Performance Testing and Modeling of Ultra- High Strength Steel and Complex Stack-Up Resistance Spot Welds

Transcription

1 Performance Testing and Modeling of Ultra- High Strength Steel and Complex Stack-Up Resistance Spot Welds THESIS Presented in Partial Fulfillment of the Requirements for the Degree Master of Science in the Graduate School of The Ohio State University By Andrea Jane Peer Graduate Program in Welding Engineering The Ohio State University 2017 Master's Examination Committee: Dr. Wei Zhang, Advisor Menachem Kimchi, Co-Advisor Dr. David Phillips

2 Copyrighted by Andrea Jane Peer 2017

3 Abstract Hot stamped boron steels, such as Usibor 1500, have been increasingly used in automotive structural components for light-weighting and impact resistance. Classified as an ultra-high strength steel, these alloys have superior strength with tensile strengths exceeding 1500 MPa. The rapid heating and cooling thermal cycle during resistance spot welding can significantly alter the martensitic base metal microstructure, resulting in formation of coarse-grained and subcritical heat-affected zones (CGHAZ and SCHAZ) with inferior mechanical properties. The martensitic CGHAZ is adjacent to the weld nugget and experiences the most time above the AC3, which allows for austenite grain growth. The SCHAZ is next to the unaffected base metal and does not reach the AC1 during welding, thus the base metal microstructure is over-tempered into cementite and ferrite. The present research aims at developing the fundamental knowledge of plastic deformation and fracture behaviors of ultra-high strength steel resistance spot welds. As a resistance spot weld comprises highly inhomogeneous microstructure, the overall research approach is based on studying the local (or microstructure-dependent) mechanical properties for individual regions in the weld as well as their interactions with weld geometry on the deformation behavior. Specifically, optimal welding parameters are determined to produce welds of appropriate nugget diameter for 2T Usibor 1500 with a ii

4 gauge thickness of 1.5 mm. Micro-hardness mapping and metallographic analysis allow for characterization of the weld metal, CGHAZ, SCHAZ, and base metal of the spot weld. Quasi-static tensile testing with digital image correlation (DIC) is used to determine the local stress-strain behaviors of each region using bulk microstructural samples created in a Gleeble thermal-mechanical simulator. Conventional and innovative resistance spot weld mechanical testing methods are used to generate more knowledge on the deformation of joints under various loading conditions. Sectioned tensile shear testing and single-sided wedge testing procedures have been established to use 2-D DIC for in situ observations of local deformation on the exposed weld cross-section during testing. A mechanical model, developed using Abaqus finite element analysis (FEA) code by incorporating the local constitutive behaviors of RSW joints, is used to better understand the effect of weld nugget profiles on the stress state present during loading. The FEA model is validated by comparing the simulated strain fields to the experimentally measured strain fields. The knowledge generated in this study can help improve the accuracy of predicting spot weld fracture of ultra-high strength steels in the automotive industry. Particularly, the fine-resolution, coupon-scale model developed in this research will be useful for implementation into coarse-resolution, full-scale models for crash simulation and optimization of vehicle components. iii

5 Acknowledgments I wish to thank my advisor Dr. Wei Zhang for his guidance, support, and inspiration. The environment in which you have created allows for individual creativity and group collaboration to achieve new, innovative solutions. I am grateful of Honda R&D Americas, Inc. for providing financial support throughout my academic career. I would individually like to thank Tim Abke for being a point of contact between The Ohio State University and Honda through the Simulation Innovation and Modeling Center (SIMCenter). You have enabled a great partnership that allows our research to impact real-world industrial problems. I would also like to thank the rest of the Welding Engineering faculty and staff for implanting their knowledge and creating an open environment for learning. Specifically, I would like to thank Menachem Kimchi for sharing his resistance welding expertise and Dr. David Phillips for serving on my committee. Both have been mentors throughout my academic career. Thank you to my fellow classmates, especially Ying Lu, and colleagues for sharing ideas, knowledge, and laughs throughout the years. Lastly, I wish to thank my parents and brother for their unconditional support. You have taught me how to be poised in the difficult times and humble in the righteous times. Thank you for always believing in me and encouraging me to pursue my dreams. iv

6 Vita Canal Winchester High School B.S. Welding Engineering, The Ohio State University M.S. Welding Engineering, The Ohio State University Publications Andrea Peer, Ying Lu, Tim Abke, Menachem Kimchi, and Wei Zhang "Deformation Behaviors of Subcritical Heat-affected Zone of Ultra-high Strength Steel Resistance Spot Welds." in 9th International Seminar & Conference on Advances in Resistance Spot Welding. Miami, (3 2016). Paper No. 12 Ying Lu, Andrea Peer, Tim Abke, Menachem Kimchi, and Wei Zhang "Heat-Affected Zone Microstructure and Local Constitutive Behaviors of Resistance Spot Welded Hot-Stamped Steel." in Sheet Metal Welding Conference XVII. Livonia, ( ). Fields of Study Major Field: Welding Engineering v

7 Table of Contents Abstract... ii Acknowledgments... iv Vita... v Publications... v Fields of Study... v Table of Contents... vi List of Tables... xii List of Figures... xiii Chapter 1: Introduction... 1 Chapter 2: Literature Review Resistance Spot Welding Welding Parameters Nugget Formation Weldability Automotive Material Selection vi

8 2.2.1 Weld Characterization Hot Stamped Boron Steel Dual Phase Steel High Strength Low Alloy Steel Galvanneal Coated Mild Steel Mechanical Properties and Testing Methods Mechanical Testing Methods RSW Failure Modes Digital Image Correlation Crash Simulation Techniques Complex Stack-Up Configurations Chapter 3: Objectives Process Optimization and Characterization Constitutive Modeling Weldment Mechanical Testing Weldment Computational Modeling Complex Stack-Up Configuration Chapter 4: Ultra-High Strength Steel Characterization Materials vii

9 4.2 Approach RSW Process Optimization Sample Preparation Microstructural Analysis Gleeble Simulation Microstructural Tensile Testing Results and Discussion RSW Process Optimization UHSS Metallographic Characterization Gleeble Microstructural Simulation UHSS Microstructure-Specific Mechanical Properties Conclusions Chapter 5: Conventional Mechanical Testing Methods Materials Approach Pure Tension Testing Tensile Shear Testing Metallographic Analysis Computational Modeling viii

10 5.3 Results and Discussion Tension Testing Tensile Shear Testing Conclusions Chapter 6: Single-Sided Wedge Testing Materials Approach Experimental Development Metallographic Analysis Computational Modeling Results and Discussion Effect of Weld Nugget Size Effect of Weld Leg Size Effect of Welding Schedule Effect of Testing Configuration Conclusions Chapter 7: Complex Stack-Up Analysis Materials Approach ix

11 7.2.1 Resistance Spot Welding Metallurgical Analysis Results and Discussion Thin/Thick 2T Stack-Up Thin/Thick/Thick 3T Stack-Up Thick/Thick/Thick 3T Stack-Up Conclusions Chapter 8: Summary and Conclusions Conclusions Ultra-High Strength Steel Characterization Conventional Mechanical Testing Methods Single-Sided Wedge Testing Complex Stack-Up Analysis Recommendations for Future Work Failure Criterion for Computational Modeling Single-Sided Wedge Testing Complex Stack-Up Analysis Impact of Research References x

12 Appendix A: Calculations Load vs. Stroke Curves for MTS System Constitutive Behavior Development xi

13 List of Tables Table 1: Potential welding defects that can affect the performance of the weld [38] Table 2: Chemical composition of hot stamped boron steel [wt %] [43] Table 3: Chemical composition of JSC 980 YL [46] Table 4: Chemical composition of JSC 590 R [55] Table 5: Chemical composition of JAC /45 [56] Table 6: Minimum Acceptable Weld Size taken from AWS D8.1M [17] Table 7: Recommended process parameters for 2T 1.5 mm stack-up Table 8: Microstructural-specific Material Properties Table 9: Common failure modes of tensile shear testing and relevant data Table 10: True fracture strains calculated by the area reduction in pure tensile samples xii

14 List of Figures Figure 1: The past and projected use of different advanced high strength steels in the automotive industry [1]... 2 Figure 2: Resistance welding configuration with typical resistance and temperature measurements for a 2T stack-up... 7 Figure 3: Bulk resistance versus temperature for common RSW materials [20]... 8 Figure 4: Simple schematic showing the four phases making up the total weld time for the resistance spot welding process... 9 Figure 5: Schematic of the dynamic resistance developed during spot welding of uncoated steel [21] Figure 6: Schematic of the dynamic resistance developed during spot welding of coated steel [21] Figure 7: Example of a current range and lobe curve for resistance spot welding Figure 8: Different lobe curve profiles depending on material selection Figure 9: Global formability diagram showing the range of properties of common steel classifications [1] Figure 10: Schematic diagram of the HAZ sub-regions compared to the Fe-Fe3C equilibrium phase diagram [40] xiii

15 Figure 11: Hardness profiles of a HSLA compared to three different AHSS [1] Figure 12: Timeline of the hot forming process [44] Figure 13: Example stress versus strain curve of a metallic material, pointing out characteristic features Figure 14: Example of how to measure the weld button size post destructive testing [17] Figure 15: Schematic of a tension test specimen [62] Figure 16: Schematic of a cross tension test specimen [17] Figure 17: Loading configuration for cross tension testing [17] Figure 18: Loading configuration for tensile shear testing [17] Figure 19: Loading versus elongation of single spot weld steel specimens subjected to tensile shear loading [1] Figure 20: Example of set-up used by Payen et al. for wedge testing [66] Figure 21: Common fracture modes seen in the automotive industry [1] Figure 22: The basic steps of the DIC algorithm implemented in Ncorr using an initial guess and iterative optimaization scheme to find a refined solution [69] Figure 23: Comparison of u (x) displacement along a specific line for Ncorr and Vic 2D Figure 24: Comparison of strain (ɛxx) contours obtained for a compressive force on an epoxy ring for Ncorr and Vic 2D Figure 25: Different loading cases [76] xiv

16 Figure 26: Some of the critical structures that must have adequate strength for passenger crash safety Figure 27: Electrodes used by Coon et al. for experimental testing [10] Figure 28: Sequencing of weld current and force used to create a nominal welding schedule by Gould et al. [12] Figure 29: Depending on sheet thickness the nugget formation begins at the geometrical center of the stack-up or the sheet to sheet interfaces Figure 30: Minimum tension shear strength requirements taken from AWS D8.1M [17] 64 Figure 31: Example of the surfaces used for micro-hardness measurements Figure 32: Temperature profiles of physically- simulated (a) CGHAZ and (b) SCHAZ. 67 Figure 33: Tensile sample geometry [mm] Figure 34: Process parameter visual showing the current, force, and time Figure 35: Current range for 2T Usibor RSW Figure 36: Cross-section of the Usibor T stack-up with labels identifying the different microstructural regions Figure 37: Microstructural regions formed as a result of temperature distribution of the weldment during the resistance spot welding process Figure 38: (a) hot stamped martensitic base metal (b) re-solidified martensitic weld nugget (c) martensitic coarse-grained heat affected zone (d) martensitic fine-grained heat affected zone Figure 39: Subcritical heat affected zone (SCHAZ) xv

17 Figure 40: Hardness profiles of the RSW Usibor 1500 (a) longitudinal plane (b) conventional transverse plane Figure 41: Hardness profile taken from the weld centerline profile Figure 42: (a) as-welded CGHAZ, (b) Gleeble simulated CGHAZ, (c) as-welded SCHAZ, and (d) Gleeble simulated SCHAZ Figure 43: Snapshot of (a) deformation and (b) axial strain distribution measured using DIC on the simulated SCHAZ sample surface during tensile testing just prior to fracture Figure 44: True stress versus strain curves for Usibor 1500 base metal, CGHAZ, and SCHAZ Figure 45: Sample geometry for pure tension testing [mm] Figure 46: Sample geometry for tensile shear testing (a) conventional and (b) modified. Dimensions in mm Figure 47: Half models with applied symmetry and mesh for mechanical testing simulations Figure 48: Succession images of tension failure in one sheet (a) initial prior to loading (b) crack initiation (c) crack propagation and the corresponding DIC Eyy strain ( y) accumulation maps Figure 49: Load and Strain versus time for a pure tension test Figure 50: Abaqus simulation of tension test with (a) analysis with homogeneous base metal properties (b) strain accumulation at crack initiation with inhomogeneous profile 94 xvi

18 Figure 51: Failed tensile shear test specimens (a) interfacial (b) partial button pull out (c) SCHAZ button pull out and (d) expulsion button pull out Figure 52: Load versus stroke curves for different failure modes Figure 53: Strain accumulation at the notch-like feature and SCHAZ locations for (a) homogeneous weld profile and (b) unique heterogeneous weld profile Figure 54: XFEM crack initiation during tensile shear testing of a weld that experimentally failed by button pull-out (top) initial fracture at interface (bottom) further pulling shows failure initiation at the SCHAZ Figure 55: Common notch profiles for 2T Usibor 1500 weldments (a) CGHAZ blunt (b) WM blunt (c) WM bi-notch and (d) WM sharp Figure 56: Abaqus simulations displaying PEEQ corresponding to the different notch geometries Figure 57: Stroke versus maximum principal strain for locations near the different notch features Figure 58: Images and strain accumulation of sectioned weld samples subjected to tensile shear loading at (a) onset of deformation and (b) immediately before ultimate failure. 106 Figure 59: Sample geometry for wedge testing [mm] Figure 60: Two different geometries used for wedge testing [mm] Figure 61: Wedges are plunged at the interface of the weld during testing Figure 62: A Nikon D7100 camera is used to record the deformation behaviors during wedge testing Figure 63: Single-sided wedge test geometry and mesh for simulation xvii

