYSICS, AND BOUNDARY CONDITIONS ...... 14 4.1.1 Boundary Conditions for Heat Transfer ...... 16 4.1.2 Boundary Conditions for the Flow Capable Region ...... 17 4.2 MATERIAL PROPERTIES ...... 18 4.3 MODEL VERIFICATION ...... 19 5. RESULTS AND DISCUSSION ...... 22

5.1 OPTICAL MICROSCOPY ...... 22 5.2 ELECTRON BACKSCATTER DIFFRACTION ...... 24 5.3 MICRO-HARDNESS DISTRIBUTIONS ...... 24 5.4 TEMPERATURE ANALYSIS ...... 26 5.4.1 Base Material Calorimetry ...... 26 5.4.2 Simulation Results ...... 27 5.4.3 Time at Temperature Evaluation ...... 34 5.5 TENSILE TESTING ...... 36 6. CONCLUSIONS AND FUTURE WORK ...... 38 7. APPENDICES ...... 40

7.1 APPENDIX A – TENSILE SPECIMEN PREPARATION ...... 40 7.2 APPENDIX B – FRACTURE TOUGHNESS SPECIMEN PREPARATION...... 42 7.3 APPENDIX C – TEMPERATURE PROFILE EVALUATIONS OF WELDING CONFIGURATIONS ...... 46 7.3.1 AS-7075B Configuration ...... 46 7.3.2 AS-2024B Configuration ...... 46 7.3.3 AS-7075C Configuration ...... 46 7.3.4 AS-2024C Configuration ...... 47 7.4 APPENDIX D – SEM FRACTURE SURFACE MICROGRAPHS FOR ALL WELDING CONFIGURATIONS ...... 48 7.5 APPENDIX E – EXAMPLE WELDING ZONE DIAGRAM ...... 49 8. REFERENCES ...... 51

iii

List of Tables

Table 1 Summary of experimental configurations and welding trial designations...... 12 Table 2 Material constants used to calculate the dynamic viscosity of the material in the flow capable region of the simulation...... 19 Table 3 Tensile testing results for all material and tool configurations...... 37

List of Figures

Figure 1 Cross-section of FSW stir and surrounding zones...... 1 Figure 2 Representations of geometries for: a the conventional tool, b the bobbin tool...... 3 Figure 3 Conventional tool simulation geometry and mesh...... 15 Figure 4 Bobbin tool simulation geometry and mesh...... 15 Figure 5 The steady-state temperature comparison of conventional tool experiment and simulation. . 19 Figure 6 Temperature as a function of distance away from the bobbin tool. A comparative plot between the simulation and experimental results...... 20 Figure 7 Optical Micrograph for, a the AS-2024B configuration and, b the AS-7075B configuration...... 22 Figure 8 Optical Micrographs for, a the AS-2024C configuration and, b the AS-7075C configuration ..... 23 Figure 9 Results of, a the SEM EBSD scan and, b the Mackenzie plot for the SZ in AS-7075C...... 24 Figure 10 Vickers Hardness Profiles for, a the AS-2024B configuration and, b the AS-7075B configuration...... 25 Figure 11 Vickers Hardness Profiles for, a the AS-2024C configuration ad b the AS-7075C configuration...... 26 Figure 12 Calorimetric data for, a 2024-T6 and, b 7075-T6...... 27 Figure 13 Bobbin tool configuration temperature comparison of AS-2024B and AS-7075B...... 28 Figure 14 AS-7075B Temperature-hardness comparison plot ...... 29 Figure 15 AS-2024B Temperature-hardness comparison plot...... 30 Figure 16 Comparison of conventional tool temperatures for AS-2024C and AS-7075C...... 31 Figure 17 AS-7075C Temperature-hardness comparison plot...... 32 Figure 18 AS-2024C Temperature-hardness comparison plot...... 33 Figure 19 Hardness Minima locations on the conventional tool optical micrographs...... 34 Figure 20 Time at temperature comparison between the bobbin and conventional configurations of AS- 2024. The x-axis represents the time of tool approach and departure from the point of interest. 35 Figure 21 Failure locations from tensile testing of all welding configurations...... 36 Figure 22 SEM fracture surfaces for: a the AS-2024C, b the AS-2024B configurations...... 36 Figure 23 Sub-size sheet-type tensile specimen...... 41 Figure 24 Sub-size sheet-type tensile specimen machining assembly...... 42 Figure 25 Compact tension-type fracture toughness specimen...... 43 Figure 26 Machining fixture assembly for C(T) specimens...... 44 Figure 27 Drawing of clevis for testing C(T) specimens...... 45 Figure 28 SEM imaged fracture surfaces of the a AS-2024C, b AS-2024B, c AS-7075C and, d AS-7075B configurations at 5 keV...... 48 Figure 29 Wide field SEM imaged fracture surfaces of the a AS-2024C, b AS-2024B, c AS-7075C, d AS- 7075B configurations at 5 keV...... 49 Figure 30 Welding zone demarcation diagram for the AS-2024B configuration...... 49 Figure 31 Composite images for: a the AS-2024B and, b the AS-7075B configurations...... 50 Figure 32 Composite images for: a the AS-2024C and, b the AS-7075C configurations...... 50

Dedication

I dedicate this work to my family.

Acknowledgements

First, I am profoundly grateful to my advisor, Dr. Carter Hamilton, for his advice, support, and direction for the duration of the thesis process. Special thanks to Dr. Anthony Reynolds and Mr. Dan Wilhelm at the University of South Carolina for their expertise and facilities in the production of all welding configurations used in this study. Additionally, Dr. Mateusz Kopyściański and the Department of Metal Engineering and Industrial Computer Science at the AGH University of Science and Technology were very generous in their assistance and guidance during the examination phase of the study. My sincere thanks to Dr. Giancarlo Corti and Dr. Fazeel Khan for their engineering expertise, general availability, and service on my examination committee. I am grateful to Karl Reiff for his trust, mentorship and training throughout my time in the mechanical engineering laboratory. I would like to thank Dr. Richard Edelmann, Matt Duley, and Brittany Cymes at the Center for Advanced Microscopy and Imaging within Miami University for their time and advice in various microscopy techniques. Finally, I would like to thank the faculty and staff of the Mechanical and Manufacturing Engineering Department at Miami University for providing me support and resources throughout my undergraduate and graduate studies.

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1. Introduction The process of friction stir welding (FSW) is a solid-state metal joining technique first developed in the early 1990s. It was primarily intended for the joining of aluminum alloys and the process has become very cost-effective in several transportation industries. Friction stir welding, friction stir processing, and friction stir spot welding are beginning to displace traditional welding methods because they feature lower heat inputs and longevity of tooling. In the case of welding aluminum, a hardened steel tool may last thousands of meters. The tool enters the material at controlled rpm and traverse speed, producing heat through friction and plastic deformation of the material. This heat allows the material to flow around the tool creating a mixed zone along the centerline of the weld. Because the material is stirred with a unidirectional rotating tool, there are distinct regions around the weld that exhibit various material properties. The heat input and physical deformation of the material produce the weld zone where stirring occurs (SZ), a thermo-mechanically affected transition zone (TMAZ) where there are both thermal and mechanical changes, and a heat affected zone (HAZ) where the dissipated heat alters the properties of the base material (BM). Due to the distinct material flow associated with tool rotation and advancement on either side of the weld, the tool produces an advancing (AS) and retreating (RS) side which have different characteristics. A schematic of the weld zone and surrounding regions is shown in Figure Figure 1 Cross-section of FSW stir and surrounding 1. zones. The low heat input of friction stir welding makes it an attractive method for joining high strength aluminum alloys such as those in the 2XXX and 7XXX series that are very difficult to join with traditional fusion welding. Fusion welding relies on the ability to melt the base material, allowing it to pool and solidify across the weld line. The spike in temperature and quick cooling can cause crack formation and other adverse properties in the high strength aluminum alloys like 2024 and 7075. Conversely, FSW can be used to join these alloys because the heat generated in the process does not reach the solidus temperature and prevents excessive phase changes in the material. Much prior work has been done to characterize the properties of these alloys joined with FSW. Literature by Mahoney, Jones, and Mishra [1]-[3] outline these topics well. There are two primary tools used in the process: a conventional tool that requires downward pressure to facilitate stirring and a self-supporting bobbin tool that stirs from the upper and lower surfaces. Most of the studies regarding the joining of aluminum have been performed with conventional tool geometries. The tool and its features have been summarized by Rai et al. [4]. However, the newer bobbin tool geometry has made an impact in the FSW community because of its unique weld stirring characteristics. Threadgill et al. [5] introduce the bobbin tool, but few studies have examined the characteristics of welds produced by the tool. Furthermore, dissimilar alloy joining is a developing interest within the field of friction stir welding, for example 2024 and 7075. This technique is potentially useful for creating structures that leverage the properties of multiple alloys. Da Silva, Cavaliere, and Khodir and Shibayanagi [6]-[8] provide useful bases for this topic. However, there is a need to examine the effects of certain variables on the process because of the asymmetric nature of the produced joints. 1

Specifically, there is limited research comparing the conventional tool geometries, bobbin tool geometries and the effects of material placement on either the advancing or retreating side of the weld line. The objective of this proposed study is to examine the properties of friction stir welded butt joints of 2024-T3 and 7075-T6, specifically comparing the effects of the bobbin and conventional tool geometries and material placement on the advancing or retreating side of the weld.

2. Literature Survey 2.1 Influence of Tool Geometry Two principal types of tools are used in friction stir welding: the conventional and bobbin (self-reacting) tools. Conventional tools consist of a single shoulder and pin with varying shoulder scrolling and pin threading geometries. These tools engage the workpiece through the top surface and rely on downward force and rotation to cause plastic flow of the material. Bobbin tools are designed with upper and lower shoulders that are separated by a pin that is usually threaded and tapered. The bobbin tool enters the material from the side because it cannot plunge and the two shoulders maintain the vertical position of the tool during welding. Simplified versions of these tool geometries are illustrated in Figure 2.

Figure 2 Representations of geometries for: a the conventional tool, b the bobbin tool. 2.1.1 Conventional Tooling Geometries There is a significant body of work on the joining of aluminum alloys with conventional tools and the effects of process parameters on weld performance. The tool geometry produces a slightly tapered stir zone with distinct shoulder and pin dominated regions. Heat input from the conventional tool begins with high downward pressure on the tool, pushing it into the material while it rotates. When the shoulder contacts the top surface, the friction created produces enough heat to start material flow. The material is taken from the AS to the RS where it is forced down toward the center of the weld by the pin. This process repeats, maintained by the downward pressure on the tool, forcing the temperature in the SZ to reach just below the solidus temperature of the alloy. Hamilton et al. [9] provide an in-depth explanation of this process. A review by Rai et al. [4] gives a thorough basis for conventional tool geometries and materials for various workpiece configurations. Kalemba et al. [10] and Ahmed et al. [11] examine the grain structures produced around the stir zone and note appreciable grain orientation differences between the advancing and retreating sides of conventional tool welds. The conventional geometry has been utilized since the inception of FSW and sets the standard for weld quality and tool performance.