19 Figure 64: 2T Usibor 1500 current range with markers indicating tested weld sizes. 115 Figure 65: Wedge testing results showing weld size versus insertion depth for 3 pulse welding schedule Figure 66: Simulated single-sided wedge testing results showing maximum plastic strain accumulation for three different weld sizes Figure 67: Failure modes of single-sided wedge testing (a) interfacial (b) CGHAZ (c) weld metal (d) SCHAZ Figure 68: Load versus displacement curves for the different failure modes of singlesided wedge testing (NOTE: tests were stopped prior to ultimate failure, except in case of interfacial failure) Figure 69: Normal strain map in the Y direction measured by DIC for (a) interfacial (b) weld metal and (c) SCHAZ fracture paths Figure 70: Single-sided wedge testing results showing leg size versus insertion depth for 3 pulse welding schedule Figure 71: Schematic comparison of the welding schedules for the three-pulse Figure 72: Single-sided wedge testing results showing weld size for 1- and 3-pulse welds versus insertion depth Figure 73: Simulated maximum plastic strain accumulation for different degrees of misalignment in single-sided wedge testing Figure 74: Complex stack-up configurations for 2T and 3T welds Figure 75: Effect of time on thin/thick stack-up (a) 4 cycles, 7.8 ka (b) 6 cycles, 8.0 ka and (b) 8 cycles, 8.0 ka xviii

20 Figure 76: Effect of current on thin/thick stack-up (a) 8 cycles, 8.0 ka and (b) 8 cycles, 8.5 ka Figure 77: Effect of current on thin/thick/thick stack-up (a) 8 cycles, 7.0 ka (b) 8 cycles, 7.7 ka and (c) 8 cycles, 8.6 ka Figure 78: Close-up of the penetration measured in the thin sheet of the stack-up Figure 79: Schematic view of the current density mismatch between the softer, more conductive material and the harder, more resistive bottom sheet seen in the 3T stack-up Figure 80: Effect of time on thick/thick/thick stack-up (a)15 cycles, 6.4 ka (b) 20 cycles, 6.8 ka and (c) 30 cycles, 7.2 ka Figure 81: Example of the output file from the MTS system Figure 82: Extensometer displacement exported from Ncorr Post Figure 83: True and engineering stress versus strain for the CGHAZ simulated sample155 xix

21 Chapter 1: Introduction With rising concerns about human-induced greenhouse gases and the always present concern for human safety, global initiatives with aggressive targets have been placed on the automotive industry to lightweight vehicles while still improving safety features of vehicles [1]. Automakers are engineering new ways of balancing performance, safety, fuel efficiency, and affordability using new materials and designs to meet these demanding requirements. In luxury vehicles, low density materials like aluminum, magnesium, and composites have been used to meet light weighting goals. While these materials can lightweight vehicles and improve fuel efficiency, their introduction into pre-existing processes is limited. To determine an effective solution, the steel industry has been developing newer materials with increased strength and formability to allow for thinner gauges to be used to reduce weight without adversely effecting safety [1]. Forming and joining processes for current materials can efficiently be altered to suit the newer classes of advanced high strength steels. The use of existing equipment and tooling can eliminate the need of excess costs associated new innovative solutions. The unique microstructures and metallurgical properties that allow for better crash performance and superior strength in new steels have created a surge in their use for automotive applications (Figure 1). One of the steels that will be increasingly used in the 1

22 automotive industry is hot stamped boron-alloyed steels, which are often referred to as ultra-high strength steels. The nominal tensile strength of such steels is greater than 1500 MPa due to the martensitic microstructure that forms after the hot stamping heat treatment (e.g., austenization by water quenching or die quenching). Figure 1: The past and projected use of different advanced high strength steels in the automotive industry [1] Resistance spot welding (RSW) is the dominant process used to join body closures and structural components of automobiles. Rapid heating and cooling during welding, alters the base metal microstructure of the advances and ultra-high strength steels. Research has heavily studied how this may affect the overall quality of RSW joints used in vehicles [2]-[7]. Several regions form in the heat affected zone (HAZ), located outside the weld nugget. The coarse-grained HAZ is the region immediately adjacent to the weld nugget and reaches temperatures well above the AC3 temperature of the steel allowing for significant austenite grain growth that transforms to martensite upon cooling [6]. The subcritical HAZ, which is farthest away from the weld nugget and next to the 2

23 base metal, experiences peak temperatures slightly below the AC1 temperature of the steel. Heating below the AC1 temperature results in martensite tempering causing decomposition into ferrite and cementite, which reduces the hardness to much lower than the martensitic microstructures of the various other regions [4]. Jong et al. measured the hardness of the subcritical HAZ compared to the base metal as 320 Hv versus 550 Hv for Usibor 1500 [6]. The reduction in hardness signifies the tensile strength of the softened subcritical HAZ is much lower than the martensitic base metal and weld nugget. Dancette et al. simulated the different HAZ microstructural regions for DP 980 using a Gleeble thermo-mechanical testing system [2]. They found the tensile strength of the subcritical HAZ was only 53% of the coarse-grained HAZ, but the elongation was much better. The microstructural regions of the heterogeneous resistance spot weld can play a large role in the overall mechanical properties of the joint. The microstructure is dependent on the RSW process parameters, material gauge, and material composition. Safety being the utmost concern, the government has established new tests and standards to rank the performance of vehicles. The National Highway Traffic Safety Administration (NHSTA) and Insurance Institution for Highway Safety (IIHS) have set vehicle safety standards for impact resistance, restraints, and fuel economy, while also encouraging automakers to improve the frontal, side, and rear impact, roof strength, and rollover ratings [8]. Burget and Sommer studied the failure behaviors of complex resistance spot welds made of a hot stamped boron steel and low-alloy steel [9]. Depending on the loading conditions, the failure mode can change from a pull-out fracture initiating in the softened HAZ of the hot stamped boron steel to a pull-out failure 3

24 in the coarse-grained HAZ. The complex failure behavior poses a major technical challenge for computer aided engineering (CAE) driven design and optimization of structures and their welds for safety and light-weighting. The ability for the vehicle structures to withstand loading in a crash event is a key design aspect that must consider the material and joint characteristics. Complex joint configurations, involving assorted materials, are being researched to allow for additional light weighting and design creativity [10]-[13]. The performance of these joints is difficult to determine without knowledge on how the materials and geometry endure numerous loading modes. Critical vehicle safety features must absorb sufficient energy to preserve the structure of the passenger compartment. CAE-driven design and manufacturing must accurately predict the deformation and failure of spot welds to be of use for crash modeling. Literature searches show performance models that have been developed to gain a better understanding of the mechanical deformation and failure of spot welds. Eller et al. has used inverse FEM optimization to determine plasticity model parameters that best fit with experimental deformation data [14]. Experimental hardness data is typically needed to linearly scale the flow stress data of the HAZ of resistance spot welds. Significant changes in welding parameters or material selection will require new experimental data to be generated. Integrating process and performance models for RSW is still evolving and must continuously be configured for newly developed materials. The main purpose of this work is to develop a quantitative understanding of the complex deformation and fracture behaviors of ultra-high strength steel resistance spot 4

25 welds comprising inhomogeneous microstructure. Chapter 2 contains a review of relevant literature on resistance spot welding, mechanical testing, and simulation techniques. Chapter 3 discusses the pertinent objectives of the subsequent chapters regarding process optimization and characterization, constitutive modeling, mechanical testing, computational modeling, and the development of complex stack-up configurations. The characterization utilizes microscopy, Gleeble simulation, and digital image correlation to develop constitutive behaviors for the unique heat affected zone regions of Usibor 1500 resistance spot welds. Mechanical testing methods using conventional techniques and innovative solutions for in situ observation are used to validate and refine computational models developed with Abaqus FEA. Finally, as described in Chapter 7, knowledge gained from the welding of ultra-high strength steels is then used to develop welding conditions for a complex stack-up consisted of three dissimilar materials. 5

26 Chapter 2: Literature Review 2.1 Resistance Spot Welding Resistance spot welding (RSW) has been a predominant welding process in the automotive industry since the 1930s [15]. There are numerous standards that exist from the company to world level that detail the common practices and procedures used for spot welding of various materials. In the work presented, International Standard Organization (ISO) [16], American Welding Society (AWS) [17], and Honda Engineering Standards (HES) [18] are compared. Each standard is sited when respectively used for welding procedure development and validation. The resistance spot welding process is developed around the resistance created when current is passed through a completed circuit. A current, either alternating (AC) or mid-frequency direct (MFDC), is passed through water-cooled electrodes at a low voltage resulting in heating of the clamped materials and fusion at the faying surfaces between the workpieces. A schematic showing the typical configuration for a two-sheet (2T) stack-up is shown in Figure 2. The heating that is generated by the current deviates throughout the stack-up due to variations in the ability of the current to pass, based on material and interfacial properties. There are two types of resistance, bulk and contact, generated during welding. Bulk resistance is generated in each of the materials, including the electrodes and 6

27 workpieces. The bulk resistance is a material specific property and is usually influenced by composition and temperature. Contact resistance is generated at the interfaces between the different parts, either electrode to workpiece or workpiece to workpiece. This value is dependent on material, surface conditions, temperature, and pressure; thus, making it a very important variable in process development [19]. Figure 2: Resistance welding configuration with typical resistance and temperature measurements for a 2T stack-up In Figure 2, the bulk resistances of the electrodes (1 and 7) are minimal and the bulk resistances of the workpieces (3 and 5) are relatively low. These values are a constant throughout the entire thickness but can vary with material and temperature, as seen in Figure 3. The resistivity of each material increases with temperature. The resistance of copper is significantly less than steel or aluminum at high temperatures, explaining why copper electrodes are used to conduct current to steel and aluminum workpieces for welding. To create a weld between the workpieces, the resistance must be 7

28 the highest at the faying surface between the workpieces. In Figure 2, at the points of contact between two different surfaces, there are peak resistance levels. This is the location of interest where melting will occur to form a weld nugget. Figure 3: Bulk resistance versus temperature for common RSW materials [20] Welding Parameters A characteristic welding schedule involves the following four successive steps: initial squeeze, application of current, hold, and electrode release (Figure 4) [21]. The heat generated under the electrodes from the concentration of current creates a molten weld nugget, due to a combination of contact resistance and bulk resistance. The molten weld nugget solidifies once the current ceases and cooling begins. The equation below, known as Joule s first law, generally represents how to calculate the amount of heat generated during the RSW process [20]. 8

29 Q = I 2 Rt (1) The total heat, Q, is calculated by multiplying the squared weld current (I), total circuit resistance (R), and weld time (t). Therefore, the key parameters for heat generation are the current, resistance, and time. The weld current and time are process variables that can be controlled with the RSW equipment. The resistance value is predominantly made up of the contact resistance at the faying interface and is dependent on the material, temperature, and welding force used. Therefore, the three variables that can be manually controlled and are often discussed in resistance spot welding are as follows: weld current, time, and force [22]. Recommended values for different material compositions and thicknesses can be found in RSW standards. Figure 4: Simple schematic showing the four phases making up the total weld time for the resistance spot welding process 9

30 Weld Current The welding current is the most influential variable for generating heat during welding. It is important to create a region of high current density through the electrodes to localize heating and melting in between the workpieces and under the electrodes. Current profiles can range from simple pulses to more complex upslopes and downslopes. Depending on the material and its weldability, a current profile is selected to widen the process window to create a more robust process. The use of sloping can be used in advanced high strength steels (AHSS) to localize heating at the faying interface. It is proposed that weld formation is based on not only the heat input, but also how the heat is applied, which has been a driving force in creating more complex current profiles for spot welding [23]. Until recently, single phase alternating current (AC) welding processes had been the major equipment used for automotive welding. With advancing technology, midfrequency direct current (MFDC) inverters are being more widely implemented for their precise control, improved reliability, and decreased power demands [24], [25]. Hofman et al. noted that AC versus DC had little effect on welding of thin gauge dual phase steels, but with thicker material the DC inverter gave larger weld nuggets and better mechanical properties [25]. Li et al. discussed that at low welding currents, there was a discrepancy in weld sizes between AC and DC processes but that the variation was decreased as current level increased to the expulsion point. The distinction in individual processes was linked to the contact resistance breakdown occurring differently due to the change in heating patterns [24], [26]. 10

31 Weld Time The weld time is usually chosen based on the material composition, coating, and thickness. The size of the molten weld nugget is based on the heat generated, which is directly proportional to the amount of weld time used. Short weld times can cause insufficient heat generation, resulting in a no weld [22]. Long weld times can produce excessive heat generation and nugget overgrowth resulting in expulsion. This can present possible defects like voids and unnecessary indentation that can have a negative impact on the mechanical performance of the weldment [27]. This is true for a nominal current level, but at low currents, longer weld times are typically required and at high currents, shorter weld times are used to compensate for the changes in heat generation [20]. Before current is even applied, a squeeze time is implemented to make sure the welding force has reached a steady state. This is important to ensure that the workpieces are in good contact with each other prior to the application of current. The hold, or cooling, time is another important variable post-current in resistance spot welding. The water-cooled electrodes help conduct heat away from the workpieces during the hold time. The enhanced cooling characteristics of the electrodes allow for quick solidification of the molten weld nugget to gain adequate strength [28]. In some cases, when martensite is formed in the weld nugget a post-weld heat treatment is needed to temper the weldment. Chuko and Gould describe how a sufficient hold time is needed to solidify the nugget and form martensite before applying a low current temper pulse to the weldment [29]. 11

32 For AC welding supplies, time is measured by cycles which are 1/60 th second for a 60 Hz North American machine. For DC welding supplies, time is measured in milliseconds. In manufacturing facilities, the weld time is of utmost importance because it can directly affect the cost and cycle time of production. In the cases of thicker or more advanced material compositions, more complicated multi-pulse welding conditions are used to create larger or tempered weld nuggets. While these conditions can enhance the quality of the weldment, they are often undesirable compared to a single pulsed weld condition due to the increased time [21] Weld Force Welding force applied by the electrodes is a unique feature to the resistance spot welding process. The electrodes utilization of force serves as in-process clamping and enhances the process by improving the joint resistance levels. The nature of weld force has been drastically altered due to the introduction of electric servo-guns. With pneumatic air-controlled guns, a single force was typically used for an entire welding process. Electric servo-guns, which use an electric motor and mechanical actuator to apply the force allow for shorter squeeze times, more precise control, and the use of higher force during the welding process [30]. It is now possible to apply multiple forces during different stages of the weld sequence to join complex stack-ups and advanced materials [12], [30], [31]. The force used during welding has a large influence on the contact resistance values at the interfaces. The welding force and contact resistance values are inversely related. A higher welding force creates more intimate contact between the workpieces 12

33 and reduces the contact resistance at the interface [32]. A reduced contact resistance can lower the heat generated during welding, which would have to be compensated by increased current or time. High electrode force is beneficial in reducing contact resistance but can also lead to excessive indentation from the electrode tips. This can result in unnecessary stress concentration and reduced mechanical properties of the joint. It can also lead to unwarranted wearing of the electrodes, which can increase the contact area and reduce the current density for future welds. On the other hand, too low of electrode force can cause premature expulsion before a sufficient weld nugget is formed Nugget Formation Each of the programmable process variables described previously must be cooperatively balanced to create suitable heat generation for welding. Welds are formed at the interface between the workpieces because the contact resistances there are much higher than the bulk resistances of the materials (refer to Figure 2). At the surface, asperities create an uneven surface that produces high resistance. As the asperities are broken down, the resistance falls to a minimum value once intimate contact is developed. The resistance of heating, increases the temperature, shown by stage 3 in Figure 5, until the liquidus temperature is surpassed. Slight decreases in resistivity are seen once the molten nugget is formed. 13

34 Figure 5: Schematic of the dynamic resistance developed during spot welding of uncoated steel [21] Due to the hidden nature of spot welds, it is difficult to analyze the formation process of the nugget. In literature, physics-based numerical analyses are completed to calculate nugget growth and temperature distribution [32]-[34]. Since the resistance welding process is electrically and thermal-mechanically coupled, it is difficult to analyze the results without comparison to actual testing samples. For experimental measurement, Cho and Rhee used high speed photography to analyze the nugget formation during welding. Simply explained, initial heating is seen in a rounded shape at the interface. The heating profile expands into a square shape, encompassing the area under the electrode. As temperatures continue to rise, a molten nugget begins to form in the center of the heated region and expands in all directions [35]. 14

35 In the automotive industry, materials are often coated for corrosion protection. The coatings can change the surface resistivity and behavior of the workpieces during RSW. In Figure 6, the dynamic resistance curve of zinc-coated steels shows a more complex profile than that of uncoated steels (Figure 5). During welding, the layer of zinc coating must be displaced to allow for intimate contact of the steel substrate. The soft, conductive nature of zinc allows for the coating to be melted and pushed to the edge of the contact interface. The change in the dynamic resistance profile alters the nugget formation. The added cycle time increases the heat generation and creates a more spherical weld nugget after solidification [36]. Den Uijl discusses that other coatings, like the AlSi coating often found on boron steels, decrease the conductivity at the surface and give different results during spot welding [37]. Figure 6: Schematic of the dynamic resistance developed during spot welding of coated steel [21] 15

36 2.1.3 Weldability While the set-up and design of resistance spot welding is relatively straightforward, there are many possible complications that can reduce the reliability of the weld. Three common issues that directly affect the resistance welding process and can lead to the welding defects seen in Table 1 are: improper control of welding parameters, material contamination, and electrode degradation [38]. The images in Table 1 show cross-sections of weldments. The hidden nature of the spot weld, in between the workpieces, does not allow for easy visual inspection and often requires special equipment and expertise for non-destructive evaluation techniques. The weldability, or capability of creating an effective weld, is important to study prior to implementing a welding procedure into production to predict how slight variation in welding parameters may affect the weld quality. In a high-volume automotive production facility, approximately 7 million welds are made per day. With this high volume, a goal of 100% spot weld quality acceptance is the aim because even 0.1% rejection would denote a high quantity of components needing repair [21]. 16

37 Table 1: Potential welding defects that can affect the performance of the weld [38] All welding parameters discussed in the previous section must be altered in unison to provide a resistance profile that allows for the joint to be welded. The resistance profile is effected by current, time, and force. When looking at effects of each variable, a window of optimal welding parameters is developed. The variables have a designated range that develops a sufficient nugget. In industry, the force value is often limited by the 17

38 equipment or more difficult to alter on a weld-to-weld basis and is held constant to simplify the window. Figure 7 shows a current range and how is it used to develop a weldability window, or lobe curve, for a given force. Current ranges show the weld growth proportional to the current. Multiple current ranges at different time increments can be put together to develop a lobe curve that gives an acceptable range of welding conditions. Slight variation that can shift the parameters will still produce a quality weld. It is more desirable to have materials with larger weld lobes as deviations from the designated currents will not significantly affect the weld quality, giving a large safety margin. The lower limit of the weldability window, represented by the left line in the lobe curve of Figure 7, is set by the minimum weld diameter that is deemed acceptable. Most North American standards use a minimum nugget diameter of between 4 t and 5 t, where t represents the thickness of the material [1]. The upper limit of the weldability window, represented by the right line in the lobe curve of Figure 7, is set by the expulsion point. Expulsion, the ejection of molten metal from the weld joint can lower the quality and strength of the joint. These limits can be determined by parameter development, or incrementally altering the heat generation to determine the critical current and time settings. The minimum weld diameter is characteristic of a low current density and the expulsion point is characteristic of a high current density. Optimum parameters are usually set just below the expulsion point where weld strengths are generally the highest [22]. 18

39 Figure 7: Example of a current range and lobe curve for resistance spot welding Every material, whether a steel, aluminum, coated, or un-coated, has its own response to welding and a unique resistance profile. The forming process of each material, like rolling and extrusion, creates slight irregularities that can also impact the resistance profiles of each weld. Figure 8 displays that by increasing the strength and alloying additions of the materials the lobe curve shifts to lower currents and decreases in width [1]. A wider lobe curve is more appealing to production because small deviations will not significantly affect the weld quality. 19

40 Figure 8: Different lobe curve profiles depending on material selection Slight misalignment of the electrodes or parts, variation in the surface conditions, and machine inconsistencies can further alter the resistance and current levels, presenting the importance of a robust weldability window. The electrodes are consumables to the process and wear over time. Wear can consist of mushrooming of the tip, reducing the current density and causing poor weld quality [20]. Proper cooling has shown to decrease the rate of electrode degradation. Paired with frequent tip dressing, process variation due to electrode geometry changes has been minimized in most industrial applications [39]. 2.2 Automotive Material Selection There are many different classifications for automotive steels that consider strength and metallurgical differences. In a broad sense, most steels fall into one of the following categories: low strength steels, high strength steels (HSS), or advanced high strength steels (AHSS). Advanced high strength steels are commonly broken down further based on yield strength levels. Yield strengths of 550 MPa to 780 MPa are 20

41 referred to as AHSS and yield strengths of 780 MPa and above are considered ultra-high strength steels (UHSS) [1]. Figure 9 compares the commonly used automotive steels based on percent elongation and tensile strength. Common metallurgical classifications of low strength steels are interstitial-free and mild steels. Conventional high strength steels (HSS) have a few metallurgical divisions including carbon-manganese, bake hardenable, and high strength low alloy steels. Typically, HSS consist of a single-phase ferritic steel. Carbonmanganese steels have the potential to form pearlite, but at minimal levels. The advanced high strength steels have different generations that meet the strength demands of certain automotive components [1]. These steels are uniquely designed to meet strength or energy absorption requirements for the key crash components of vehicles. First and second generation advanced high strength steels are metallurgically divided into the following: dual phase, complex-phase, ferritic-bainitic, martensitic, transformation-induced plasticity, hot formed, and twinning-induced plasticity. AHSS have more complex microstructures that are composed of phases other than ferrite, pearlite, and cementite like martensite, bainite, and austenite. Different combinations and the unique mechanical properties give AHSS more capabilities for greater strength and ductility compared to conventional HSS. Third generation advanced high strength steels are being researched. These steels are being developed to have an even better strength-ductility combination to further advance steel as an innovative solution for light-weighting and increased safety requirements. 21

42 Figure 9: Global formability diagram showing the range of properties of common steel classifications [1] The introduction of new materials in the automotive industry has presented some concerns pertaining to the process and performance of resistance spot weldability. Some process issues that have arose are smaller acceptable current ranges and decreased electrode life. The acceptable current range is generally limited due to the increased strength and hardness levels of AHSS. More complicated welding schedules are needed to grow the weld nugget to sufficient size. The electrode life has been reduced due to increased electrode forces needed and more complex geometries needed to provide ample heating Weld Characterization The resistance spot weld is hidden at the interface of the workpieces. While surface appearance, temperature discoloration, sheet separation, and indentation can give some insight on the weld quality, many costly defects are hidden within the nugget. In many cases, welds must be destructively tested to determine maximum loads and failure 22

43 mode prior to qualifying the welding schedule and stack-up as satisfactory for production [22]. By welding, a thermal cycle is induced on the workpieces, creating a weld nugget and heat affected zone (HAZ). The welding nugget exceeds the melting temperature, thus upon solidification a completely new microstructure is created. The HAZ does not reach temperatures above the melting, therefore the previous base metal microstructure is altered depending on the peak temperature. In gas metal arc welding (GMAW), there is a similar heating profile that creates HAZ sub-regions with different microstructural features and properties. In Figure 10, the peak temperatures and different regions are compared to the Fe-Fe3C phase diagram. In a GMAW low carbon steel weld, the HAZ created during welding can be divided into four main regions [40]. The first region, closest to the re-solidified weld metal is the grain growth zone, which sees peak temperatures above the AC3. The preexisting base metal microstructure is transformed into austenite and cooled back to room temperature. The peak temperatures are well above the AC3, allowing for austenite grain growth to occur. The second region is the recrystallized zone, which slightly exceeds the AC3 temperature. Since the peak temperatures are closer to the AC3, the amount of time at austenitic temperatures is less, resulting in a finer recrystallized microstructure. The last region experiencing temperatures above the eutectoid temperature is the partially transformed zone, which sees peak temperatures between the AC1 and Ac3 temperatures. The base metal microstructure is partially transformed into austenite, while the rest of the 23

44 microstructure is tempered. The last region, the tempered zone sees peak temperatures below the Ac1, which tempers the base material microstructure. While the RSW process is much different than the GMAW process, the resultant microstructural variation follows the same trend. The weld nugget consists of material that has melted and completely re-solidified, subsequently undergoing phase transformations. The base material composition and the cooling rate directly influence the microstructure produced in the re-solidified weld nugget [41]. The orientation of the grains is another important aspect of weld nugget solidification. The water-cooled electrodes create a high temperature gradient that align the solidification trajectory toward the electrodes, away from the weld centerline. Often, weld centerline defects (shrinkage voids, cracking, lack of fusion, etc.) can be significant weld nugget features. The HAZ, like in GMAW, consists of material that has not been melted but sufficiently heated to alter the base metal microstructure. It can be broken into unique regions, as described by Figure 10, but the span of each is much smaller. Resistance spot welding uses a high current density to create a localized process directly between the welding electrodes. The RSW joint is characteristically made between thin (<5 mm) sheet materials. Thus, the weld nugget spans an average distance of 3-9 mm. The HAZ region connects the weld nugget to the unaffected base metal and extends roughly 2 mm along the periphery of the nugget. 24

45 Figure 10: Schematic diagram of the HAZ sub-regions compared to the Fe-Fe3C equilibrium phase diagram [40] Metallographic testing is often used to measure weld nugget diameter, HAZ width, penetration, indentation, and detect possible weld nugget defects. Resistance spot welds are usually sectioned at the weld centerline, polished, and chemically etched to reveal the microstructural features [17]. The weld sizes and flaw measurements can be compared to standards to determine if the weld is acceptable for a given application. Hardness testing is then conducted on sectioned weldments. Hardness profile measures the response of a material to indentation and can be related to its strength. Hardness 25

46 profiles of an HSLA steel with three different AHSS is shown in Figure 11. The weld nugget has higher, or equivalent hardness values than the base metal. The fast cooling times and characteristic compositions of these materials infer that martensitic weld nuggets are forming. Depending on the material, the gradient between the weld nugget and unaffected base metal is different. Figure 11: Hardness profiles of a HSLA compared to three different AHSS [1] Hot Stamped Boron Steel Usibor 1500, a hot stamped boron-alloyed steel produced by ArcelorMittal, has high mechanical strength and good hot formability properties which allows for weight savings without loss in integrity [42]. The material composition, Table 2, allows for a fully martensitic microstructure upon quenching from the austenitic state, a common trait of UHSS. During the stamping process, complex geometries can be shaped in the high 26

47 temperature austenite regime (above AC3), with strength being regained after forming and quenching to martensite (Figure 12). The sheet material is often produced with a Al-Si coating to enhance the oxidation resistance during hot stamping and the corrosion protection. Table 2: Chemical composition of hot stamped boron steel [wt %] [43] C Mn P S Si Al Nb Ca < Ni Cr Mo V Ti Cu N B <0.003 < Figure 12: Timeline of the hot forming process [44] The hot stamped boron steels have been deemed weldable using the RSW process, but the significant reduction in hardness in strength has become an area of focus in literature. Jong et al. performed a study examining the mechanical properties and microstructural variation seen in resistance spot welding of Usibor It was found that the nugget diameter increased proportionally with increasing heat input, which also increased the strength in shear tension. In lower strength steels, the strength can be directly related to the nugget size [5]. In higher strength materials, such as Usibor 1500, microstructural changes can adversely affect the strength of the joint, creating a weak point where failure can occur at lower than expected loads. 27