2.1.2 Bobbin Tooling Geometries In contrast to experiments performed with the conventional geometry, there are relatively few studies that examine aluminum welds produced with the bobbin tool. However, the geometry produces a unique microstructure and has certain advantages over the conventional tool. Rather than producing a tapered cross-section, Threadgill et al. [5] describe the bobbin tool weld zone as being hourglass shaped. Similar stir patterns to the conventional tool are observable, but the two shoulder dominated regions affect the stir zone more. In the same work, Threadgill et al. compare the tool forces required for bobbin and conventional geometries. Overall, conventional tools require much higher downward forces to keep the tool in plane with the workpiece than the bobbin tool. The normal forces on the upper and lower shoulders of the bobbin tool act opposite to each other and reduce the plunging force required. Threadgill et al. indicate that 14 kN of down-force is applied to the bobbin tool whereas 62 kN is applied to the conventional tool. This reduction in force eliminates the need for a thick backing plate standard in processing with a conventional tool. Although the bobbin tool does not require significant downward force, the heat input and stirring action are created by the thermal expansion of the aluminum between the two shoulders of the tool. The two shoulders act together to push material toward the center of the weld. To produce good results, the bobbin tool may utilize a slower rotation speed than the conventional tool. This speed reduction may affect the performance characteristics of the weld as discussed in section 2.2. Studies have compared the bobbin and conventional tool [12], discussed process parameters and weld defects [13], [14], and examined the effects of stir zone bulge on weld performance [15]. These give a strong basis for understanding how the bobbin tool affects the welding process and prove its usefulness. However, comparative studies examining the difference in characteristics between the bobbin and conventional tool welds are still necessary. 2.2 Influence of Processing Temperature The heat generated by friction stir welding subjects the material to elevated temperatures that conduct into the base material outside the joint line. For alloys that are age-hardenable and strengthened by secondary phases, like the 2xxx and 7xxx aluminum alloys, the welding temperature in relation to the phase transformation temperatures, i.e., dissolution and precipitation, is critical to the final properties of the weld. For example, Reynolds [16] states that if the welding temperature does not reach the solution heat treatment temperature (TSHT) for an age-hardenable alloy, the secondary phases in the matrix coarsen and produce a low material hardness. In contrast, if the peak temperature in the weld nugget is above the TSHT, producing a single-phase solid solution, then recovery of base mechanical properties is possible upon cooling. Further, a weld left in the partial solution heat treated condition can be hardened by the Guinier- Preston (GP) zones found in a full solution heat treated alloy, but at the same time, coarsened precipitate phases are still present, causing a post-weld aged hardness between that produced by the overaged and single-phase solid solution conditions. Since this work aims to examine the behavior of 2XXX and 7XXX series heat treatable alloys, the focus of the literature survey is limited to these two alloys. Several works discuss the correlation between the hardness profiles measured and the precipitation behavior of the aluminum over the temperature distribution experienced in FSW.

2.2.1 Precipitation Behavior of 2024 For naturally aged aluminum alloys containing copper and magnesium, like 2024, several precipitate phases can be produced based on the peak temperature of the weld. As the temperature increases, the precipitates leave the Guinier-Preston (GP) phase formed in quenching and begin to grow into the S’ transition phase which is coherent in the matrix and the alloy retains some of its hardness. However, the alloy soon reaches the S transition phase where the precipitates grow coarser and become incoherent from the matrix. Hardness values measured when the material peaks at this temperature will be the lowest [17]. For 2024, the TSHT is 493 °C and it is likely that the HAZs of 2024 joints produced with FSW contain S transition phases. In a study by Jones et al. [2], a plot of hardness as a function of distance from the weld centerline is presented that contains a local maximum surrounded by two minima. A 2024 weld is examined for a relationship between microhardness values across the weld centerline and the material microstructure. According to Jones et al. [2], the changes in hardness are due to the dissolution and coarsening of precipitates as the peak temperature decreases with distance away from the weld nugget. The significant dip in hardness closest to the weld line is due to the presence of overaged S phases in the matrix. The sharp rise in hardness immediately following the dip is caused by S’ phases and the second dip farthest away from the weld line is possibly due to the dissolution of the GP zones into the solid solution. 2.2.2 Precipitation Behavior of 7075 For artificially aged aluminum alloys containing zinc and magnesium, such as 7075, a somewhat similar set of precipitation phases is possible as temperature increases from room temperature. First, distributed GP phases begin to coarsen and form the η’ transition phase. Unlike the S’ phase, η’ is not fully coherent with the matrix and may not produce the same increase in hardness for 7075. At higher temperatures, below the TSHT, the precipitates coarsen and form the η phase [17]. Outside of formed GP phases, the η and η’ phases produce reduced hardness values in 7075. The TSHT temperature for 7075 is between 466 °C and 482 °C and it is likely that the HAZs of 7075 welds reaches a peak temperature below this point, significantly coarsening precipitates and reducing hardness. Although 7075 friction stir welds show a decrease in hardness outside of the weld nugget, the η’ and η precipitate phases do not produce a complex hardness profile like the S’ and S phases in 2024. Chao et al. [18] show that the welded 7075 produces a single hardness dip in the heat affected zone likely from the coarsening of precipitates in the matrix and suggests that the highest strain occurs in the area with the lowest hardness of the 7075 welds. Mahoney et al. [1] also show that tensile failure in the welded sample occurs in the area with the highest strain. The study indicates that the temperature distribution around the failure zone is well below the TSHT for 7075. Together, Chao et al. and Mahoney et al. prove that a low-temperature region around the weld promotes precipitate coarsening which causes high strain and eventual localized failure in tension testing. 2.2.3 Heat Input, Process Temperature and Material Flow Several approaches have been taken to estimate the heat input during friction stir welding of metallic materials in order to model the temperature distribution and flow around the weld, tool, and base material. Torque-based and shear layer approaches may both be applied to the conventional and bobbin tool configurations. Alternatively, a combination of the thermal modeling techniques has been used to eliminate some of the difficult input parameters while maintaining a more realistic solution.

A torque-based approach was taken by Khandkar et al. [19] to describe the heat input into a 6061-T651 friction stir weld with a conventional tool by the sum of the torque of the shoulder- workpiece interface, the vertical pin-workpiece interface, and the bottom of the pin-workpiece interface. This total torque was related to the average power input by multiplying the total torque by the angular velocity of the tool. The torque equations follow as:

푟표 푇푠ℎ표푢푙푑푒푟 = ∫ (휏푟)(2휋푟) 푑푟, (1) 푟𝑖

푟 푇 = 𝑖(휏푟)(2휋푟) 푑푟, (2) 푃𝑖푛퐵표푡푡표푚 ∫0

푇푃𝑖푛푆푢푟푓푎푐푒 = (휏푟𝑖)2휋푟𝑖ℎ, (3)

푃푎푣 = 푇푡표푡푎푙휔, (4) where τ is the shear stress, r is the radius of the involved geometry, h is the length of the pin, and ω is the angular velocity of the tool. The shear stress was found by utilizing the above equations and the measured torque outputs from the machine. This model was applied by Hamilton et al. [20] to characterize the material flow and temperature distributions of a 7136-T76511 weld. The equations utilized in this model describe a conventional tool geometry in a butt weld configuration. The total torque of the shoulder, side surfaces of the pin, and pin bottom was expressed as: 2 푟0 푟𝑖 (5) 푇푡표푡푎푙 = 2휇퐹 ( + 2 ℎ), 3 푟0 where, r0 is the radius of the tool shoulder, ri is the pin radius, h is the pin height, and μ is the coefficient of friction between the tool and the workpiece. Subsequently, the energy input to the weld was expressed with the following equation where ω is the angular velocity of the tool and vw is the velocity of the weld: 휔 퐸푙 = 푇푡표푡푎푙 . (6) 푣푤 Then, starting with the governing heat equation for a moving heat source, Hamilton et al. [20] developed a torque-based expression characterizing the heat flux generated by the tool- workpiece interaction. This equation is presented below:

푞 = 훿휇푃푁휔푟 (7) where, δ represents the friction slip factor, μ is the coefficient of friction, PN is the normal pressure along the tool face, and r is the radial distance away from the tool centerline. The slip factor was calculated utilizing the following formula that takes the ratio of energy input to the effective energy input for a weld at the solidus temperature of the material. This relationship was developed previously by Hamilton et al. [21] as: 퐸 훿 = exp [− 푙 ], (8) 퐸푙푀푎푥 where El is the effective energy per unit length of weld and ElMax is the energy of the material when the welding temperature is equal to the workpiece solidus temperature. Effectively, the slip factor is determined by the energy input to the weld and its ratio to the maximum energy input corresponding to bringing the weld to the solidus temperature of the material. As the heat

6 increases, the friction heat transfer becomes less effective because the interface between the tool and the workpiece softens and sliding friction dominates over sticky friction. This is evident in the expression because when El is low, the factor is close to 1 and as El approaches ElMax, the factor approaches 0.37, reducing the overall heat transfer term. With this torque-based approach, the pressure of the shoulder and the bottom of the pin dominate the model and the pressure normal to the sides of the pin is comparatively negligible. Therefore, the average heat flux across the shoulder and bottom of the pin provide the necessary components of a heat source to the model. These area averaged equations are presented below as: 2 훿휇푃 휔(푟3 − 푟3) 푁 0 𝑖 (9) 푞푃𝑖푛퐵표푡푡표푚 = 2 2 3 푟0 − 푟𝑖 and 2 푞 = 훿휇푃 휔푟 . (10) 푆ℎ표푢푙푑푒푟 3 푁 𝑖 The torque-based approach was proved to be accurate by Khandkar et al. [19] and Hamilton et al. [20]–[22] and is a reliable method for the thermal modeling of conventional tool friction stir welds. However, the method requires inputs that are difficult to accurately predict like the coefficient of friction, μ. Another type of thermal model uses the shear layer methodology (SLM) to represent the tool and workpiece interaction. This approach eliminates the need for the coefficient of friction and any machine force inputs because it is based on a thermal pseudo-mechanical (TPM) heat source. The TPM method, developed by Schmidt and Hattel [23], relies on the friction shear stress at the welding interface and yield shear stress of the material to characterize the model. It begins with the understanding that heat is generated between the tool and workpiece, principally because of friction. Due to the inefficiencies of continued frictional heat transfer as the heat increases near the solidus temperature, a surface contact state variable was introduced to account for this. Hilgert et al. [24] expressed equation 11 for the state variable, δ1: 휔푚푎푡푟𝑖푥 훿1 = (11) 휔푡표표푙 where, ωmatrix is the angular velocity of the material matrix around the tool and ωtool is the angular velocity of the tool. When δ1 approaches 1, sticking friction dominates because the rotational velocity of the matrix and tool coincide. As δ1 decreases, the interaction becomes more dominated by sliding friction determined by the discrepancy between the matrix and tool rotations. Schmidt and Hattel conclude that even for a small value of the state variable, δ1, the frictional shear stress must equal the yield shear stress of the material because even the smallest amount of sticking friction causes plastic deformation in the material. The total heat flux was defined by the shear rate, shear stresses, and tool rotational velocity. With the incorporation of the state variable relationship and Coulomb’s frictional law, the total heat flux was simplified to equation 12 by Schmidt and Hattel [23] and was also shown to be the same by Hilgert et al. for case of the bobbin tool [24]:

푞 = 휔푟휏푦𝑖푒푙푑, (12) where ω is the angular velocity of the tool, r is the shoulder radius, and τyield is the yield shear stress of the material. This model simplifies the input parameters significantly compared to the torque-based approach and is easily adaptable for a variety of tool and workpiece configurations. In a recent work, Hamilton et al. [25] prove that the shear layer and the torque-based approaches can yield extremely comparable results even with such different base methodologies. 7

The work highlights the advantages and disadvantages of each of the methods, explicitly pointing out the difficulty of finding input parameter values in the torque-based method and the fact that the shear layer approach limits the heat generation to the front of the tool as disadvantages. The work aims to modify the methodology of the shear layer approach for the bobbin tool geometry to account for the entirety of the tool/workpiece. The original SLM approach assumes that the heat generation between the tool and workpiece is 100% efficient, causing a discrepancy between the two methodologies. To correct this, an efficiency factor was incorporated into the general heat flux equation detailed by Schmidt and Hattel [23], producing the following result:

푞 = 휔푟(휂푝휆휏푦𝑖푒푙푑 + 휂푓(1 − 휆)휏푓푟𝑖푐푡𝑖표푛) (13) where, ηp and ηf are efficiency factors associated with the plastic deformation and frictional heat generation in the system, and λ is the ratio between the material and tool rotational velocity. This equation could be then simplified based on the understanding that even for small values of λ the frictional shear stress will be equal to the yield shear stress of the material. Experimentation by Heurtier et al. [26] indicates that λ is approximately 0.01. With such a small value of λ, the plastic deformation term of the heat flux equation becomes negligible. The result of these simplifications is presented as equation 14:

푞 = 휂휔푟휏푦𝑖푒푙푑. (14) Hamilton et al. [25] estimate a reasonable value for the heat transfer efficiency by setting the ratio of the heat generation for the torque based and modified shear layer approaches equal to 1. This produced an efficiency factor, η, of 0.5. Effectively, Hamilton et al. [25] modified the SLM approach to account for all surface interactions by negating the assumption that the heat transfer between the surfaces is 100% efficient and equating the heat flux to that produced by the torque-based modeling approach. This resulted in a modified SLM methodology that maintains the realism of the torque-based approach but does not require the inclusion of the input parameters of the torque-based model. 2.3 Weld Characterization Ultimately, friction stir welds are characterized through microscopy techniques and mechanical testing procedures. Separately, each analysis method highlights a factor of quality for the joint, but a full picture of the SZ, TMAZ, HAZ, strength, and toughness characteristics and how they interact is clear when the methods are combined. Microscopy can be utilized to identify the SZ grain structure, TMAZ location, and presence of defects that have an effect on the tensile and fracture toughness testing results in a study. Meanwhile, micro-hardness testing can be used to confirm the failure locations by identifying the weakened areas of the weld due to the sharp thermal gradient caused by tool rotation and translation. These methods have become standard techniques for evaluating aluminum alloy friction stir welds. A survey of the application of the methods and results from similar alloy FSW form the foundation for dissimilar alloy welding characterization. 2.3.1 Microscopy Methods In similar alloy welding, the observation of FSW cross-sections provides useful information about possible weld defects, stirring patterns and microstructural changes across the weld line. Optical (OM), Scanning electron (SEM), and transmission electron (TEM) microscopy have

8 been used by many authors to analyze the entire microstructure. Standard practices of anodizing, etching, and polishing are employed to reveal details and grain boundaries in the specimens. Optical micrographs are useful for heat treatable alloys by indicating grain structures throughout the cross-section and emphasizing defects. Shah and Badheka [27] provide optical micrographs of 7075-T6 that highlight the TMAZ and SZ boundaries on both the AS and RS. The boundary is well defined on the AS but much more gradual on the RS visible due to the stirring patterns and higher heat on the AS. Khodir et al. [28] examined weld defects and microstructural features of 2024-T3 also using micrographs. More direct analysis of grain size and orientation can be completed through EBSD mapping with SEM microscopes. Kalemba et al. [10] performed a quantitative analysis of grain size and orientation for the AS and RS of a 7136-T76 weld. The study highlights the asymmetrical nature of the FSW process and shows the more abrupt transition between SZ and TMAZ on the AS compared to the RS. The observation of material flow patterns, weld defects, and microstructural changes are critical to the full characterization of dissimilar alloy welding. Several works have shown that contrasting stirring patterns are obtained when material configurations are compared. For example, Guo et al. [29], Simar et al. [30], and Kalemba-Rec [31] indicate this phenomenon for several alloy configurations by suggesting that the stirring patterns may be material dependent. Research in FSW frequently utilizes optical microscopy techniques to observe and evaluate these patterns using an anodizing reagent and polarized light. Proper selection of the reagent allows the dissimilar alloys to appear differently under the microscope which helps the observer to identify important features. Some works include SEM and TEM micrographs to highlight EDS spectroscopy and precipitate changes across the welded cross-section. These results are used to validate observed stirring patterns in optical micrographs and phenomena recorded in mechanical testing regimes. 2.3.2 Mechanical Methods Hardness Mapping As discussed in relation to the precipitation behavior during welding, hardness mapping of weld cross-sections is a valuable tool for understanding the post-weld makeup of the weld nugget, thermo-mechanical, and thermally affected surrounding zones. da Silva et al. [6] studied dissimilar welding using a conventional tool with 7075-T6 on the advancing side and 2024-T3 on the retreating side. Overall, the hardness map shows the general hardness trends of 7075 and 2024 as the distance increases from the weld center. Hardness dips near the weld nugget boundaries are seen and there is a slight hardness recovery in the stir zone, which likely due to dynamic recrystallization. The work by Cavaliere [7] corroborates the hardness results found by da Silva et al., showing that there is a general hardness interface between the 2024 and 7075 in the stir zone. Again, dynamic recrystallization is observed in the stir zone and hardness minima are seen as measurements reach the HAZ on either side of the weld line. However, neither work considers the effect of alloy placement on the advancing or retreating side. This variable is examined by Khodir and Shibayanagi [8], but hardness profiles for both material configurations are not provided. Park et al. [32] show that hardness profiles differ between material configurations for 5052-H32 (non-heat treatable) and 6061-T6 (heat treatable) welds because the material mixing and thermal properties in the stir zone are affected by material placement. Although Park et al. make valuable comparisons between the hardness profiles of 5052-6061 welds, the fact that 5052-H32 is non-heat treatable makes the comparison to hardness mapping

9 of heat treatable 7075-2024 welds difficult. Thus, further work of examining the hardness profiles produced by 7075-2024 welds in both material configurations is needed. Tensile Testing Tensile testing of friction stir welded 7075 shows weld strength lower than the base metal properties. In the transverse orientation, Mahoney et al. [1] found that specimens fractured in the heat affected zone, probably on the retreating side of the weld. The yield strength of the sample was approximately 54% and the ultimate tensile strength was 75% of the base metal. The elongation percentages for the as-welded samples were 7.5% and were around half of the base metal properties. Overall, the fracture in the heat affected zone suggests material softening due to heat input during the welding process. The material approached 350 oC in the failure zone causing precipitate coarsening and local weakening of the matrix. Additionally, Mahoney et al. note that the HAZ is wider at the top of the weld due to the tool input and shrinks at the bottom of the weld from heat loss to the backing support. Tensile testing of friction stir welded 2024 indicated similar trends to 7075. Khodir et al. [28] reported tensile fracture data for a variety of welding parameters using a conventional tool. For a welding speed of 1250 min-1 and feed of 50 mm min-1, the transverse tensile specimen fractured in the HAZ on the retreating side of the weld. The fracture occurred at 402 MPa, indicating an 88% joint efficiency from the base metal strength of 455 MPa. Additionally, the yield strength of the welded specimen was close to 300 MPa, compared to a 312 MPa measured value of the base material. Overall, Khodir et al. indicate that failure resulted from coarsened precipitates in a thermally affected area of the welded joint. The reported elongation percentage for the welded 2024 showed a slightly reduced value of 12% compared to 14% for the base metal. In the transverse to weld direction orientation, tensile specimens contain the SZ, TMAZs, and HAZs on each side of the weld. If a defect is present, testing in this orientation can cause failure near that location. With defect free welds, a common failure location is found across multiple studies and alloys. Works by Khodir et al. [28], Hamilton et al. [20], Sato and Kokawa [33], and Liu et al. [34] all indicate failure in the HAZ of the retreating side. The tensile testing results of dissimilar alloy welding are consistent across most works covering mechanical characterization of the process. Instead of specimens fracturing in the HAZ of the RS like similar alloy welds, most tend to fail on the weaker alloy side in the HAZ of the weld. The works by da Silva [6] and Cavaliere [7] both show that 7075-2024 welds failed in the HAZ region of the 2024 side. Khodir and Shibayanagi [8] show that this is independent of material placement for 7075-2024 joints when a defect-free weld is produced. Amancio-Filho et al. [35] and Park et al. [32] show similar results for different aluminum alloy configurations. Fracture Toughness Mishra and Ma [3] present a collection of basic fracture toughness properties for 7XXX series aluminum alloys. In general, it was found that the crack tip opening displacement (CTOD) values were higher than base metal properties. A large CTOD value indicates a greater ductility of the alloy as it undergoes plastic deformation and eventually reaches failure. Mishra and Ma’s [3] review noted the CTOD of 7075 in the weld nugget to be 0.024 mm which was double that of the base material. For 7002, Strombeck et al. [36] reported a value of 0.48 mm in the weld nugget compared to the base metal value of 0.40 mm. These results indicate that the friction stir welding process improved the fracture toughness properties of the 7XXX alloys being welded. In addition to the data for the 7XXX series, Mishra and Ma [3] provide data for fracture toughness of 2024-T3. In the review, Strombeck et al. [36] reports that the weld nugget CTOD

10 value was 0.22mm, slightly less than the 0.30 mm of the base metal. This data indicates that the welding process has the opposite effect on fracture toughness properties for 2024-T3 compared to 7075-T6. The fracture toughness of dissimilar aluminum alloys joined by FSW is a relatively unexplored mechanical characterization. However, standard techniques outlined by the ASTM standard in volume 03.01 [37] and precedents from testing done on similar alloy welds serve as a basis for this process.