48 The weld nugget and directly surrounding HAZ were composed of martensite with a hardness exceeding 500 Hv. A tempered zone was created adjacent to the unaffected base metal where a hardness drop of over 200 Hv was seen. The microstructure in the softened region was composed of tempered martensite and ferrite [6]. Vignier et al. have investigated the hardness profile of UHSS, looking particularly at the tempered microstructure where the reduction in hardness is seen. They have successfully applied the Rosenthal equation to determine the thermal history during welding and used the results to make hardness predictions along the weld profile [7] Dual Phase Steel Dual phase (DP) steels, like JSC 980 YL, exhibit high strength and are commonly used in automotive applications for weight reduction. The microstructure of DP steels contains islands of martensite surrounded by a ferrite matrix [45]. The strength of the grade, signified by the number designation, is proportional to the volume fraction of hard martensite, meaning the more martensite present in the microstructure the more strength the steel has. The soft ferrite allows for DP steels to have increased formability, allowing such steels to be shaped into structural components. Table 3: Chemical composition of JSC 980 YL [46] C Mn P S Si Al Nb Sn Ni Cr Mo V Ti Cu N B The composition of JSC 980 YL is shown in Table 3 and there is no coating on the surface. This grade has a lower carbon equivalent and high molybdenum content to produce a low yield ratio [47]. Cold rolled dual phase steels are produced by the 28

49 continuous annealing process. The steel is heated to the intercritical temperature regime and saturated to nucleate and grow the austenite grains. The entire part is then slowcooled to a quench temperature and then rapidly cooled to form martensite, producing the signifying dual phased microstructure [48]. Dual phase steels have been successfully welded in production for over ten years [10], [23], [25], [49]-[51]. The copper electrodes allow for quick heat dissipation and quenching of the weld nugget. In the heat affected zone, near the unaffected base metal, there is a region of base metal where the martensite is tempered, much like the hot stamped boron steel described earlier. In this case, the tempering of martensite can improve the overall ductility of the weld, but creates a region of localized necking where early failure occurs in tensile shear loading [51]. Interfacial failure, which is usually deemed unacceptable in lower strength materials is commonly seen in dual phase steels. The hard, brittle microstructure produced in the solidified weld nugget is susceptible to such failure [52]. Choi et al. found that interfacial failure predominantly occurs at RSW joints composed of dual phase and hot stamped boron steel. Combinations of stress concentration of the notch-like feature and the presence of voids within the weld nugget were commonly cited as the source of failure. The loading carrying capacity of the joints remains sufficiently high compared to industrial standards High Strength Low Alloy Steel High strength low alloy steels (HSLA), such as JSC 590 R, are commonly used for structural components in automobiles. Classified as an AHSS, this C-Mn steel has 29

50 improved formability and higher strength than the conventional mild steels. The goal of creating a weldable alloy with relatively low carbon and alloying additions started the development of HSLA steels [53]. Higher strength HSLA can replace lower strength steels with thinner gauges and added strength and energy absorption at impact [53], [54]. The composition of JSC 590 R is shown in Table 4 and there is no coating on the surface of the material. This material has added titanium and niobium to facilitate grain boundary pinning by carbides to create a finer microstructure [47]. Typical JSC 590 R is cold rolled and has a microstructure consisting of a ferrite matrix with islands of bainite acting as the strengthening phase. Table 4: Chemical composition of JSC 590 R [55] C Mn P S Si Al Nb Ca Co Ni Cr Mo V Ti Cu N B Sn JSC 590 R has been spot welded effectively with a single pulse weld schedule. Khan et al. compared the weldability and performance of 590 R to AHSS and conventional HSLA steels. They found that after testing, failure occurs in the unaffected base metal, providing insight that the RSW of this material is stronger than the unaffected base metal [3] Galvanneal Coated Mild Steel Thin sheet mild steels are commonly used for outer body paneling. These steels usually have a ferritic microstructure and have been widely produced over the past decade [1]. Mild steels are widely used in many different applications and are deemed weldable with the resistance spot welding process. 30

51 Table 5: Chemical composition of JAC /45 [56] C Mn P S Si Al Nb Ni Cr Ti Cu N B V For corrosion protection, the mild steel is dipped in a galvanneal coating. The galvanneal contains 9-12% bulk iron and % aluminum. The iron forms three zinciron phases, where the high-iron phase (roughly 15-28% iron) is near the steel substrate and the low-iron phases (5-6% iron) are closer to the surface. The aluminum is more uniformly spread throughout the coating thickness, but is more pronounced near the steel substrate where it is concentrated before the reheat cycle that allows for diffusion [57]. Khan et al. studied the weldability, microstructure, and mechanical properties of hot dipped galvannealed HSLA to DP steel. The differing compositions of the steels resulted in a smaller current range than each material in a homogeneous stack-up configuration, but still maintained a current range greater than 1 ka [58]. 2.3 Mechanical Properties and Testing Methods A key mechanical property of the material can be defined by its stress strain curve (Figure 13), which is generated from a uniaxial tension test. Many of the defining features of a material, like strength, ductility, and toughness can be taken from key features of its unique curve. Deformation of a metallic material can be elastic or plastic. Elastic deformation occurs as very small strains and can be completely recovered upon unloading. Plastic 31

52 deformation is permanent and occurs when the strain cannot be linearly related to the stress based on Hooke s law [20]. The stress at which deformation changes from elastic to plastic is desirable to know and usually used as a maximum in the design of structures. The exact transition point is difficult to accurately pin-point and can be estimated by measuring the 0.2% offset, commonly referred to as the yield strength. After yielding, the stress needed to continue plastically deforming the material increases until it reaches a maximum. The maximum stress a material can endure is also known as its tensile strength. For steels, the tensile strength is proportional to the hardness and a materials ability to withstand plastic deformation. At this point, necking begins and deformation is subjected only to the necking region where final failure occurs. Figure 13: Example stress versus strain curve of a metallic material, pointing out characteristic features 32

53 The degree of plastic deformation at failure is referred to as ductility. A material that can endure severe deformation is known as ductile, while a material that cannot is known as brittle. Ductility is an important trait of a material from a design perspective, giving insight of the degree of plastic deformation before fracture. For fabrication of sheet metal, it is often used to determine the allowable deformation during forming. For static testing conditions, the area encompassed under the curve can be described as the toughness of the material. For a metallic material to be tough, it must display both strength and ductility. Toughness represents the amount of plastic deformation and energy a material can absorb prior to fracture. The term toughness can also be used to describe a materials resistance to fracture when a preexisting stress concentration or flaw is present [59] Mechanical Testing Methods The demands of a resistance spot welded joint for a given application determine the needed performance level. The performance of the joint is often related to the weld nugget diameter, but the chemical and microstructural composition of the material, thermal profile, inhomogeneous material properties, residual stresses, and specific loading conditions can alter the performance behavior [60], [61]. The base metal microstructure before welding helps determine the strength of the material, but upon welding the base metal microstructure is altered considerably at the weld nugget and heat affected region [40]. As Callister explains, the mechanical behavior of a material reflects the relationship between its response or deformation to an applied load or force [59]. This relationship is severely altered during welding; therefore, it is important to use 33

54 mechanical testing methods to determine qualitative and quantitative weldability characteristics of resistance spot welds that meet the requirements of a specific application. Testing of welded samples is different than testing uniform material samples because of the geometry. Spot welds are characteristically considered a singular unit and the strength and elongation is displayed in force and displacement, respectively [20]. The unit of a spot weld consists of the weldment and the surrounding heat affected zone region. Therefore, the measurement of strength is not solely determined by the weld nugget, but is also influenced by the HAZ and base material. It is always necessary to provide information on the base material when discussing the strength of spot welded joints. To meet design criteria, the spot weld must exceed the strength of the weakest base metal being joined [17]. Different testing methods are used to replicate different loading conditions that could be seen in application. In the case of RSW for automotive use, testing parameters employ combinations of various loading rates and directions to mimic the crash analyses used. The test results are typically displayed as failure mode, weld button size, and weld strength [20]. The failure mode is a qualitative measure of weld quality and can give information on whether the failure was brittle or ductile in nature. The weld button size is a measure of the size of the button-like material that remains joined after destructive testing. An example of how the button measurement is taken is shown in Figure 14. The weld strength can be quantitatively measured in many ways. The peak load is the maximum force endured by the weld prior to failure. The ductility gives insight on the 34

55 energy that can be absorbed by the joint prior to failure. The fatigue limit measures the number of cyclical cycles to failure in a specific loading condition. Figure 14: Example of how to measure the weld button size post destructive testing [17] There can be confusion on what the weld size describes in resistance spot welding. The weld button (Figure 14) and nugget (material that has reached a molten state) width are often interchangeably used in describing the weld size, but different values are expected from each instance. Measurement of the weld button can be completed post-destructive testing but measurement of the nugget diameter requires cross-sectioning. The nugget diameter is more consistent and can be easily compared between material classes and welding parameters, but is not used as often as the weld button due to convenience [20]. Many different conglomerates are often interested in the mechanical behaviors of different materials and joints. Procedures for common mechanical tests are described in standards to define a consistency between companies and different market sectors. There are a wide variety of standards available which may have slight modifications but in general address a similar set of guidelines [16]-[18]. Due to the geometry of the spot 35

56 weld, testing procedures and specimen preparation are necessary to minimize the bending of the specimen. The stress concentration, or notch-like feature, of the weldment at the interface of the workpieces plays a large role in the deformation and fracture. From standardized testing methods, it is easy to obtain results that can be used to quantify the weld quality. But each manufacturer has unique joint design and quality control that can alter the acceptance levels of spot welds subjected to different loading conditions. If bending occurs, the original loading mode is not maintained during the entire test. The naming of the tests describes the original loading mode that is used in the procedure Tension Testing Tension tests are used to determine the strength and ductility of materials. There are three different tension tests used to assess spot weld quality: conventional tension, cross-tension, and U-tension. The peak load, maximum displacement, button diameter, and failure mode are recorded and compared to similar samples. The tension test can be implemented on a single sheet that is spot welded to a backing plate. The one sheet is pulled in tension until failure. The strength and width of the base metal have a large influence on the load able to be endured by the sample. A tension test can also be implemented at the weld centerline by eliminating out-of-plane warping in lap welded specimen (Figure 15). Often a U-brace is used to eliminate any rotation of the overlap region [62]. The cross-tension test is implemented on a cross shaped specimen loaded in the direction normal to the sheet interface, as seen in Figure 16. This test gives insight on the opening failure behavior of a spot weld. The two sheets are pulled apart from one 36

57 another, creating a large stress concentration at the notch-like feature at the interface. Special jigs (Figure 17) are needed to properly clamp the sample and minimize slipping. If the sample is not centered perfectly between the two opposing sheets, pure tension may not be attained. Ghassemi-Armaki et al. show that the cross-tension strength increases linearly with weld diameter and sheet thickness for DP980 [63]. Figure 15: Schematic of a tension test specimen [62] Figure 16: Schematic of a cross tension test specimen [17] 37

58 Figure 17: Loading configuration for cross tension testing [17] Radakovic and Tumuluru performed cross-tension testing on DP 790 and 980 to obtain a better understanding of the material behavior [64]. From finite element modeling and experimentation, weld sized from below 4 t to expulsion failed in the button pullout mode. The maximum stress is in the HAZ and is perpendicular to the loading direction. Radokovic and Tumuluru compared the cross-tension results to actual crash tested vehicle welds and found that while deformation seems similar, the actual loading condition is very different. The loads and failure modes seen in cross-tension testing could not be related to those seen in crash testing Tensile Shear Testing One of the most common testing methods used for RSW joints in industry and research is the tensile shear test. Unlike cross tension testing, the exact location of the weld is less significant to the results. The test can be conducted at a quasi-static or dynamic loading rate. Dynamic loading conditions require specialized equipment and are not very reliable and repeatable, often quantified with trends in data [3]. Therefore, to 38

59 evaluate the weldability of new materials and joint configurations, static tensile shear tests are often used [20]. The test specimens are welded in a lap configuration, as shown in Figure 15. Shims are often used in the set-up (Figure 18) to minimize bending that would deviate the loading mode from shear. The amount of deformation and rotation that is often seen in testing is dependent on the thickness and size of the weldment. Upon deformation, the nugget begins pulling form the base plate creating a moment. Tensile shear testing provides information on the ultimate strength and failure mode of RSW joints. Failure is primarily a factor of weld diameter, but can also be influenced by expulsion, defects, and heterogeneous microstructures. Figure 18: Loading configuration for tensile shear testing [17] Although the tensile shear test gives good indication of the welds strength and deformation behavior, it cannot completely simulate the loading conditions seen during crash testing. The tensile shear specimen is easy to fabricate, requires little fixturing, and 39

60 exhibits much less experimental variability, which are why this test is commonly used to validate RSW joints. Figure 19 shows the load versus elongation, typical tensile shear testing output, of single steel spot welds. The button, or weld nugget, size was 5 mm to allow for direct comparison between results. The higher strength materials have less elongation. Currently, new materials are being researched and developed to increase the ductility while still maintaining the high strengths. While strength is important, without ductility, the failure mode is typically brittle which correlates to very little energy absorption at failure. Figure 19: Loading versus elongation of single spot weld steel specimens subjected to tensile shear loading [1] Chao has shown that while the lap-shear test specimen is subjected to a shear load globally, the failure mechanism of the weld at the microstructural level is tensile [60]. He 40

61 has also shown that the cross-tension test is subjected to a normal load but the failure mechanism seen is shearing. This explains why the cross-tension test often sees a lower overall load than the tensile-shear test for the same given weldment and base material Pry-Checking/Chisel Testing Once welding conditions are verified and implemented in production, on-going quality assurance is needed. Pry-checking is considered a non-destructive method that gives insight on no welds versus good welds during production assembly. Often in production, a screw-driver or chisel ( 11 angle) is inserted at the interface between two welds to ensure the bond. Tsai et al. has studied how material properties, pry angle, and pry distance can cause a weld stress that may cause damage to good welds [65]. It was found that the parametric effects are more sensitive to strain variations due to the large plastic deformation seen during testing Wedge Testing Wedge testing is a novel spot-weld testing procedure developed by Payen et al [66]. The procedure mimics the opening loading mode of the cross-tension tests but can be completed in the confined space of a scanning electron microscope (SEM) chamber. The purpose of the wedge test is to be able to observe the crack initiation and propagation of an RSW sample in situ. Due to the symmetry of a spot weld, tests were constructed on spot welds sectioned at the weld centerline. The sample was polished, etched, and placed into the SEM chamber with two tool steel wedges, as seen in Figure 20. The wedges are inserted at the interface of the weld to a predefined load and displacement to observe crack growth. 41

62 Figure 20: Example of set-up used by Payen et al. for wedge testing [66] The developed testing method is like cross-tension testing but gave the ability to observe in situ crack propagation of the DP780 spot welded sample. The direct observation of strain fields gives information on the differences between interfacial and button pull-out failure modes and the fracture strength of the different region of the joint. Failure seen in the region between base metal and heat affected zone shows a ductile failure mode that causes decohesion of martensite-ferrite interfaces. When failure occurs in the fusion zone, it is brittle and very small inhomogeneous local strains are seen [66] RSW Failure Modes Many company, national, and global procedure and specification manuals determine weld quality based on visual inspection of the fracture. The automotive industry uses the eight descriptions in Figure 21 to describe the fracture mode at the joint 42

63 level [1]. Of the six, the three fracture modes commonly referred to are as follows: button pull out, partial button pull out, and interfacial failure. Figure 21: Common fracture modes seen in the automotive industry [1] Button pull out fracture occurs when the joint itself does not fail, but the HAZ or base material at the border of the joint fails. Failure in the HAZ occurs when the welding process has weakened the HAZ by tempering the microstructure or excessively growing the region. Failure in the base metal occurs when the welding process has strengthened the weld above the base metal strength. Partial button pull out fracture occur when the RSW joint cannot endure the load. Part of the weld nugget remains intact, but the crack propagates through a portion of the nugget. There are many different variations depending on the route of the crack, as seen in Figure 21, but these are usually lumped 43

64 together as partial button fracture. Interfacial fracture occurs when failure is in a straight line along the interface of the workpieces. Button pull out fracture is deemed an acceptable failure mode and gives evidence that the welding process did not affect the materials ability to withstand loading. There is often more part deformation, indicating more energy absorption at failure, when button pull out is the fracture mode. At a material level, failure is usually ductile in nature and exhibits localize necking of the material surrounding the nugget. Partial button pull out fracture is indicative of the welding process altering the materials ability to endure a load. The joint is the weakest link and a crack propagates partially through the solidified weld nugget. There are many different factors that can contribute to partial button pull out and it is often deemed an unacceptable fracture mode for production. With the advancements of materials, and brittle martensitic weld nuggets, partial button pull-out can be deemed acceptable in situations where a maximum load is reached prior to failure. Interfacial fracture leaves little to no warning signs before complete propagation along the weld centerline in a brittle fashion. There is often no part deformation at failure, indicating very low energy absorption. In the past, interfacial fracture was always deemed unacceptable, but like partial button pullout has been qualified as acceptable for some advanced high strength materials. Zhang discusses many experiments that collectively show a strong relationship between failure mode and weld size [20]. This relationship is not clearly defined in common RSW specifications, but is implicitly stated as smaller welds tend to fail 44

65 interfacially while large welds tend to fail as button pull out. The scale between small and large welds continuously changes based on material thickness and strength. The safety demands of a car rely on the assembled parts. Many crash specific features, like door beams for impact safety in side collisions, are spot welded. The interaction between the type of loading, weld joint geometry and material properties makes it difficult to determine the mode or location of failure in advance [67]. In an ideal situation, the weld would never fail and material properties could be used to define failure of the unaffected base metal. From the standardized testing methods and tests described previously, a lot of information can be gathered about the failure mode and strength of RSW joints. But, these tests individually cannot simulate the complex loading conditions that are seen crash testing. As materials get stronger and more complex, the failure modes become more compound. With heterogeneous weld and hardness profiles (see Figure 11), failure studies must consider the changing properties in the weld nugget and heat affected zone, compared to the base metal. Softening phenomenon in the HAZ can lead to a decrease in peak load and can raise concerns for a joints ability to absorb energy in crash related incidents. Though button pull out would still be the failure mode in this instance, literature is discussing the importance of looking at the material and joint failure contributions to determine the efficiency of a joint [2], [3], [67], [68] Digital Image Correlation Digital image correlation (DIC) is a non-contact technique that uses image processing algorithms to determine relative displacements of material points to measure 45

66 material deformation [69]. The core computation algorithms have been improved in recent years to increase the usage among researchers and industry alike. The accessibility to open-source 2D DIC software, like Ncorr, has allowed researchers to continue innovating the modern algorithms to better suit new challenges faced in industry and materials science. Blaber et al. gives examples of materials such as biological, metal alloy, porous, polymer, and shape memory alloys that have been studied using DIC techniques. Each material system requires its own unique modifications, but the core computational techniques remain the same. DIC uses algorithms to register relative points of material displacement between an untested reference image and a deformed current image. The breadth of DIC for various material systems is due to its lack of scale sensitivity. Researchers can study deformation at different length scales, from meters to nanometers, if proper pattern and imaging practices can be executed [70]. The reference image is partitioned into subsets that are small enough that deformation is uniform within each subset. Each subset is tracked from the reference image to the sequential images. In Ncorr, an open source 2D DIC MATLAB software, the subsets are originally circular groups of points with integer pixel locations in the reference image. As the region is deformed, the transformation of the subset coordinates is expressed as a linear, first order transformation. Coordinates are taken from the reference image and parameterized by displacements and their derivatives to correspond with the current image of interest. The exact equations used can be found in the core DIC algorithm description by Blaber and are described in Figure 22 [69]. 46

67 Figure 22: The basic steps of the DIC algorithm implemented in Ncorr using an initial guess and iterative optimaization scheme to find a refined solution [69] Harilal and Ramji conducted testing to determine the accuracy of Ncorr, an open source DIC software, compared to commercial DIC packages [71]. The random pattern used was created by applying a thin layer of white paint across the surface of the sample, followed by an airbrushed black paint to create a random black and white speckle pattern. The random pattern created by the disorganized nature of the speckle is what is used to track the displacements of the sample. Two different geometries and loading conditions were used to determine the correlation between Ncorr and the most widely used commercial software, Vic 2D. It was found that the displacements and strain values closely matched one another as seen in Figure 23 and Figure 24. The displacement is taken along a line, to show that the displacement values are accurate across the entire span of the sample. The strain is shown in map form and shows very similar trends between the Ncorr, free software, and Vic 2D, commercial software. 47

68 Figure 23: Comparison of u (x) displacement along a specific line for Ncorr and Vic 2D Figure 24: Comparison of strain (ɛxx) contours obtained for a compressive force on an epoxy ring for Ncorr and Vic 2D 48

69 2.4 Crash Simulation Techniques Automotive companies use commercially available finite element software to simulate the situational crash testing that is required by the National Highway Traffic Safety Administration and Insurance Institution for Highway Safety. Due to the complexity of the entire simulation, spot welds are often simplified for full vehicle FEA models. Currently, RSW joints are modeled by flexible springs, bars, rigid elements, and multipoint constraints. Advancements and increased availability of high-performance computing systems has allowed for larger, more detailed models to run efficiently. Designers and analysists have set new goals of increased accuracy, correlation, and validation between crash simulations and actual crash testing. Improved accuracy is based on the demand for developing ways of precisely simulating the thousands of spot welds present on a vehicle, without over complicating the entire FEA model [72]. As previously mentioned, there are many ways of modeling spot weld connections in commercial FEA software. These techniques do not require detailed information about the joint and simplify the analysis. This is a beneficial practice because it drastically cuts down on the computational time since there are thousands of spot welds present on a single vehicle. Current crash models use a shell model with a very coarse mesh. During crash testing and simulation, the spot weld plays a very integral role in the performance of the vehicle. The National Crash Analysis Center (NCAC) has created FEA models of vehicles for crash simulation and conducted validation studies during the 49

70 development of the models to determine the performance and accuracy of various techniques. They have found significantly different results in simulations and actual testing when as little as three spot welds are removed from key locations in the vehicle [73]. Lamouroux et al. discuss that the main source of inconsistencies between numerical and experimental crash behavior is due to the inaccurate predictions of the mechanical and fracture behavior of joined structures [74]. For efficiency and to simplify spot weld connections in crash simulations, rigid node-to-node connections are used to join components. Unfortunately, this type of connection does not consider the amount of detail needed to accurately make predictions with the newer higher strength materials. These materials have a distinct base metal, heat affected zone, and weld nugget that behave differently and effect the overall behavior of the joint. Lamouroux et al. has used hardness testing on cross-sectional specimens to estimate the mechanical behaviors of the different regions [74]. Using the respective mechanical behaviors, they could obtain good correlation between experimental and modeled force displacement curves for different geometries. Wung discusses the use of a force-based failure criterion that can be directly related to stress-based criterion for a specific thickness and weld nugget [75]. Destructive testing must be completed to obtain the failure properties of a spot welded joint. Different loading angles (Figure 25) and rates are used to determine static, dynamic, and impact behaviors of crash-like influences [76]. Simulations of such tests have given more information on the behavior of spot welds and by including this information into crash models can attribute to an even greater accuracy. The thermal 50

71 cycle during welding creates a heterogeneous weld profile and overload failure for ductile materials involves large strains and geometry changes. These variables and conditions create a complex task of predicting failure in resistance spot welds. Zuniga and Sheppard performed experiments to characterize fracture, strain hardening, and strain rate sensitivity of HSLA steel RSW joints [67]. Figure 25: Different loading cases [76] Palmonella et al. used three different spot weld modeling techniques to see their applicability to crashworthiness applications [77]. The first technique implemented a single element connection, which is the simplest discretization for FE spot weld modeling. The second technique uses a single hexahedron element connection. The 51

72 stiffness of the single element depends on the position of the connection nodes in regards of the shell surface and cannot account for any torsional loading. The third technique uses four hexahedron elements to simulate a connection. This necessitates small element sizes, which decreases the time step needed for numerical stability. This could then lead to additional mass scaling to allow for larger time steps, but would inhibit the accuracy of the simulation. While these methods discuss how these methods compare in regards to simulation times and model size, they do not discuss the comparison of the simulation to actual crash testing. 2.5 Complex Stack-Up Configurations The most common joints on an automotive body are resistance spot welds that adhere two different sheet materials together. Due to complexities in structural design, equipment access to joint locations, and light-weighting demands there are many applications where it would be beneficial to join dissimilar materials of differing thicknesses, making joints composed of two, three, or even four sheet materials together with a single weld [78]. Increased efforts have been made to research potential methods for applying complex stack-ups to some of the critical structures seen in Figure 26, which include front longitudinal rails, bulkhead to inner wings, A-, B-, and C-pillars [79]. 52

73 Figure 26: Some of the critical structures that must have adequate strength for passenger crash safety As described previously, the most basic form of resistance spot welding is a twosheet combination of similar thicknesses. When materials of two different thicknesses are joined, the heat conducted away from the joint by the electrodes is unequal for each sheet. The thermal profile is altered and the melting will occur in the thicker material (less heat sink from the electrode), away from the interface. If not enough heat is generated at the interface, lack of fusion can occur. Researchers have discovered ways to alter the welding schedule and electrode properties to develop more uniform thermal profiles that generate peak heating at the interface. With many standards and 53

74 recommended procedures developed for 2T welding, an increased interest has been developed for 3T and 4T welding. These multifaceted stack-ups involving three and four sheets present more interfaces that add complexities to the resistance welding process. With added sheets and interfaces, one common problem faced is insufficient nugget growth due to insufficient heating of the large stack-up. With the locations of these multilayered welds being structurally important, large weld nuggets are needed to provide adequate mechanical properties to the joint. The literature has presented information regarding the nugget growth of 3T welds and the failure behaviors associated with these joints [10], [12], [80]. Very few organizations have information on 4T welding [11], [49], [78]. Generally, guidelines for 3T and 4T resistance spot welding are based on rule of thumb and trial and error results. Standards and common procedures are not available to help determine process parameters that create proper nugget growth and failure behavior for industrial application. Coon et al. looked at a hot dipped galvanized DP steel stack-up consisting of a thin (0.8 mm) sheet joined to two thick (1.9 mm) sheets [10]. They looked at the effect of different electrode geometries (Figure 27) on the weldability and endurance of the welded joints. The dome shaped electrodes produced a larger current range for the given stack-up using an AC power supply. The dome shaped electrodes were also able to stabilize more quickly, due to the higher current density (smaller face diameter), which allowed for 54