3. Experimental Procedure The design of the experiment contains two primary elements. First, performance characteristics are compared between the bobbin and conventional tool geometries when they are used with dissimilar alloys. Second, the effect of the placement of the 2024 and 7075 alloys on the retreating (RS) or advancing side (AS) was assessed. Comparisons can be made between the two tool types on the basis that all four configurations are successfully joined without defects. Tensile testing, micro-hardness maps, and optical microscopy were used to fully characterize each weld trial and compare the effect of tool choice and material placement on weld quality. Table 1 summarizes the experimental configurations and provides identifying nomenclature for each welding trial. Table 1 Summary of experimental configurations and welding trial designations.

Trial Desig. AS Mat. Tool AS-2024B 2024 Bobbin Tool AS-7075B 7075 AS-7075C 7075 Conv. Tool AS-2024C 2024

Aluminum sheets of 2024-T351 and 7075-T6 were selected, conforming to the ASTM B209 standard, at a thickness of 6.35 mm. Each sheet was 101.6 mm wide by 609.6 mm long before processing and the mating surfaces were machined for proper alignment. The welding was done in two stages at the University of South Carolina on an MTS ISTIR machine to ensure the process validity for each material and tool configuration. After the first half of each trial was completed, a three-point bend test was done to test for process defects. Then, each trial was finished using the same or corrected parameters if needed. In practice, a crosshead feed of 127 mm min-1 and a spindle speed of 120 RPM was used for all material configurations with the bobbin (self-reacting) tool. A feed of 203.2 mm min-1 and speed of 300 RPM was used with the conventional tool geometry for all material configurations. The machine operated with a 0- degree tool tilt angle. The bobbin tool used had 17 mm shoulders and a 6.35 mm threaded pin and the conventional tool had a 17.78 mm shoulder diameter with a 4.1 mm scroll pitch and a 6.35 mm threaded pin. Temperature measurements were made in configurations with the bobbin tool by embedding a thermocouple in the base sheet near the weld nugget and the conventional tool had a thermocouple embedded in the center of the tool body. A Zeiss Axio Imager M1 light optical microscope was used to observe the microstructural features of a cross-section from each weld configuration. Each specimen was prepared through standard polishing practice and etched with Barker’s Reagent at a 1.8% concentration of fluoroboric acid (HBF4) in water for 2 min at 20 V DC. For all welding configurations, the advancing side is always presented on the left-hand side of the image. Vickers micro-hardness (HV) testing of the weld cross-sections for each configuration was performed using 1 kg load for 10 s on a Wolpert-Wilson Tukon 2500 machine. For the conventional tool configurations, one measurement row was taken along the mid-thickness line of each cross section. Measurements with a 1 mm spacing interval were taken to a 20 mm distance on each side of the weld centerline. A total of three measurement rows were taken for 12 each bobbin tool configuration: 1 mm below the top edge of the cross-section (upper shoulder), the mid-thickness line, and 1 mm above the bottom edge of the cross-section (lower shoulder). As before, a 1 mm spacing was used between indentations out to 20 mm on each side of the weld centerline. Tensile specimens were machined transverse to the weld direction on a Haas TM-1 CNC milling machine. A sub-size sheet-type specimen design was selected to conserve space on the welded sheets and to ensure the tensile trials could be accomplished on an MTS Exceed E44 with flat, mechanical grips and a load capacity of 26,000 N. The specimen design conformed to the ASTM E8-04 standard and each specimen was roughly machined to basic dimensions using a carbide tool under flood coolant to ensure proper cooling during cutting. The exact dimensions and exhaustive information on preparation are provided in Appendix A. A 7.9375 mm carbide endmill was used to achieve the final dimensions with a finish pass of 0.05 mm the average surface roughness in the gage region for the group of specimens was 0.2 μm. Five specimens were machined for each material configuration and each was engraved with a serial number for identification. The tensile testing was performed after four months of natural aging at a strain rate of 0.012 min-1. An MTS 634.12E-54 extensometer with a gage length of 25.4 mm was used during each trial. All welded material and samples were kept below 0oC after completion of the tensile testing to prevent further natural aging before or during the subsequent hardness experiments. Two-sample t-tests assuming unknown and unequal variance were used to identify any trends in the data following tensile testing. The two-sided hypothesis test type with a significance level, α, of 0.05 was selected. The null hypothesis, h0, for comparing the population mean, pm, of each sample group, was set to pm1 = pm2. The alternative hypothesis, h1, was set to pm1 ≠ pm2. A Zeiss Supra 35 VP FEG Scanning Electron Microscope was used to collect micrographs of the fracture surfaces from tensile samples for each welding configuration. Secondary electron detection at 5keV was used for the imaging. Each specimen was sonicated in 95% ethanol for 20 minutes prior to placement in the microscope and each sample was secured to the stage by double-sided carbon tape. Differential scanning calorimetry of was performed using a PerkinElmer DSC 8500. Base material specimens from both 2024-T3 and 7075-T6 were sections in the form of 4 mm disks. Testing of the samples occurred over a -50 °C to 500 °C heat range at a rate of 20 °C min-1. Fracture toughness specimens were machined with the notch centerlines parallel to the welding direction. A compact tension specimen design conforming to the ASTM E1820-06 standard was chosen with dimensions that kept the entire height of the sample within the shoulder-affected region of the weld. The machining was done on a Haas TM-1 CNC mill with a carbide tool and flood coolant. A 7.9375 mm carbide endmill was used with a 0.05 mm finish pass to reach the specimen’s target outer dimensions. An Excetek W350G wire EDM was selected for machining the notch section and tip of each specimen. Micro EDM wire with a 0.100 mm diameter was selected for these features. Due to unforeseen complications with the EDM machine, fracture toughness testing was not completed for this thesis. Exact dimensions for the specimen, the test clevis design, and information regarding the planned testing is provided in Appendix B.

4. Computational Modeling Approach Models for each welding configuration were designed in COMSOL Multiphysics using the torque-based approach developed by Khandkar et al. [19]. This approach was applied by Hamilton et al. [22, 25] whose methodology is the basis for this model. The purpose of the modeling was to identify the peak temperatures and temperature profiles of the models after steady state configurations were reached. These temperature results were validated by the experimental results and then used to relate the mechanical behavior of the welds to the precipitate behavior discussed in section 2.2. The precipitation behavior of the alloys directly influences the material hardness which affects the tensile results. 4.1 Geometry, Meshing, General Physics, and Boundary Conditions One model was used as the basis for each welding configuration. The geometry creation, meshing techniques, and model physics were the same regardless of material configuration and tool type. Similar to the modeling done by Kalemba-Rec et al. [31], three main geometric domains were created to capture the heat generation and material flow: solid-state regions for the base material and backing plates, a fluid capable region for the material directly interacting with the tool, and the tool geometry. The geometry was split along the centerline to create domains for both materials. The flow capable region extended 1 mm past the radius of the tool shoulder, transverse to the movement direction, to allow for flow around the tool. The interaction between the tool and the material require both a general heat transfer and laminar flow physics. These physics and geometry interaction factors were considered for the building of the mesh. The mesh consisted of free tetrahedral elements that were continuous across the various solids. Non-union contacting faces maintained the mesh continuity by projecting the face mesh from one solid to the matching face of the other solid. The mesh was refined in the flow region around the tool and a much coarser mesh was generated outside of the region of tool interaction. Figure 3 and Figure Figure 4 highlight the geometric regions and meshes constructed for the conventional and bobbin tool configurations, respectively.

Figure 3 Conventional tool simulation geometry and mesh.

Figure 4 Bobbin tool simulation geometry and mesh.

4.1.1 Boundary Conditions for Heat Transfer In general, the heat transfer physics capture the heat generation between the tool interaction occurring in the flow zone and the solid-state regions on either side of the weld line. The frictional slip factor, δ, and the heat generation equation used in this simulation were presented by Kalemba-Rec et al. [31] and Hamilton et al. [22], respectively. The frictional slip factor follows as: 푇 훿 = exp (훽 ( )), (15) 푇푠 where, the β term represents an empirically determined scaling factor that Kalemba-Rec et al. found by including it in a previously calibrated model of similar materials. The value of β was determined to be -4/7. The Ts represents the solidus temperature of the aluminum alloy and T represents the instantaneous temperature calculated in the simulation. The slip factor was calculated for both the AS and RS of the flow region, depending on the placement of 2024 and 7075. For the heat generation in the models, both tool configurations used Equation 7 (2.2.3). However, due to the functional differences between the tools, the manifestation of the normal pressure, PN, in conventional tool simulation is different from that of the bobbin tool simulation. The conventional tool generates heat through downward pressure (2.1.1), so PN was taken as the average down-force exerted by the machine in experimental testing applied over the area of the tool shoulder. The normal pressure created in the bobbin tool simulation was developed by Hamilton et al. [25]. Because the bobbin tool shoulders are fixed and only exert a small amount of pressure on the workpiece initially, the thermal expansion of the aluminum causes the pressure to increase as the aluminum heats up from friction with the pin. This pressure will increase with temperature until the workpiece reaches the temperature dependent yield stress and will remain there until the completion of the weld. Therefore, the simulation takes the minimum quantity between the pressure due to thermal expansion [25] (equation 16) and the temperature dependent yield stress [38] (equation 17) to capture this behavior, shown as: 푃(퐸푥푝푎푛푠𝑖표푛) = 훼퐸(푇 − 푇푟표표푚), (16) and

푇−푇푟표표푚 휎푦(푇) = 휎푦,푟푒푓 (1 − ), (17) 푇푠−푇푟표표푚 where, α is the coefficient of thermal expansion, E is the elastic modulus, Troom is the room temperature (293.17 K) and T is the instantaneous temperature in equation 16. Equation 17 uses the room temperature yield stress, σy, multiplied by a temperature dependent scaling factor to produce the temperature dependent yield stress, σy,ref. For use in COMSOL, the heat generation equation was applied to the surfaces used to generate the heat. The shoulder and pin bottom boundaries were the main contributing surfaces for the conventional tool. The tool shoulders and front side of the pin were incorporated for the bobbin tool. A point-based approach was taken, meaning that the heat generated at each point in the assigned area was calculated during every time step by 2 2 (18) 푞 = 훿휇푃푁휔√푥 + 푦 , where x and y are the coordinates of the specific point with respect to the center of the tool in the plane of the tool-material interaction domain.

4.1.2 Boundary Conditions for the Flow Capable Region The amount of material flow around the tool in the flow capable region of each simulation is determined by a set of boundary conditions at the interaction faces and the dynamic viscosity of the flowing material. In the actual experiments, the tools had scrolls scribed into the shoulder to encourage flow in the shoulder dominated region of the weld and threaded pins to pull material from the surface into the center of the cross-section. These geometries are captured in the simulation through the velocity boundary conditions. The shoulder scroll was recreated using the following velocity field condition: 휔 푥 푢 = 휔푦 − ( ) 푝푠 ( ) 2휋 √푥2+푦2 푉푒푙표푐𝑖푡푦 퐹𝑖푒푙푑 = 휔 푦 , (19) 푣 = −휔푥 − ( ) 푝푠 ( ) 2휋 √푥2+푦2 { 푤 = 0 where ω is the angular velocity of the tool and ps is the pitch of the scroll on the tool shoulder. The u, v, and w variables represent the velocity magnitudes in the x-, y-, and z-directions, respectively. These velocity conditions were used for both the conventional and bobbin tool simulations; the condition was applied to both shoulders of the bobbin tool simulation. The threaded pin was recreated on the pin side boundaries using the following velocity field condition: 푢 = 휔푦 푣 = −휔푥 푉푒푙표푐𝑖푡푦 퐹𝑖푒푙푑 = { 휔 , (20) 푤 = −푝 ( ) 푝 2휋 where pp is the pitch of the threading on the pin. This velocity field was applied to both the conventional and bobbin tool geometries. In the case of the conventional tool, the bottom of the pin also received a velocity field, since material flow would occur from the tool pressure and rotation there. This portion of the tool was not scrolled, so the following simplified velocity field condition was used: 푢 = 휔푦 푉푒푙표푐𝑖푡푦 퐹𝑖푒푙푑 = {푣 = −휔푥, (21) 푤 = 0 The velocity fields of boundaries between the tool shoulders, pin sides, and pin bottom (in the case of the conventional tool) allowed the tool to be modeled as a stationary heat source. The angular velocity of the tool was included in the boundary conditions so that the geometry would not need to actively rotate in time. To model the translational movement of the tool in the material, a velocity condition was assigned to the flow capable region that moved the material past the tool at the traversing speed. Thus, an upstream inlet boundary had the following velocity field condition:

푢 = 푢푤푒푙푑 푉푒푙표푐𝑖푡푦 퐹𝑖푒푙푑 = { 푣 = 0 , (22) 푤 = 0 -1 -1 where uweld is the traverse speed of 127 mm min for the bobbin tool and 203 mm min for the conventional tool in the respective simulations.

The dynamic viscosity of the flowing material in each simulation used to calculate the velocity fields was calculated by 휎 휇 = 푒, 3휀̇ (23) where σe is the flow stress and 휀̇ is the maximum effective strain rate. Due to the differences in tool geometries of the conventional and bobbin tools, the dynamic viscosity calculations contain different estimations of the respective strain rates. The conventional tool strain rate was kept constant through the thickness of the tool and depended on the shoulder radius as presented by Hamilton et al. [22]. It is calculated through the following equation: 휔푟 휀̇ = 푠ℎ표푢푙푑푒푟, 3ℎ (24) where rshoulder is the shoulder radius of the tool and h is the thickness of the workpiece. The strain rate used in the bobbin tool configuration was not held constant for the thickness of the weld. Because the tool shoulders act from both sides of the pin and create an hourglass shaped stir zone, a variable strain rate dependent on the vertical position within the thickness of the workpiece was used. Hamilton et al. [25] estimate that the strain rate decreases linearly based on the radius of the tool as the point of interaction moves from the top surface of the weld ( i.e., the tool shoulder) to the middle of the cross-section (i.e., the pin). This strain rate at the upper shoulder, middle of the pin, and lower shoulder is calculated by 푟푈푆휔 푟푃𝑖푛휔 푟퐿푆휔 휀푈푆̇ = ℎ , 휀푃𝑖푛̇ = ℎ , 휀퐿푆̇ = ℎ , ( )√3 ( )√3 ( )√3 (25) 2 2 2 where rUS is the radius of the upper shoulder, rPin is the radius of the pin, and rLS is the radius of the lower shoulder. Every time step in the simulation, the heat transfer physics calculate the temperatures in the model and pass those values to the laminar flow physics to calculate the velocities fields which are dependent on the temperature sensitive dynamic viscosity calculation. The dynamic viscosity is temperature dependent because the flow stress calculation is dependent on temperature. The flow stress was proposed by Sheppard and Wright [39] and formulated as 1 1 푍 휎 = sinh−1 [( )푛], 푒 훼 퐴 (26) where α, A, and n are material constants (section 4.2) and Z is the Zener-Holloman parameter. The Zener-Holloman parameter captures the effect of temperature on the flow stress through the following equation: 푄 푍 = 휀̇ exp ( ), 푅푇 (27) where Q is the activation energy of the alloy and R is the universal gas constant. The activation energy is also provided in section 4.2 for both 2024 and 7075. 4.2 Material Properties For the thermal calculations in the model, temperature dependent values for thermal conductivity, k, and the specific heat capacity, cp, were taken from Mills [40] for 2024 and Mills [41] for 7075. The Zener-Holloman parameters used to calculate the dynamic viscosities used in the flow calculations of all welding configurations are presented below in Table 2.

Table 2 Material constants used to calculate the dynamic viscosity of the material in the flow capable region of the simulation.

Material Const. 2024 Value 7075 Value Units Reference Jmol- Q 148880 155336 1 A e19.6 e27.54 s-1 MPa- α 0.016 0.012 1 n 4.27 7.8 n/a

4.3 Model Verification Simulation temperature results for the bobbin and conventional tool configurations were compared to experimental temperatures to verify the accuracy of the models. A thermocouple was placed inside the tool tip during conventional trials and a representative node near the tool tip within the simulation recorded the temperature during the model execution. A comparative plot is presented in Figure 5.

Figure 5 The steady-state temperature comparison of conventional tool experiment and simulation.

In Figure 5, the experimental curve reaches the steady state condition much faster than the simulated result. However, the goal of the simulation was to model the steady state condition of welding. Thus, all the data used from the simulation was taken from the final time step (130 s). The peak temperature reached at 130 seconds was 328.8 °C. The steady state temperatures for the experimental welds were approximately 335 °C. Overall, the simulation proves to be a good approximation of the experimental results. Similar to the conventional tool, temperatures during welding with the bobbin tool were measured during processing. The thermocouple was placed on the advancing side, 20.5 mm from the weld centerline. Due to the stationary thermocouple, the verification was done by comparing the temperatures recorded by the thermocouple and the simulation as a function of distance from the tool centerline. This comparison is shown in Figure 6.

Figure 6 Temperature as a function of distance away from the bobbin tool. A comparative plot between the simulation and experimental results. Again, the simulation closely follows the experimental results peak temperature. Although the heat generation on the tool approach and subsequent dissipation of the simulation curve does not exactly match the experimental curve, the peak temperatures are within 2 °C. This heating is the most critical in the way that it alters the mechanical properties of the weld. Although the support tooling, such as backing plates, used in the welding were included in the simulation, it is difficult to capture all of the interactions away from the weld that lead to the correct heat dissipation

20 behavior. This results in the discrepancy between the simulation and experimental curves during the cooldown after the tool has moved past the point of interest.

5. Results and Discussion 5.1 Optical Microscopy The microstructures of the welds utilizing the bobbin tool geometry are presented in Figure 7 (a and b) where 2024 is on the advancing side and retreating side respectively. Figure 7a shows the 2024 as optically darker compared to the 7075. Both configurations contain the typical stir, thermal-mechanical, heat affected and base material zones that characterize most friction stir welds. Fine grain refinement occurs in the stir zone, likely due to dynamic recrystallization as described by Reynolds [16]. Both configurations show abrupt transitions between the SZ and TMAZ on the AS compared to the RS which do not depend on material location. However, the AS TMAZ in the AS-2024B configuration is reduced in size compared to the same feature in the AS-7075B weld. Measurements of the grain transformation between the recrystallized and base zones on the AS in the AS-2024B weld occurs within 0.4 mm whereas the AS-7075B weld transitions within 1.5 mm. A similar trend is seen on the RS with the AS-2024B transition being more gradual than that of AS-7075B. Here the transformation occurs over 6.5 mm and 5.0 mm respectively, measuring on the mid-section lines of each cross-section. Overall, the optical microscopy shows the 2024 alloy transitions between fine recrystallized grains in the stir zone to unaffected grains over less distance on both the AS and RS than the 7075 alloy. Little material mixing is observed in both material configurations, with the most mixing occurring along the mid-section line. Bands of material are present throughout the stir-zone, indicating that some material transferred across the centerline during processing.

Figure 7 Optical Micrograph for, a the AS-2024B configuration and, b the AS-7075B configuration.

The optical micrographs of the conventional tool welds show similar patterns to the micrographs of the bobbin tool configurations. Figure 8 (a and b) presents these results with 2024 appearing darker than 7075. The advancing side transition between the SZ and TMAZ is more abrupt than the corresponding retreating side for both the AS-2024C and AS-7075C welds. Like the welds in the bobbin tool configuration, the 2024 microstructure transitions over less distance than that of 7075 when compared on the same side. Figure 8a shows that it takes around 0.1 mm for 2024 to transition on the AS and 3.0 mm for 7075 to transition on the RS. Conversely, the opposite configuration shown in Figure 8b indicates that 7075 takes 0.45 mm near the mid-section line to transition on the AS. The 2024 transitions span over 1.5 mm on the RS.

Figure 8 Optical Micrographs for, a the AS-2024C configuration and, b the AS-7075C configuration In all welding configurations, the 2024 transition from stir zone to base material grain structure occurs over less distance than that of 7075, regardless of advancing or retreating side placement. The bobbin tool configuration produced a wider region of dynamically recrystallized grains due to the symmetric shoulder affected regions and simpler mixing patterns of the dissimilar alloys compared to the conventional tooling. However, grain size and orientation within the stir zone are identical for both tool geometries. Similar patterns are found in the TMAZ zones of both tool geometries where the advancing side exhibits a significantly more abrupt transition between the stir zone and base material than the retreating side. Additionally, both tool configurations show advanced stir patterns in the shoulder affected region of the advancing side when 2024 is placed on the retreating side before welding. In these configurations, bands of 2024 are observed penetrating deeper into the center of the cross- sections and indicate more thorough material mixing in these cases. Works from Reza-E-Rabby

23 et al. [43] and Koilraj et al. [44] support these observations by concluding that in dissimilar welding situations, placing the harder material on the advancing side can reduce stirring defects and can improve mechanical properties of the joint. For visual reference, an example of how the zones demarcate the optical micrographs is presented in Appendix E. Unmarked composite images for all four welding configurations are also contained within the same appendix. 5.2 Electron Backscatter Diffraction In addition to the optical microscopic evaluation of the welded cross-sections, electron backscatter diffraction (EBSD) was utilized to evaluate the grain size and orientation of the stir- zone produced by the conventional tool. The EBSD scan of a region of the SZ from the AS- 7075C is shown in Figure 9a. Each color and boundary represents a different grain orientation measured by the EBSD. Consequently, a result with colors spanning the entire spectrum and occurring randomly across the scanned area indicates a highly misaligned grain structure. Therefore, the SZ grains in AS-7075C captured by the EBSD are observed to be randomized and equiaxed.