75 more consistent welds and less power consumption. For this specific stack-up, pulsing had no effect on the overall acceptable current range. Figure 27: Electrodes used by Coon et al. for experimental testing [10] Gould et al. used a dissimilar stack-up consisting of thin (0.7 mm) galvanneal 270 MPa steel joined to two thick (2.0 mm) galvanneal DP 590 steel. They determined a nominal style of welding schedule, as seen in Figure 28, and altered the process and electrodes to determine the most efficient manner of welding the 3T configuration. The nominal measure of weld quality, per industry usage, is based on the nugget penetration into the thin sheet. This measure is most sensitive to many process and electrode variations. The easiest way to enhance the nugget penetration into the thin sheet was to use electrode variations to alter the heat balance and pull the nugget toward the thin sheet material. Though the effect is much more pronounced, changing electrode diameters and compositions is not a viable solution for industrial application due to the mass production of resistance spot welds. Gould et al. developed a welding schedule that would result in 50% penetration into the thin sheet. The welding schedule was short and consisted of a single pulse before forging followed by a higher force, higher current forge cycle to complete the joint [12]. 55

76 Figure 28: Sequencing of weld current and force used to create a nominal welding schedule by Gould et al. [12] Pouranvari and Marashi completed an in-depth study on the nugget growth of 3T stack-ups. Experimentation was conducted to determine the effect of material thickness in 3T stack-ups of low carbon steel. Welds made on a 3T stack-up of steel thinner than 1.5 mm, the nugget formation began at the geometrical center of the stack-up. The nugget would then grow radially outward, but was largest at the center point. Welds made on a 3T stack-up of material thicker than 1.5 mm exhibited nugget formation beginning at the sheet to sheet interfaces. Individual nuggets would grow at each interface, creating slightly larger diameters at the interfaces rather than the geometrical center point. This concept can be visually seen by the schematic in Figure

77 Figure 29: Depending on sheet thickness the nugget formation begins at the geometrical center of the stack-up or the sheet to sheet interfaces Investigations by Jung et al. [78] have considered joining three and four sheets by RSW, but only analyzed the mechanical behavior of the joint and did not consider the nugget growth behavior. Wei et al. [49] completed a study looking at the nugget size and tensile shear behavior of RSW 2T, 3T, and 4T stack-ups. Their work discussed the welding parameters for each joint configuration but the individual sheets were all homogeneous in composition and thickness. Eizadi and Marashi [11] investigated the nugget development and failure behavior of a dissimilar 4T RSW stack-up. The stack-up consisted of a thin (0.7 mm), thick (1.2 mm), thick (1.2) mm, thin (0.9 mm) configuration where all sheets were low carbon steel. When welding a 4T stack-up many problems with nugget formation occur due to the increased number of interfaces. The bonding mechanism seen in resistance spot welding is directly related to the heat input. Eizadi and Marashi showed that welding currents above a critical threshold created an acceptable bond at the middle interface (between the two thick sheets), but it was difficult to get fusion between the thick/thin interfaces on each side [11]. In some instances, a solid-state metallurgical bond was found at the thick/thin interface. Since the thin sheets of materials are on the outside and 57

78 in direct contact with the water-cooled welding electrodes, it requires a higher input to overcompensate for the quick heat dissipation through the electrodes. Thin materials are often used for body exteriors, thus when creating complex stack-ups are typically found on the outside of the configuration. The difficulty of joining a thin material to thick materials was also addressed as a problem in 3T welding. It is of an utmost importance to determine critical current levels that allow the nugget to penetrate the thin sheet to achieve an acceptable joint integrity. 58

79 Chapter 3: Objectives The overarching goal of this work is to predict the performance of resistance spot welded ultra-high strength steel in conventional and complex stack-ups. To be able to make reliable predictions a thorough understanding of the hot-stamped boron steel metallurgy, material constitutive behaviors, and joint performance is needed. The work presented in this thesis is considered preliminary and must be continued to complete the overarching goal. The specific research tasks are listed in greater detail as follows: 3.1 Process Optimization and Characterization Trials will be conducted to determine conditions to weld a two sheet Usibor 1500 to Usibor 1500 stack-up. A window will be determined in which welding achieves quality nugget size and strength. Samples will be sectioned and prepared for metallographic characterization to observe microstructural variances between the base metal, weld metal, and heat affected zone. Special interest will be taken in analysis of the coarse-grained HAZ and subcritical HAZ. Micro-hardness mapping will be used to measure the hardness distribution of the weldment. Quantitative microstructural information gained will be used as a baseline comparison to simulative HAZ microstructural samples produced using the Gleeble thermal-mechanical simulator. 59

80 3.2 Constitutive Modeling Bulk microstructural samples will be produced using a Gleeble thermalmechanical simulator to mimic the various HAZ regions of resistance spot welds made of Usibor A special interest will be taken in the coarse-grained heat affected zone and subcritical heat affected zone. The bulk microstructural samples will be compared to the microstructural region of the weldment qualitatively with microscopy and quantitatively with hardness data. Once bulk samples provide an adequate simulation of the weldment regions, quasi-static tensile testing will be conducted to determine the unique microstructural material properties. Digital image correlation (DIC) will be implemented on the samples to provide virtual extensometer data that can be correlated into strain. The constitutive behaviors of the unique microstructural regions will then be implemented into FEA. 3.3 Weldment Mechanical Testing Mechanical testing will be conducted on Usibor 1500 resistance spot welded specimens to determine the performance of the joints under multiple loading conditions. Testing will be conducted at the coupon level and new techniques will be developed to use in situ digital image correlation to determine the localized mechanical behavior of the differing regions of the weldment. The failure type, peak loads, strain distributions, and crack propagation path will be analyzed to determine what causes different failure modes. Conventional testing methods, like tensile shear testing, will be conducted and altered to gain a better sense of the properties of the joint. New innovative testing methods will be 60

81 developed to give a better overall understanding of the differences between failure modes in resistance spot welded joints. 3.4 Weldment Computational Modeling Coupon level, fine-resolution, computational models will be constructed using the commercial Abaqus FEA package. The models will incorporate the constitutive behaviors found by the bulk Gleeble simulated to account for the heterogeneous profile of the weldment. Coupon level models will be used to compare directly to the mechanical testing methods developed. 3.5 Complex Stack-Up Configuration Materials of interest consist of Usibor 1500, JAC 590 R, JAC 980 YL, and JAC /45, in an assortment of stack-up configurations. The thicknesses of each material are different, complicating the joint further. Welding conditions that can effectively weld the four materials in 2T and 3T stack-ups and are suitable to a production setting are to be developed. Metallographic analysis will be conducted to determine the nugget sizes at each interface. 61

82 Chapter 4: Ultra-High Strength Steel Characterization This chapter presents details of the work that was completed to better understand ultra-high strength steel and resistance spot weld characteristics. A metallographic investigation was used to determine the unique microstructural regions of the joint and mechanical testing was used to develop the constitutive behaviors of each. For completeness, results generated by Ying Lu are included in this chapter. 4.1 Materials Usibor 1500, which had previously been hot stamped, was the material of interest for characterization. Welds were created between two sheets of 1.5 mm thick Usibor, coated in aluminum silicon. Gleeble microstructural simulations were also created on 1.5 mm thick Usibor, but with the coating removed prior to heat treatment. 4.2 Approach Process optimization was first performed to determine the specific microstructural regions developed during welding and relative locations to the weld nugget. From microstructural analysis, the different regions were further investigated to determine Gleeble parameters to duplicate each microstructural region in bulk. Bulk microstructural samples were then mechanically tested to determine individual constitutive behaviors of the respective regions. 62

83 4.2.1 RSW Process Optimization RSW tests were first conducted on a Taylor-Winfield single-phase 60 Hz alternating-current (AC) resistance spot welding machine with a Medar control unit. The electrode tips used were dome-shaped RWMA class II with 6 mm diameter face. The two sheet (2T) Usibor to Usibor welding schedule was developed using the ISO standard (ISO :2004) [16]. Destructive testing, including sectioning and lap-shear testing, were used to measure the nugget size and joint strength, respectively. It was established from the American Welding Society Standards (AWS D8.1M) that the optimal welding parameters were those that exceeded minimum values for nugget size and joint strength [17]. The minimum acceptable nugget sizes for 2T stack-ups can be found in Table 6, where the 1.5 mm thickness is highlighted. While the AWS minimum nugget size is 5.0 mm for the given thickness, many industry standards have a more rigorous requirement of 5 t as the minimum nugget diameter. The minimum strength requirements for the given thickness, per AWS standards, is marked with a green dotted line and measured with tensile shear testing (Figure 30). The optimal welding parameters that met both minimum nugget diameter and strength requirements were used to fabricate duplicate samples for subsequent microstructural characterization. The parameters were carefully held constant for the entirety of experimentation to minimize any variation of temperature distribution and final microstructure in the HAZ [81]. 63

84 Table 6: Minimum Acceptable Weld Size taken from AWS D8.1M [17] Figure 30: Minimum tension shear strength requirements taken from AWS D8.1M [17] Sample Preparation The resistance spot welded samples were sectioned transversely, through the center of the weld nugget, and mounted in bakelite powder with a Leco PR-32 mounting press. Mounted samples were polished using 240 to 800 grit SiC metallographic paper lubricated with water. Following the grit procedure, 3- and 1-micron diamond paste 64

85 lubricated with diamond compound extender was used to get a mirror-like finish. For scanning electron microscopy, a 0.5-micron colloidal silica suspension was used as a finishing polish for a nearly scratch-less surface. Between each polishing step, samples were sprayed with ethanol, cleaned in an ultrasonic bath for 2-3 minutes, sprayed again with ethanol, and dried with a hot air source. The polished sample were etched with 2% Nital, for seconds depending on the desired degree of etching Microstructural Analysis Optical microscopy was conducted using an Olympus GX51 microscope. This type of analysis could be used for visual determination of nugget size and broad comparison between samples. A Quanta 200 scanning electron microscope (SEM) was used to characterize the various microstructures of a weldment, including the weld metal, base metal, and various regions of the heat affected zone. Micro-hardness mapping was completed using a Leco micro-hardness tester LM100AT with a 0.5kg load and 300 µm grid spacing [14]. The mapping was performed in two different orientations, as depicted by Figure 31. The first map was completed on the conventional transverse plane (XZ), and the second map was completed on the longitudinal plane (XY), parallel to the steel surface. For the latter, the spot weld was grinded below the electrode indentation to expose a profile of the weld nugget near the interface of the 2T stack-up, marked by the red dotted line in Figure

86 Figure 31: Example of the surfaces used for micro-hardness measurements Samples were placed in the fixturing device, leveled flat, and locked into place with a set screw. The surface was mapped and indentation paths were set to analyze hardness variations across the profile of the weldment. Indentations and hardness readings were automatically generated and exported to an excel file containing the hardness value and location (in x and y coordinate format) of each indent. A color-coded map was developed from the information gathered to determine hardness trends and values across the sample Gleeble Simulation The microstructural simulations were completed by Ying Lu using the OSU Welding Engineering Gleeble thermo-mechanical simulator. The differing regions of the heat-affected zone are of interest to determine the mechanical behavior of the joint. To generate constitutive behaviors of the individual regions of the heat-affected zone, where potential failure may occur, bulk microstructural samples were created using the Gleeble Samples were clamped with copper or stainless steel grips and heated resistively. A K-type thermocouple was percussion welded at the middle of the sample to monitor the temperature profile. 66

87 For the CGHAZ simulation, half contact stainless steel grips were used to clamp the samples. This allowed for a slow and uniform heating rate of 100 C/s to a peak temperature of 1300 C. The temperature profile can be seen in Figure 32a. For the SCHAZ simulation, copper grips were used to clamp the samples. These grips allowed for a high heating rate of 2000 C/s to a peak temperature of 700 C, as seen in Figure 32b. The SCHAZ sample was held for 0.5 seconds. After heating to and holding at peak temperature, both the CGHAZ and SCHAZ samples were quenched through sprayed water. This method of cooling can exceed rates of 2000 C/s, which simulates the watercooled electrodes used during spot welding. Figure 32: Temperature profiles of physically- simulated (a) CGHAZ and (b) SCHAZ Microstructural Tensile Testing To determine the constitutive behaviors of the different microstructural regions, tensile tests were conducted on the bulk Gleeble samples. The Gleeble specimens where machined to the geometry shown in Figure 33. Tensile testing was performed using an MTS 810 which applied a quasi-static loading condition on the sample. 67

88 Appendix A provides sample calculations on determining the stroke rate of the MTS system. Due to the small size of the sample, digital image correlation (DIC) was used to determine the local displacement and strain distribution. A fine coating of white spray paint is applied to cover the entire cross-sectioned surface. Once the white spray paint is dry, a random black speckle is applied to the cross-sectioned surface and allowed to dry before testing. A Nikon D7100 camera was focused on the speckled pattern and used to record the testing. Images were extracted from the video and analyzed using Ncorr and Post_Ncorr, open-source Matlab software programs [69]. A 2-mm-long virtual extensometer was used during analysis for calculate the strain localization in the sample via the displacement. Figure 33: Tensile sample geometry [mm] The displacement data for each image was fit to a general curve and used to determine the engineering strain accumulation during the tensile test. The MTS system measured the force accrued during testing. Engineering stress is the quotient of the measured force divided by the initial cross sectional area of the reduced section. By compiling the engineering stress and strain data, constitutive behaviors of the materials were developed. The engineering stress and strain are more convenient to determine with 68

89 experimental procedure, but the material behavior is better described by the true stress and strain values [82]. The engineering stress and strain values found by experiment can be converted into the true values by using the equations below. σ true = σ eng (1 + ε eng ) (2) ε true = ln(ε eng + 1) (3) 4.3 Results and Discussion During welding, microstructural changes occur to the weldment that can alter the properties of the weldment and surrounding material. The RSW process optimization allowed for reproducible weldments that met the AWS requirements. These welds were characterized metallographically and the individual regions were reproduced using Gleeble simulation to gain information on the constitutive behaviors RSW Process Optimization Per the ISO standard (ISO :2004) [16], the welding schedule consisted of the recommended process parameters listed in Table 7. The squeeze cycles are used to ensure that the electrode force is at the recommended value before welding. The hold cycles are used to cool and quench the weldment. The actual welding parameters consist of three 11 cycle pulses that are separated by 2 cycles of cool time with a 450 kgf weld force applied (Figure 34). 69