Figure 9 Results of, a the SEM EBSD scan and, b the Mackenzie plot for the SZ in AS-7075C. A Mackenzie plot of grain orientations within the scanned region supports these observations as shown in Figure 9b. The number fractions of specific grain orientations closely follow a random distribution. Overall, the EBSD scan and Mackenzie plot indicate that the stir zone of the conventional tool weld experienced continuous dynamic recrystallization (CDRX) as outlined by Reynolds [16]. 5.3 Micro-Hardness Distributions The micro-hardness profiles of the bobbin tool configuration AS-2024B are presented in Figure 10a. Hardness minima are recognized on the advancing and retreating sides of the weld with values of 110 HV and 120 HV, respectively. Their locations indicate that the maximum material softening occurs near the edge of the SZ, otherwise marking the HAZ of the weld. Away from the centerline on both sides, the material gradually returns to the material properties of the base material. Additionally, the profiles indicate a degree of hardness recovery in the SZ from dynamic recrystallization and the material reaching temperatures needed for either full or

24 partial recovery of 2024 or 7075. Figure 10a shows a distinct hardness step across the weld centerline from AS to RS. The shoulder dominated profiles mark the midpoint step as biased 3-4 mm to the AS from the centerline while the midsection profile records the step occurring 1-2 mm toward the AS from the centerline. The range of measurements taken across the midsection step is less severe than in the shoulder dominated profiles, indicating better material mixing in the region. This is reinforced by the optical micrographs above.

Figure 10 Vickers Hardness Profiles for, a the AS-2024B configuration and, b the AS-7075B configuration. The micro-hardness profiles of the bobbin tool configuration AS-7075B are presented in Figure 10b. This weld configuration produced profiles similar to its counterpart (AS-2024B) with two hardness minima near the edges of the SZ with values of 121 HV and 108 HV on the advancing and retreating sides, respectively. The minima are shifted slightly toward the AS and the base material on the 7075 side of the weld achieves better recovery when compared to the AS-2024B configuration. In both the AS-2024B and AS-7075B configurations, the lowest hardness minimum was found accompanying the 2024 side of the weld. The same hardness step is present in the AS-7075B configuration, but it shows better dynamic recovery of the 7075 and gradual decline of the 2024 when compared to the step of the AS-2024B weld. Again, a less severe step is present along the midsection profile than in the shoulder dominated regions. The conventional tool configuration AS-2024C produced the midsection hardness profile shown in Figure 11a. Here, one hardness minimum was recorded on the 2024 side of the weld near the edge of the pin diameter with a value of 129 HV. A less severe minimum is also present on the 7075 side outside of the pin diameter but within the shoulder diameter. Inside the SZ, measurements varied significantly and indicate some material mixing along the midsection of the weld. This is supported by the optical micrograph showing bands of material present across the width of the stir zone. The midsection hardness profile for the AS-7075C conventional tool configuration exhibits similar characteristics to the AS-2024C weld, shown in Figure 11b. One hardness minimum is present on the 2024 side measured at 120 HV and a less severe minimum is present on the 7075 side between the extents of the shoulder and pin geometries. The profile of this configuration also shows a hardness step across the midsection of the weld, highlighting the interface between 2024 and 7075. The same trend is likely present in the AS-2024C weld, but its presence may be

25 hidden due measurements taken in adjacent material bands across the SZ that obscure this trend.

Figure 11 Vickers Hardness Profiles for, a the AS-2024C configuration ad b the AS-7075C configuration. Both conventional weld configurations show one severe hardness minimum on the 2024 side and a less pronounced minimum in the 7075. This is distinctly different from the bobbin tool configurations where there are severe hardness minima on both the 2024 and 7075 sides. This difference is likely due to contrasting heat inputs by the two tool geometries. 5.4 Temperature Analysis 5.4.1 Base Material Calorimetry Differential scanning calorimetric (DSC) analysis of the base metal was performed to find the dissolution and precipitation temperatures of the alloying phases in the 2024-T3 and 7075-T6 matrix. The 2024-T3 and 7075-T6 calorimetric curves are presented in 10 (a and b). The calorimetry curve for the 2024-T3 base material taken from the welded samples closely resembles established data. A study by Yang et al. [45] describes the temperatures of the exothermic and endothermic peaks caused by precipitate formation and dissolution in detail. Figure 12a presents an endothermic reaction at 228oC, an exothermic release at 265oC, and another endothermic reaction at 398oC. This corresponds to the GP zone dissolution (215oC), S’ and S phase precipitation (257oC), and S’ and S phase dissolution (300-400oC) outlined in Yang et al.’s work [45]. Figure 12b shows the calorimetry curve for 7075-T6 taken from the base metal of a welded sample. An endothermic reaction peaking at 198oC, exothermic releases at 235oC and 250oC, and long endothermic reaction peaking at 396oC describe the precipitate formation and dissolution profile for the tested alloy. Jung et al. [46] found GP zone dissolution (endothermic) at 200oC, η’ and η precipitation (exothermic) at 230oC and 255oC, and precipitate dissolution (endothermic) continuing past 350oC. Both curves follow previous studies and provide clear indications of both the dissolution and precipitate reactions discussed by Hatch [17] in section 2.2.

Figure 12 Calorimetric data for, a 2024-T6 and, b 7075-T6. 5.4.2 Simulation Results The primary data extracted from the welding simulations is the temperature profile transverse to the weld line at the location of tool interaction. These profiles predict the temperatures of the material in the BM, HAZ, TMAZ, and SZ on both sides of the weld. The temperatures can be directly related to both the optical micrographs and hardness distributions for each configuration. The temperature results for welds of different material placement configurations with the same tool are compared. The temperature results of welds between tool types are also compared. In both the material type and tool type comparisons, the mid-thickness results are examined for all

27 welding trials. Although hardness measurements were taken near the upper and lower shoulders of the bobbin tool trials, the mid-thickness results are adequate for the individual comparisons. The simulation cross-section results for the bobbin tool AS-2024B and AS-7075B weld configurations are presented in Figure 13. There was not a significant temperature difference between the two simulations. However, the AS peaked at a higher temperature than the RS, similar to studies by Hamilton et al. [25], [47].

Figure 13 Bobbin tool configuration temperature comparison of AS-2024B and AS-7075B. By separating the individual temperature distributions into two plots and superimposing the mid-thickness hardness profiles onto the figures, the effects of temperature on the hardness can be observed. Figure 14 and Figure 15 show these profiles and include reference lines of the precipitation and dissolution temperatures for the secondary phases of both alloys. The reference thresholds were sourced from the exothermic and endothermic peaks on the DSC curves discussed in section 5.4.1. Figure 14 shows the AS-7075B weld configuration. On the advancing side, the temperature reaches the GP phase dissolution threshold near the -20.5 mm mark. The 7075 hardness is decreasing rapidly at this point, which is expected when the GP phase dissolves. The hardness continues to decrease as the temperature rises through the η and η’ phase precipitation zones located near the -17 mm and -15 mm marks. There is agreement between the hardness and temperature profiles in this region because 7075 does not usually recover hardness during these

28 events. The hardness reaches its local minimum on the 7075 side at -14 mm. Simulation data show that material in this area peaked at a temperature above the η precipitation threshold. This suggests that precipitate coarsening has occurred, which likely caused the material softening indicated by the hardness. On the retreating side, similar congruence between the hardness and temperature profiles is observed. At the +18 mm mark, the simulation temperature reaches the GP phase dissolution threshold and the hardness starts to decrease. The temperature crosses the S’ and S precipitation threshold around +13 mm which marks the beginning of the precipitate growth on the 2024 side. The hardness minimum is located at +6.5 mm where the peak temperature is well above the threshold and causes significant precipitate coarsening.

Figure 14 AS-7075B Temperature-hardness comparison plot Together, the hardness profile of the AS-7075B cross-section and the simulation temperature curve support one another and confirm that the low hardness regions in both the 7075 and 2024 were a result of precipitation coarsening. Furthermore, this is the case for all four welding configurations. A full discussion on each remaining hardness-temperature plot is contained in Appendix C. The hardness data presented in section 5.3 indicates that the most severe hardness minima always occurred on the 2024 side. These minima are much more likely to affect the outcome of tensile testing and overall weld performance than the low hardness regions dominated by 7075. Moving forward, the focus of the analysis is on the 2024 side of each welding configuration.

The temperature-hardness comparison plot for AS-2024B is shown in Figure 15. The hardness and temperature interactions are observed to be similar to the those of the AS-7075B plot. Specifically, the 2024 side hardness minimum of each welds was ~110 HV, and the corresponding temperatures located at each minimum were 320 °C and 318 °C for the AS-7075B and AS-2024B configurations, respectively. This indicates that the low hardness regions for both configurations experienced the same overall heating for the same amount of time (both were welded with the bobbin tool) which lead to identical hardness reductions in the 2024.

Figure 15 AS-2024B Temperature-hardness comparison plot. Along the mid-thickness line, the maximum temperature was approximately 345 °C for both configurations. This peak occurs within the sir zone of the weld and is close to the dissolution temperature of both the 2024 S and 7075 η precipitates. Upon cooling, some of these dissolved secondary phases reprecipitate as finely dispersed GP phases which contribute to the hardness recovery throughout the stir zone. Compared to the results from the bobbin tool simulations, the conventional tool simulations resulted in less overall material heating. This difference is shown in Figure 16.

Figure 16 Comparison of conventional tool temperatures for AS-2024C and AS-7075C. Both material placement configurations welded with the conventional tool peaked about 30 °C lower than their bobbin tool counterparts. Like the bobbin tool simulations, little temperature difference was seen between the material configurations and the advancing side reached a higher temperature than the retreating side. The lower peak temperatures experienced by the conventional tool welds affected the precipitate behavior in the 2024 on both the AS and RS when compared to their bobbin tool counterparts. The temperature-hardness comparison plots for the AS-7075C and AS-2024C configurations are presented in Figure 17 and Figure 18, respectively. The temperature-hardness plot for AS-7075C in Figure 17 shows how the temperature affects the hardness for the conventional tool configurations. The temperature on the 2024 side of the cross-section reaches the GP phase dissolution threshold at the +15 mm mark. There is not a negative hardness trend once the GP phases reach the dissolution temperature but there is a small hardness increase when the temperature reaches the S’ and S phase precipitation threshold at +11 mm. As the temperature increases closer to the stir zone boundary, the precipitates coarsen and reach the hardness minimum of 120 HV at +3.5 mm.

Figure 17 AS-7075C Temperature-hardness comparison plot. The AS-2024C configuration shows similar trends to the AS-7075C configuration with one significant difference. In Figure 18, the hardness on the 2024 side reaches a minimum of 129 HV at -3.5 mm on the cross-section. This is significantly higher than the 120 HV recorded in the AS- 7075C case and the reason for the discrepancy was not immediately clear. The higher hardness configuration, AS-2024C, recorded a temperature of 313 °C at its minimum location and the lower hardness case of AS-7075C recorded a 308 °C temperature at its minimum location. The temperatures, by themselves, seem to predict a different outcome than that of the hardness. The higher temperature should indicate a more severe hardness dip because it would cause more significant S phase coarsening than the lower temperature counterpart. However, the more severe hardness minimum was found in the lower temperature AS-7075C configuration indicating that other factors must affect the ~10 HV hardness discrepancy.

Figure 18 AS-2024C Temperature-hardness comparison plot. The specific locations of the minima were examined against the optical images of the conventional tool weld cross-sections to explain the difference in hardness. Both minima occurred at 3.5 mm away from the centerline, but each was on a different side of the weld. The analysis of optical images presented in section 5.1 indicated that, for 2024, the TMAZ transitions on the AS and RS took 0.1 mm and 1.5 mm respectively. With the minima occurring on the edge of the stir zone in both material configurations, it is possible that the more gradual but wider TMAZ on the RS affected the minimum hardness of the 2024 when placed on that side. A discussion on 2XXX series alloys by Mahoney [16] states that, for conventional tool welding, the conductive heat transfer through the thickness could be different on the AS and the RS. Mahoney’s reasoning is that through thickness temperature variation is more evident in areas of a weld cross-section that experiences more mechanical deformation. The SZ and TMAZs will experience the temperature variation, but the HAZ and BM will not. Figure 19 presents the locations of the hardness minima overlaid on the optical images of the conventional tool configurations.

Figure 19 Hardness Minima locations on the conventional tool optical micrographs. When the 2024 was placed on the advancing side of the weld, the hardness minimum fell on the edge of the TMAZ in the HAZ. With 2024 on the retreating side of the weld, the hardness minimum fell inside the TMAZ. It is likely that the different sizes of the TMAZ zones on the AS and RS caused the hardness minima discrepancy. The RS has more residual mechanical deformation than the AS and this may have caused a temperature variation through the weld thickness large enough to lower the hardness on the RS. 5.4.3 Time at Temperature Evaluation To thoroughly analyze the difference between the bobbin and conventional tool welding configurations, evaluations were made on both the processing temperature and time at temperature. The two tool geometries required different processing parameter sets to create defect-free welds. These parameter sets had some effect on the 30 °C peak temperature difference between the bobbin and conventional welding trials. Although the peak temperatures influence how the secondary phases likely behave within the matrix, they do not account for the amount of time that the phases are exposed to the significant temperature. The bobbin tool configurations were welded with a feed rate of 127 mm min-1 compared to the 203 mm min-1

34 feed rate of the conventional tool welding trials. An arbitrary point of interest in the material would experience less heating time during a conventional tool trial than a bobbin tool trial because of the feed rate differences. As the tool moves past the point of interest, a heating curve against time with sharp lead in, peak, and slow dissipation temperature characteristics describes the interaction with the point of interest. The longer this curve stays above a precipitation temperature threshold, coarsening of the secondary phases will be more severe. The precipitation threshold for the 2024 is 264 °C and this time at temperature was calculated for an arbitrary point of interest along the mid-thickness line in both the conventional and bobbin tool configurations. This plot is shown in Figure 20. Material welded with the conventional tool experienced ~7 seconds above this threshold, while material welded with the bobbin tool experienced this for ~11.5. Overall, the 2024 in the bobbin tool configurations was exposed to higher peak temperatures for 4.5 seconds longer than 2024 in the conventional tool configurations. These two factors were likely the reason the minimum hardness of the 2024 was 10 to 20 HV lower for the bobbin tool trials when compared to the conventional tool results. The difference in hardness between the tool configurations and the discrepancy observed within the conventional tool trials significantly affect the performance of the welds during tensile testing.

Figure 20 Time at temperature comparison between the bobbin and conventional configurations of AS-2024. The x-axis represents the time of tool approach and departure from the point of interest.

5.5 Tensile Testing The average result for the tensile testing of each welding configuration is presented in Table 2. All trials produced failures on the 2024 side of the weld and the failures were close to the location of the hardness minimum in all cases. Compiled images of completed tensile tests are provided in Figure 21.

Figure 21 Failure locations from tensile testing of all welding configurations.

To confirm that the samples failed in a ductile region of the welds, fracture surfaces were examined under an SEM for each configuration. Ductile failures were identified in all cases and a micrograph from both tool configurations is provided in Figure 22. Images for all configurations are in Appendix D for reference. The ductile failures indicate that location was in a region of low hardness and support the results found from the hardness mapping and precipitate behavior estimates from the simulations.

Figure 22 SEM fracture surfaces for: a the AS-2024C, b the AS-2024B configurations.

Table 3 Tensile testing results for all material and tool configurations.

St. Dev UTS Trial Desig. Description UTS (MPa) YS (MPa) Fracture Loc. (MPa) AS-2024B Bobbin, 2024 AS 359 245 HAZ 2024 16.95 AS-7075B Bobbin, 7075 AS 383 246 HAZ 2024 16.60 AS-7075C Conv., 7075 AS 426 287 HAZ 2024 1.40 AS-2024C Conv., 2024 AS 444 302 HAZ 2024 3.01

The tensile testing results suggest the same trends observed in the hardness and temperature profiles in terms of congruency between hardness values and tensile values. In the case of the bobbin tool configurations, the hardness minima are remarkably similar across the two material configurations and it was expected that the welds would fail under the same stress. The mean ultimate tensile strength results of the AS-2024B and AS-7075B configurations are 24 MPa different. However, the standard deviations of these results indicate a significant variance between the trials in both configurations. A t-test, outlined in section 3, was performed on these data sets to determine the significance of these differences. The test returned a p-value of 0.0552 for a significance level of 0.05 which rejects the alternative hypothesis that the population means were different. Therefore, the differences in the results were concluded to be statistically insignificant. The test held a 49.7% power due to the small sample size of 5 specimens and a high standard deviation for each data set. Approximately 12 trials per material configuration would be required to achieve a 90% power, reducing the possibility of type II error. However, the space constraints on the welded sheets prevented more testing in this study. The hardness discrepancy between the AS-2024C and AS-7075C configurations is evident in the ultimate tensile strength results. Again, 2024 on the RS resulted in the lower hardness and produced the lower ultimate tensile strength result of 426 MPa. With 2024 on the AS, the tensile results were 444 MPa on average. Unlike the bobbin tool, a t-test indicated that this 16 MPa difference was statistically significant. A p-value of 3.221E-05 with a significance value of 0.05 and a test power of 100% verify that the null hypothesis was rejected. The low variance between the trials in both material configurations likely produced the high power of the test.

6. Conclusions and Future Work Aluminum alloys 2024-T3 and 7075-T6 were joined by FSW using bobbin and conventional tools while placing both materials on the AS and RS. The basis for comparison between the configurations was the production of defect-free weld zones in all cases. The resulting welds were examined for microstructural and mechanical differences using various microscopy techniques, microhardness mapping, and tensile testing. Each welding configuration was simulated in COMSOL Multiphysics to generate temperature profiles used to determine the extent of temperature effects on the post-welded aluminum. Comparing the bobbin tool configurations, the microstructures of both the AS-2024B and AS-7075B cross-sections show no visible defects. The 7075 shows more grain deformation and larger TMAZ transitions between the BM and SZ than the 2024. The extent of material mixing in the two configurations was similar, although deeper penetration of the RS alloy into the thickness of the SZ on the AS is observed in the AS-7075B configuration. It is unclear whether these mixing differences affected the results of subsequent mechanical testing. The hardness mapping of the cross-sections indicated that 2024 produced lower hardness minima than 7075 regardless of AS or RS placement. For both bobbin tool configurations, the 2024 minima were similar, around 110 HV. This was explained by the temperature profiles produced by the simulation, which showed that the bobbin tool produced enough heat to cause significant precipitate coarsening in the HAZs and TMAZs on both sides of each weld. The 2024 experienced similar heating peaks regardless of its placement and it was determined that precipitate coarsening was the main cause of the hardness minima. Tensile testing of the configurations resulted in similar weld strengths for both material placement configurations. The differences in mean UTS values for the AS-7075B and AS-2024B were not statistically significant. Overall, the welds failed on the 2024 side in the ductile, low hardness region expected from the hardness mapping and temperature simulations. For the conventional tool configurations, the microstructures of the cross-sections were free of defects. The same trends were seen in the transition zones as the bobbin tool configurations and the AS-7075C configuration showed deeper penetration of the RS material into the AS in the SZ than the AS-2024C configuration. The hardness mapping of the cross-sections showed that the 2024 produced lower hardness values than the 7075 on both sides of the weld. However, the severity of the minimum was more significant when 2024 was placed on the RS. The simulated temperature profiles did not explain the discrepancy because the predicted temperatures at the locations of the minima were within 5 °C of each other. This meant that the precipitate coarsening of the secondary phases was similar on both the AS and RS and that other factors were contributing to the discrepancy. Evaluating where the locations of the 2024 minima occurred on the cross-sectional micrographs led to the possibility that extensive grain deformation on near the minimum on the retreating side caused the lower hardness because of a through-thickness temperature gradient during cooling. Tensile tests on both material configurations showed that the lower hardness AS-7075C configuration resulted in lower ultimate tensile test results than its counterpart configuration. A t-test confirmed this conclusion that the results were statistically significant. As expected, the welds failed in the low hardness region on the 2024 side. Using the achievement of defect-free welds as the basis for comparison, the conventional and bobbin tool configuration resulted in welds with significantly different tensile properties. Because of differences in tool geometry and interaction with the material, the process parameters selected for the bobbin tool were significantly slower than those of the conventional tool. The

38 simulation predicted that the peak temperatures from the bobbin tool welds were 30 °C hotter than those of the conventional tool welds because of the processing, geometry, and interaction differences. The hotter peak temperature and related time at temperature for material in the bobbin tool configurations resulted in more significant precipitate coarsening of both the 2024 and 7075. This resulted in hardness minima between 10 and 18 HV lower than those of the conventional tool and translated to significantly lower average ultimate tensile and yield strength values. The standard deviations among the sampling groups suggest that the conventional tool welds were more consistent than the joining performed with the bobbin tool. The reasoning for the large variance in the bobbin tool tensile testing trials is recommended for future work. It is the author’s hope that work on this project will continue, as there are several areas for continued exploration. First, continued analysis of the bobbin tool welds by performing additional tensile testing would be valuable. The low statistical power of the hypothesis testing and the significant variance between the test specimens warrants further examination. Hardness measurements and microstructural examination from multiple cross-sections along the length of the welded plates might yield more information about material differences between test specimens. Second, the final notch preparation and testing of the fracture toughness specimens should be completed. This is an unexplored area in dissimilar alloy friction stir welding and will complement the mechanical characterization already completed. Further use of the wire EDM and hydraulic tensile testing machine are needed to proceed with testing. Third, further EBSD analysis in the SZ of the remaining three welding configurations could better classify the behavior of the material flowing around the tool during processing. Fourth, a more in-depth exploration of the welding microstructure for each configuration would complement the thermal modeling and hardness mapping already completed. Use of the TEM to examine the precipitate behavior around the TMAZ of each configuration may highlight important differences between the specimens.