90 Table 7: Recommended process parameters for 2T 1.5 mm stack-up Squeeze 70 cycles Weld 3 x 11 cycles Cool 2 cycles Hold 20 cycles Electrode Force 450 kgf Figure 34: Process parameter visual showing the current, force, and time By using these parameters as guidelines, the amperage was varied to determine a current range that produced acceptable weld nuggets. The three pulses were altered in unison, meaning that the current values were identical for each of the three pulses. Current values were varied from 4.3 ka to 6.9 ka. In Figure 35, the nugget diameters for the trial are given and the minimum nugget diameter, 5.0 mm, determined by AWS is indicated by the red dashed line. Current levels below 5.25 ka gave nugget diameters below the AWS minimum value. Sporadic expulsion was seen with welds at or above 6.4kA. While these welds still give adequate nugget size, expulsion creates much greater variability. 70

91 The acceptable current range that was considered from testing is 5.5kA to 7.0 ka. To ensure that all cross-sectioned welds experienced the same heating and cooling cycle, welds for experimentation were made at 6.5 ka. To ensure that adequate strength was met, welds made at this current were subjected to tensile shear testing. On average, the loads experienced were 27 kn, which is well above the minimum value per the AWS standards (Figure 30). The samples failed by button pull-out. Figure 35: Current range for 2T Usibor RSW UHSS Metallographic Characterization The joint consists of many different regions, depending on the heating and cooling cycle seen at a given location. The typical cross-section of the 2T Usibor 1500 weldment can be seen in Figure 36. The unique metallographic regions that form during 71

92 welding are as follows: weld nugget, coarse-grained heat affected zone, fine-grained heat affected zone, intercritical heat affected zone, and subcritical heat affected zone. Figure 36: Cross-section of the Usibor T stack-up with labels identifying the different microstructural regions The peak temperatures are highest at the weld nugget and well above the melting temperature, as depicted by the welding temperature distribution profile (Figure 37). The water-cooled copper electrodes act as sinks, extracting the heat from the joint rapidly after welding, thus quenching the material directly in contact with the electrodes. The molten weld nugget solidifies rapidly, transforming into austenite at elevated temperatures which in turn completely transform into martensite (much like the transformation seen in the base metal during hot stamping). For comparison, the unaffected base metal and weld metal microstructure are shown Figure 38(a) and (b). Both are composed of a fully martensitic microstructure with fine martensite laths. The material surrounding the re-solidified weld nugget comprises various heat affected zone regions. The heat affected zone experiences a dynamic heating and cooling 72

93 sequence that alters the microstructure without being melted. In the coarse-grained heat affected zone, CGHAZ (Figure 38c), there is austenitic grain growth due to an extended period spent above the AC3 temperature. The CGHAZ is directly surrounding the resolidified weld nugget and experiences temperatures slightly below the melting point of the material. In the fine-grained heat affected zone, FGHAZ (Figure 38d), which is further away from the weld nugget, peak temperatures are lower, thus reducing the amount of time available for austenite grain growth. Due to the rapid cooling rates encountered in RSW [7] and high hardenability of the base metal steel, all regions shown in Figure 38 allow austenite to transform completely into martensite which is supersaturated with carbon. Figure 37: Microstructural regions formed as a result of temperature distribution of the weldment during the resistance spot welding process 73

94 Figure 38: (a) hot stamped martensitic base metal (b) re-solidified martensitic weld nugget (c) martensitic coarse-grained heat affected zone (d) martensitic fine-grained heat affected zone The heat affected zone also encompasses regions of material that are heated below the AC3 temperature. The first HAZ region to be subjected to such temperatures is known as the intercritical HAZ (ICHAZ), that is partially transformed austenite and partially tempered martensite. The ICHAZ serves as a transition from the recrystallized martensitic microstructure of the weld metal, CGHAZ, and FGHAZ to the over-tempered microstructure of the subcritical HAZ (SCHAZ). 74

95 The microstructure of the SCHAZ, located immediately adjacent to the unaffected base metal is shown in Figure 39. It is composed of ferrite grains and cementite precipitates; the latter decorate along the prior austenite grain boundaries and the interlath regions of prior martensite. The formation of the SCHAZ is due to over-tempering of martensite, consisting of the decomposition of metastable martensite into ferrite and cementite when heated to peak temperature below the AC1 temperature [45]. The tempering is often detrimental to the strength and hardness that is often associated with a martensitic microstructure. Figure 39: Subcritical heat affected zone (SCHAZ) The existence of SCHAZ can be more clearly observed in the two micro-hardness maps shown in Figure 40. The hardness of base metal and weld metal is above 500 HVN. On the other hand, the hardness of SCHAZ is only 300 HVN, significantly softer than 75

96 that of the base metal and weld metal. In Figure 40(a), the SCHAZ appears as a 0.7-mmwide annulus on the longitudinal plane. The ring-shaped region is due to the axisymmetric distribution of electric current and temperature in RSW. The throughthickness profile of SCHAZ is plotted in Figure 40(b), which is slightly concave inwards following the contour of the weld nugget. Figure 40: Hardness profiles of the RSW Usibor 1500 (a) longitudinal plane (b) conventional transverse plane 76

97 To more clearly illustrate the hardness profile, a line plot of the hardness in the weldment is shown in Figure 41. A drastic decrease of hardness in the SCHAZ is observed (40% of the weld metal and unaffected base metal). The softening phenomenon shows a severe drop that gradually rises to the 500 HV hardness level for the unaffected base metal. The tempering of martensite is most severe at the point directly below the AC1 temperature. During tempering the martensite is decomposed into cementite and ferrite. Closest to the ICHAZ, the microstructure is fully decomposed, but near the unaffected base metal, the SCHAZ is partially decomposed and only undergoes slight overtempering. The gradual rise in SCHAZ hardness is directly related to the amount of untempered martensite in the microstructure. Figure 41: Hardness profile taken from the weld centerline profile 77

98 4.3.3 Gleeble Microstructural Simulation Gleeble simulations were conducted on all heat affected zone microstructures discovered in the above microstructural analysis, but in-depth studies were completed on the CGHAZ and SCHAZ because failure is routinely noted in these regions. Figure 42 shows the as-welded versus Gleeble simulated microstructures for the CGHAZ and SCHAZ. In both instances the bulk microstructure in the Gleeble samples are consistent with the microstructure in the actual weld. The simulated CGHAZ (Figure 42b) depicts large grains of freshly formed martensite. The simulated SCHAZ (Figure 42d) shows tempered martensite composed of ferrite with cementite precipitation at the prior austenite grain boundaries and martensite inter-laths, like the as-welded SCHAZ. Hardness values for the as-welded and simulated microstructures were consistent which is essential to the validity of constitutive behavior development of the heat affected zone regions. 78

99 Figure 42: (a) as-welded CGHAZ, (b) Gleeble simulated CGHAZ, (c) as-welded SCHAZ, and (d) Gleeble simulated SCHAZ UHSS Microstructure-Specific Mechanical Properties Once the microstructures of the Gleeble simulated samples were found to be consistent with the as-welded heat affected zone regions, the samples were tensile tested utilizing 2D DIC techniques to measure the strain distribution. Figure 43 is an example of an image taken prior to extreme necking for DIC and the axial strain distribution found after analysis. As shown, the strain distribution is not uniform across the gauge section. This is expected due to a temperature gradient along the axial (or longitudinal) direction 79

100 of the sample. Even with the gradient there is a 2-mm-long band, encompassed by the green crosses, in the center of the sample that gives a relatively uniform strain reading. This is the true gauge section used for measurement and calculation of the constitutive behavior of the sample. A virtual extensometer of 2 mm in length was used to determine displacement data at the center of the sample. The displacement data and force data acquired during testing were analyzed using techniques described by Huang et al. [83] to obtain the true stress versus strain curve for the SCHAZ, CGHAZ, and base metal (See Appendix A). The curves are plotted on the same graph in Figure 44 and the microstructural-specific material properties are summarized in Table 8. The SCHAZ yields at 618 MPa and has a tensile strength of 866 MPa. While still high compared to other steels, the yield strength and ultimate tensile strength are much lower than those of the Usibor 1500 base metal (yield = 1179 and tensile = 1464 MPa). The CGHAZ has a slightly higher yield and tensile strength compared to the base metal. The fracture strain of the SCHAZ is much greater than the fully martensitic base metal and CGHAZ. The SCHAZ sample fails at a strain close to 60% while the CGHAZ fails at a strain of 20%. 80

101 Figure 43: Snapshot of (a) deformation and (b) axial strain distribution measured using DIC on the simulated SCHAZ sample surface during tensile testing just prior to fracture Table 8: Microstructural-specific Material Properties Yield Strength (MPa) Ultimate Tensile Strength (MPa) Base Metal CGHAZ SCHAZ

102 Figure 44: True stress versus strain curves for Usibor 1500 base metal, CGHAZ, and SCHAZ 4.4 Conclusions The rapid heating and cooling of RSW leads to microstructural changes in the weld nugget and surrounding material. The weld nugget and directly adjacent HAZ regions are heated above the AC3. The resulting microstructure is cooled quickly from the austenite regime to form a fully martensitic microstructure. Directly adjacent to the unaffected base metal is a region of the HAZ that is heated to temperatures directly below the AC1. This temperature is not high enough to cause any phase transformation to austenite and results in over-tempering of the base metal martensite. The martensite is 82

103 decomposed into ferrite and cementite, which adversely affects the mechanical properties of this region. Failure of resistance spot welded UHSS is often seen in the weld nugget, CGHAZ, or SCHAZ. It is important to determine the microstructural and constitutive differences of these regions to get a better understanding of why different failure mode occur. From the simulated bulk microstructures, material properties could be determined for each of the regions of interest. The CGHAZ has a slightly higher yield and tensile strength than the base metal, but fails at a lower fracture strain. The SCHAZ has a significantly lower yield and tensile strength compared to the base metal, but fails at a much higher fracture strain. When discussing failure and crash simulation, it is important to determine the energy that can be absorbed. In addition to the tensile strength, the fracture strain is an important parameter. Though the SCHAZ has a lower strength, it can endure a higher strain accumulation (more ductile). The measured properties were calculated using DIC, allowing for data to be gathered beyond the uniform elongation limitations of ASTM and extensometer procedures [83]. The extended stress-strain curves take the place of empirical extrapolation methods and are utilized in the FEA model to predict the deformation behavior in the composite weld in the following chapter. 83

104 Chapter 5: Conventional Mechanical Testing Methods Mechanical testing of welded joints is an important step in determining the weldability of new materials and joint designs. Important information regarding the weld characteristics, failure mode, and strength of the joint can be gathered from various testing methods [84]. 5.1 Materials Two sheet stack-ups of Usibor 1500 were used for development and validation of different testing methods. The hot stamped boron steel sheets were 1.5 mm thick and had an aluminum silicon coating, as described in Chapter Approach Testing methods that are currently used in industry were completed to determine the joint strengths of 2T Usibor 1500 stack-ups. These testing methods were also altered to allow for more information to be gathered on the typical failure modes observed. In situ observation, post-test microscopy, and computational modeling techniques were used to determine the mechanisms of failure in RSW joints Pure Tension Testing Pure tension testing was performed on welded samples using an MTS 810 testing system. A weld was made on one regular coupon of Usibor 1500 and a smaller backing sheet of Usibor 1500, as seen in Figure 45. The sample was pulled in tension on the 84

105 100.0 mm sheet. To ensure quasi-static loading, a displacement rate of 2.54 mm/min was used for each sample. The top surface of the weldment was prepared for DIC analysis by speckling the surface with a random black and white pattern. A fine coating of white spray paint was applied to cover the entire top surface. Once the white spray paint was dry, a random black speckle was then applied to the same surface and allowed to dry before testing. A Nikon D7100 camera was used to take a video during testing. Images were extracted from the video and analyzed using Ncorr and Post_Ncorr, open-source Matlab software programs [69]. Figure 45: Sample geometry for pure tension testing [mm] 85

106 5.2.2 Tensile Shear Testing Tensile shear testing was performed on welded samples using an MTS 810 testing system. Displacement rates between mm/min were used for each sample. The mechanical properties of the weldment were assessed by measuring the peak load at failure during conventional tensile shear testing. To ensure shear loading and minimize the rotation of the sample, the geometry seen in Figure 46a was used for testing. The hidden nature of the weldment and constant rotational movement make in situ observation of local deformation difficult for tensile shear testing. Post-weld microscopy inspection of the cross-section or fracture surface can give insight on where failure occurred, but offer little information regarding strain accumulation throughout the weldment. A modified tensile shear test sample was prepared by removing the material enclosed by the red dotted boxes in Figure 46b. The sectioned surface of the sample where the welding nugget was exposed was speckled with a random black and white pattern for DIC examination. As described earlier, a fine coating of white spray paint was applied to cover the entire cross-sectioned surface. Once the white spray paint was dry, a random black speckle was then applied to the cross-sectioned surface and allowed to dry before testing. Tensile shear testing was completed using the conventional methods explained above and the test was recorded using a Nikon D7100 camera, focused on the speckled cross-section. Images were extracted from the video and analyzed using Ncorr to determine the local displacement and strain distribution. 86

107 Figure 46: Sample geometry for tensile shear testing (a) conventional and (b) modified. Dimensions in mm Metallographic Analysis The tested samples were sectioned transversely, through the center of the weld nugget, and mounted in bakelite powder with a Leco PR-32 mounting press. Mounted samples were polished using 240 to 800 grit SiC metallographic paper lubricated with water. Between each polishing step, samples were sprayed with ethanol, cleaned in an ultrasonic bath for 2-3 minutes, sprayed again with ethanol, and dried with a hot air source. The polished sample were etched with 2% Nital, for seconds depending on the desired degree of etching. 87