7. Appendices 7.1 Appendix A – Tensile Specimen Preparation The tensile testing specimens were designed to the ASTM E8-04 standard and conform to the sub-size classification. The dimensions (in inches) for each specimen are provided in Figure 23. Serial numbers were engraved in the gripping region on both sides of the weld centerline to identify the AS and RS alloy. The plates (12” x 8”) were fixtured in the milling machine using the long-extruded edge as a reference edge, collinear to the long table axis of the machine. Serial numbers, saw-cutting guides, and fixturing holes were engraved and drilled from the plate in this first fixture. The fixturing holes, shown in the gripping regions of the specimen in Figure 23, were added for use in fixturing the sample blanks once they were extracted from the plate. This ensured that a reference was still available to keep the final machining of the blank perpendicular to the weld centerline. The clamping fixture designed, manufactured, and used for the rough and final machining of the samples is depicted in Figure 24. The clamping shoulder screws are shown in red, the specimen in blue, the bottom fixture plate in grey, and the top clamp in black.

Figure 23 Sub-size sheet-type tensile specimen.

Figure 24 Sub-size sheet-type tensile specimen machining assembly. 7.2 Appendix B – Fracture Toughness Specimen Preparation The fracture toughness specimens were designed to the ASTM E1820-06 standard and conform to the compact tension type. The specimens were designed to fit entirely inside the width of the shoulder affected region of the weld. The exact dimensions of the specimens are provided in Figure 25. The specimens were machined in the same fixturing of the plates as the tensile specimens. Again, reference saw-cut guides and re-fixturing features were engraved and machined in the plate during the first fixture. The holes intended for the cyclical testing doubled as locating features on a second fixture used to bring the specimens to the final outer dimensions. This fixture assembly is depicted in Figure 26. The clamping screws are shown in red, the specimen in blue, the bottom plate in grey, the alignment slide in black and its clamping screw in gold. Although the starter notch is labeled on the drawings and in the rendered image, the machining of the feature was not completed. It was intended that the notch be machined in the wire EDM using 0.010” wire. Due to the low clamping forces on the specimen during this machining, the specimen could be cantilevered off of the table edge for this process and secured by a single toe clamp. The holes in the specimens should be used to align the machining axis so the notch is collinear to the weld centerline. Two testing clevises were designed and manufactured for the testing of the fracture toughness specimens. The clevises conform to the ASTM E1820-06 standard and were manufactured out of A2 tool steel that was heat treated after machining. They were made specifically for use on an MTS machine with hydraulic grips capable of securing 1.18-inch diameter tools and specimens. The exact dimensions of the clevises are provided in Figure 27. Flat spots on the horizontal surfaces of the loading holes were included so that the specimen can rotate during testing.

Figure 25 Compact tension-type fracture toughness specimen.

Figure 26 Machining fixture assembly for C(T) specimens.

Figure 27 Drawing of clevis for testing C(T) specimens.

7.3 Appendix C – Temperature Profile Evaluations of Welding Configurations 7.3.1 AS-7075B Configuration This description supplements Figure 14, presented in section 5.4.2.

On the advancing side, the temperature reaches the GP phase dissolution near the -20.5 mm mark. The 7075 hardness is starting to decrease rapidly at this point, which is expected when this phase dissolves. The hardness continues to decrease as the temperature continues to climb through the η and η’ phase precipitation zones located near the -17 mm and -15 mm marks. Again, this is expected because 7075 does not usually recover hardness during these events. The hardness reaches its local minimum on the 7075 side at -14 mm after the secondary phases have precipitated out and begun to coarsen. Based on the surveyed literature in section 2.2.2, the 7075 is expected to have a minimum hardness in this region. On the retreating side, similar trends are present when the temperatures are examined along with the hardness profile. At the +18 mm mark, the simulation temperature reaches the GP phase dissolution threshold and the hardness starts to decrease. The temperature crosses the S’ and S precipitation threshold (264 °C) around +13 mm which marks the beginning of the precipitate coarsening on the 2024 side. The hardness minimum is located at +6.5 mm and the simulation profile indicates that significant precipitate coarsening has occurred from a temperature of 320 °C at the location. 7.3.2 AS-2024B Configuration This description supplements Figure 15, presented in section 5.4.2.

On the advancing side of the weld, the 2024 reaches the GP phase dissolution temperature near the 18 mm mark. The hardness decreases shortly after this and continues to decrease until it reaches the minimum. Between the GP dissolution threshold and the minimum, the temperature reaches the S’ and S phase precipitation value at -14 mm. Unlike the work by Jones et al. [2], no hardness increase is observable at or near this point. However, 2 mm before the threshold is reached, the downward hardness trend levels out briefly. The temperature of this configuration does not approach the solution temperature of 2024, so it is possible that the temperatures required for S’ and S phase coarsening were not reached long enough to cause the hardness recovery recorded in other work. The hardness minimum at -9.5 mm is well past where the simulation predicts the temperature passes the coarsening threshold of the secondary phases, which is the expected result. On the retreating side, the 7075 reaches the GP phase dissolution temperature at the +20.5 mm mark. The hardness begins decreasing shortly after, in a direction oriented toward the weld centerline. The temperature crosses the η’ and η precipitation thresholds at the +16.5 mm and +15 mm marks and the hardness minimum follows shortly after at the +13.5 mm mark likely due to significant precipitate coarsening. 7.3.3 AS-7075C Configuration This description supplements Figure 17, presented in section 5.4.2.

The 7075 side of the AS-7075C configuration shows good agreement between the simulation temperatures and the recorded hardness values. At the -20.5 mm position, the temperature of the simulation crosses the GP phase dissolution temperature threshold. The trend of hardness oriented toward the weld center sharply decreases which indicates that the GP phase dissolution decreased the strength of the 7075. At the -15.5 and -13.5 mm marks, the temperature of the

46 simulation reaches the η’ and η precipitation thresholds. Again, the hardness responds with a declining trend until it reaches the minimum on the AS at -5 mm. The precipitates formed after the threshold coarsen and soften the material until the SZ begins. On the RS of the weld, the 2024 maintains the base material hardness until the +16.5 mm mark. Here, the hardness dips slightly. Simultaneously, the simulation temperature reaches the GP phase dissolution temperature. The 2024 is expected to decrease hardness when the strengthening GP phase dissolves into the matrix, so the simulation temperature and hardness values agree at the +16.5 mm mark. The hardness starts to decline sharply at the +9.5 mm mark until it reaches the minimum at +3.5 mm and the simulation temperature crosses the S’ and S phase precipitation threshold close to the 9.5 mm mark. The hardness and simulation agree through the hardness minimum because the temperature climb coarsens the precipitates significantly which softens the material. Overall, the hardness profile matches the precipitation and dissolution events that occur as a result of the simulated temperature on both sides of the weld. 7.3.4 AS-2024C Configuration This description supplements Figure 18, presented in section 5.4.2.

The 2024 side of the AS-2024C configuration shows good agreement between the hardness profile and the simulated cross-sectional temperature. On the advancing side, the hardness is stable until it starts to climb at the -15 mm mark. The simulated temperature reaches the GP phase dissipation threshold at the -16 mm mark, but no accompanying decrease in hardness is observed. Although this result was not expected, the material near this point only experienced temperatures above the GP phase dissolution and below the S’ precipitation thresholds for a short time. There was likely not enough time at temperature to cause a significant hardness dip on the AS. This contrasts the AS-7075C configuration where the 2024 experienced a small hardness dip on the RS when the temperature in the simulation reached the GP phase dissolution threshold. The increase in hardness around the -15 mm mark occurs near when the temperature reaches the S’ and S phase precipitation threshold. This can be explained by the strengthening that occurs shortly after the S’ phase precipitates out of the aluminum matrix (section 2.2.1). Then, the 2024 decreases sharply, reaching the hardness minimum at the -3.5 mm mark. The temperature continues to rise after reaching the precipitation threshold, which results in material softening due to precipitate coarsening. This indicates good agreement between the simulation and hardness trends. Moving to the RS of the weld, the 7075 hardness decreases sharply, beginning at +17.5 mm. The simulation temperature reaches the GP phase dissolution threshold at -19.5 mm which is shortly before the hardness decline. The hardness decline can be explained by the dissolution of the GP phases. The hardness continues to decline until it reaches the minimum at +6.5 mm. The temperature here exceeds the η’ and η precipitation thresholds, which occur at +14 mm and +12.5 mm. Again, the low hardness and high temperature agree, suggesting material softening occurred from precipitate coarsening.

7.4 Appendix D – SEM Fracture Surface Micrographs for All Welding Configurations

Figure 28 SEM imaged fracture surfaces of the a AS-2024C, b AS-2024B, c AS-7075C and, d AS-7075B configurations at 5 keV.

Figure 29 Wide field SEM imaged fracture surfaces of the a AS-2024C, b AS-2024B, c AS- 7075C, d AS-7075B configurations at 5 keV. 7.5 Appendix E – Example welding zone diagram This diagram, Figure 30, shows how the welding zones demarcate the AS-2024B configuration. It is intended as an example and is not to exact scale. Similar zones can be observed on the other three configurations.

Figure 30 Welding zone demarcation diagram for the AS-2024B configuration.

Figure 31 Composite images for: a the AS-2024B and, b the AS-7075B configurations.

Figure 32 Composite images for: a the AS-2024C and, b the AS-7075C configurations.

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