108 Optical microscopy was conducted using an Olympus GX51 microscope. The magnification of the optical microscope was high enough to measure crack distances and behaviors Computational Modeling Stress models were developed using Abaqus, a commercial FEA software package, to aid in the development of failure criterion for resistance spot welds of ultrahigh strength steels. Figure 47 shows the geometry, half samples with symmetry applied, and mesh near the spot weld of the model. Variations were developed using the crosssectional profiles of experimental results to incorporate disparities in the weld geometries. The results of micro-hardness mapping (Figure 40) were used to partition and assign different constitutive behaviors to the specific regions of the heat affected zone. It is noted that a spot weld contains a steep microstructure gradient. As a first step toward developing a comprehensive model, SCHAZ, i.e., the HAZ region with the lowest strength and CGHAZ, i.e., the HAZ region with the highest strength, are assigned the stress-strain curves measured from Gleeble specimen (Figure 44). The rest of the regions (FGHAZ, ICHAZ and weld metal) are assigned to the same stress-strain curves as the base metal for simplicity. The elastic properties (Young s modulus and Poisson s ratio) are the same for the entire spot weld as they are not strongly dependent on the microstructure. For pure tension testing, one side of the sample was constrained while the other was pulled along one plane. For tension shear, the sample was also constrained to only allow motion in one plane, simulating the shims used in experimental testing. For 88

109 simplicity, residual stresses form the resistance spot welding process were not considered. Abaqus Standard, an implicit solver, was used to understand the effect of the inhomogeneous microstructure on deformation behaviors of the weldment. Models using uniform base metal properties for the entire weldment were also ran for comparison purposes. Figure 47: Half models with applied symmetry and mesh for mechanical testing simulations 5.3 Results and Discussion Conventional testing methods can be used to develop welding parameters that result in acceptable failure modes. Weldments along the entire current range developed in 89

110 Figure 35 show strengths greater than that required by the AWS standards. Though strength requirements are met, a variety of failure modes are seen Tension Testing The tension test is useful for determining the local deformation of a spot weld when subjected to a simple load in one direction. By using a simple tension test, there are no rotational effects to align the spot-welded specimen. Figure 48 shows the crack initiation and propagation in a tension specimen during testing. The red dashed circle represents the indentation and a rough estimate of the nugget that is hidden in between the pulled sheet and backing sheet. The crack initiation is outside of the nugget, in the softened region of the heat affected zone. The SCHAZ, or softened region, is the weakest part of the resistance spot weld. Once the crack initiation has taken place, final failure is seen at a 45 angle through the base material. The strain localization can be observed along the same path as fracture. Small strain build-up is seen encompassing the weld nugget, as if it was a rigid button. Maximum strain is located outside the weld. The top of the sample is held stationary and the bottom of the sample is pulled continuously. One probable cause of local accumulation at the bottom of the nugget is that the properties are not completely symmetric about the weld center due to factors such as slight misalignment of the test specimen. The localization of normal strain in the Y direction is shown on the right side of Figure 48. From hardness profiles, the weld nugget and SCHAZ/base metal radii are 3.11 and 3.81 mm, respectively. From the center of the nugget, is was confirmed that failure occurs in the SCHAZ of the weldment. 90

111 The maximum strain locally seen in the SCHAZ before the crack initiation is but quickly jumps to On the other hand, the global strain in the bulk of the sample which is primarily base metal is around The plot of load and strain versus time is seen in Figure 49. Crack initiation occurs 108 seconds into the test, which corresponds with the quick drop in load and rise in strain. The strain is measured by a virtual extensometer at the SCHAZ where the onset of failure occurs. There was no gauge section to force failure on a specific region. Therefore, the force before failure is much higher than a typical spot weld load because of the excess base metal on each side. 91

112 Figure 48: Succession images of tension failure in one sheet (a) initial prior to loading (b) crack initiation (c) crack propagation and the corresponding DIC Eyy strain ( y) accumulation maps 92

113 The equation below can be used to convert the force to stress. The cross-sectional area of the entire plate (1.5 mm by 40 mm) is used because the load is carried by the entire plate. Stress [MPa] = Force [N] Area [mm 2 ] (4) At the peak load, prior to failure, the stress imposed on the plate is 1348 MPa. This stress is less than the ultimate tensile strength of the base material, depicting that the softened SCHAZ, which has a lower ultimate tensile stress, is lowering the ultimate strength of the joint. These results are consisted with similar testing done by Burget and Sommer [85]. Figure 49: Load and Strain versus time for a pure tension test 93

114 In Abaqus FEA, simulating the experimental results with a homogeneous weld (uniform property of base metal) (Figure 50a), shows a strain accumulation at the weldment, but does not predict the local degradation in the SCHAZ. By simply incorporating the local constitutive behaviors (Figure 50b), a much better depiction of strain accumulation can be seen in the Abaqus simulation. While this testing method is repeatable and useful for verifying model accuracy, it is not a common testing method in the automotive industry because the failure mode produced is not commonly seen in structural applications. Figure 50: Abaqus simulation of tension test with (a) analysis with homogeneous base metal properties (b) strain accumulation at crack initiation with inhomogeneous profile 94

115 5.3.2 Tensile Shear Testing Effect of Weld Schedule The size of the resistance spot weld can alter the stress-state of the joint and therefore change the mode in which it fails. In Figure 51, four different failure modes corresponding to different currents along the 2T Usibor 1500 range (Figure 35) are shown. The nugget diameter, current, and peak load for the corresponding weld failures are listed in Table 9. Figure 51: Failed tensile shear test specimens (a) interfacial (b) partial button pull out (c) SCHAZ button pull out and (d) expulsion button pull out 95

116 Table 9: Common failure modes of tensile shear testing and relevant data Failure Mode Nugget Diameter [mm] Welding Current [ka] Peak Load [N] Interfacial Partial BPO BPO Expulsion BPO Interfacial failure is typically seen for small welds but can also correlate to welding schedules that exhibit expulsion. Interfacial failure (Figure 51a), is an abrupt failure that occurs along the interface with a straight, flat fracture path. Partial button pull-out (Figure 51b) occurs when the notch-like feature causes high stress concentration, but the ultimate failure is shared between the interface and a region outside of the nugget, typically the SCHAZ. Full button pull-out is used to describe failure modes that occur outside of the weld metal. The cross-sectional images of Figure 51c and Figure 51d show two different button pull-out failure modes. In SCHAZ button pull-out ultimate failure is away from the weld nugget, but when expulsion is present, the failure path follows circumferentially along the weld nugget and CGHAZ boundary. From the load versus stroke curves for tensile shear testing, in Figure 52, partial button pull-out and SCHAZ button pull-out have comparable maximum loads at failure. Both see the same ultimate failure predominantly occurring in the softened SCHAZ. The expulsion condition has a button pull-out failure mode, but the maximum load endured by the joint is much lower than the loads seen by the previously mentioned failure modes. When expulsion occurs, the ability to predict the failure mode is difficult because it depends so highly on the formation of the expulsion and amount of material expelled. 96

117 For conventional tensile shear testing, the spot weld of nominal size and current (6.5 ka) fractured through the softened SCHAZ, the same button pull-out failure observed by Burget and Sommer [9]. The cross section of such fracture can be seen in Figure 51c. The button pull-out failure is a result of the rotation of the weldment after the initial pull. The rotation aligns the nugget with the loading line, which acts as a rigid button. Thus, the softened SCHAZ surrounding the button experiences the bulk of the load, therefore being the site of deformation and subsequent fracture. In the crosssectional image, small cracks formed at each side of the weld interface, at the notch-like feature. Geometrical stress concentration at the notch-like feature gives a competing failure path to the ultimate failure of the spot weld. However, for welds of nominal size, the fracture through the SCHAZ leads to the ultimate failure before the crack at the weld interface has a chance to propagate further to allow for weld nugget failure. Figure 52: Load versus stroke curves for different failure modes 97

118 Effect of Inhomogeneous Microstructure To further understand the effect of considering the SCHAZ mechanical properties on the deformation behavior of a UHSS resistance spot weld, two simulations were made using the FEA model for tensile shear testing. The first simulation used uniform mechanical properties of only base metal and the second simulation called out the unique constitutive behaviors of the differing regions of the weld profile. The maximum plastic strain was monitored at the two locations where competing failure mechanisms are seen in Figure 51c: the SCHAZ and the notch-like feature at the weld interface. In Figure 53, the strain accumulation at different force levels is seen for the uniform property and inhomogeneous profile, respectively. For the uniform profile (Figure 53a), the calculated maximum plastic strain is much higher at the weld interface than at the SCHAZ location. If the SCHAZ and other unique mechanical properties are considered (Figure 53b), the competing failure behaviors show similar build-up of strain at increasing forces initially. To approximate the probability of crack initiation, the plastic strain is compared to the true fracture strain seen in base metal (grey dashed line) and SCHAZ (black dashed line) microstructural samples. These values are calculated by the area reduction seen in the bulk microstructural tensile specimens and displayed in Table 10. The true fracture strains are calculated based on Equation 5 and 6 for the bulk microstructural tensile specimens and are meant to serve as simple failure criterion. True fracture strain is analogous to the total strain at fracture on the engineering stress-strain curve. These values serve as an initial set of failure criteria and many assumptions must be made for use in simulations [61]. It is assumed that the level of stress triaxiality and 98

119 strain rate are comparable for the welded and bulk microstructural specimens. These values can also only be used for ductile failure, which was the bulk tension failure mode of the bulk samples, and cannot be used as criterion for brittle interfacial failure through the centerline. Dancette et al. use a similar method to avoid using more advanced damage models which cannot be easily implemented for the inhomogeneous weld profile [61]. ε f = ln q 1 q (5) q = A 0 A f A 0 (6) In the simulated homogeneous weld profile, the strain accumulation at the notchlike feature severely exceeds the elongation limit of the base metal but does not reach the limit in the SCHAZ, indicating failure would not initiate at the SCHAZ location. In other words, it indicates that using a homogeneous profile, interfacial failure originating at the notch-like feature would be the preferred failure path for the RSW joints subjected to tensile shear loading. Under the same loading conditions but with a more representative heterogeneous profile, the plastic strain at the SCHAZ and notch-like feature exceed the elongation limits for the corresponding constitutive behaviors simultaneously. This indicates that fracture initiation is likely to occur at both locations, much like the crack formation seen in Figure 51c. By considering the local constitutive behaviors of the cross-sectional profile, simulated resistance spot welded joints can more accurately predict the deformation and failure patterns seen under various loading conditions. 99

120 Table 10: True fracture strains calculated by the area reduction in pure tensile samples Microstructural Region True Fracture Strain Subcritical HAZ Coarse-Grained HAZ Base Metal Figure 53: Strain accumulation at the notch-like feature and SCHAZ locations for (a) homogeneous weld profile and (b) unique heterogeneous weld profile An initial attempt at modeling damage and failure was made using the XFEM mode in Abaqus Standard. With the inhomogeneous microstructure/property profile, failure initiation is competing between the notch location and SCHAZ. The true fracture strains (Table 10) were used as the failure criterion. Initially, in the top image in Figure 54, a crack is seen propagating at the interface of the joint. Experimentally, this is also seen in Figure 51c. Upon further loading, a new fracture initiation is seen in the SCHAZ, where ultimate failure was seen experimentally. 100

121 Figure 54: XFEM crack initiation during tensile shear testing of a weld that experimentally failed by button pull-out (top) initial fracture at interface (bottom) further pulling shows failure initiation at the SCHAZ Effect of Notch Geometry The inhomogeneous microstructure profile of a weldment is consistent from weld to weld when the same parameters are used, but a level of uncertainty exists with the geometric profile of the weld nugget at the interface. When welding Usibor 1500, there are four different notch profiles that can be developed and are shown in Figure 55. Each of these profiles are present in welds within the acceptable current range and below the expulsion limit. It is unclear exactly what causes each unique notch. Possible interactions of the electrode force, molten nugget expansion forces, and coating could combine to create a complex phenomenon. 101

122 Figure 55: Common notch profiles for 2T Usibor 1500 weldments (a) CGHAZ blunt (b) WM blunt (c) WM bi-notch and (d) WM sharp These different notch features create a discrepancy on the force present at the weld interface during testing. Abaqus simulations were completed with the various notch formations to determine how the failure mode could be affected. In Figure 56, a visual of the strain accumulation during tensile shear testing can be seen. In the case of a weld metal blunt notch or weld metal bi-notch, the strain accumulation is primarly at the interface in the weld nugget. If a layer of CGHAZ can protect the weldment, the strain accumulation is primarily in the softened SCHAZ. A sharp notch leading to the weld nugget shows competing strain accumulation between the CGHAZ adjacent to the nugget and the SCHAZ. 102

123 As a first approximation of the failure iniation, Figure 57 shows the numerical values of strain accumulation at the notch, CGHAZ, and SCHAZ for the different notch geometries. The red crosses symbolize the location that reaches the ultimate fracture strain (see Table 10) first. It can be seen that depending on the local notch geometry, failure iniation is seen in the weld metal, CGHAZ, and SCHAZ. This adds a level of uncertainty in the studying the failure modes and more research needs to be completed to determine the causes of different notch geometries in the weldments. Figure 56: Abaqus simulations displaying PEEQ corresponding to the different notch geometries 103

124 Figure 57: Stroke versus maximum principal strain for locations near the different notch features Modified Test Specimen Tensile shear testing is the most widely used method for determining the weldability and failure behavior of resistance spot welds. Though widely used, it still contains some flaws that hinder the amount of information able to be gathered. To better investigate the deformation behavior of RSW joints subjected to a tensile shear load, half weld testing was conducted where a transverse section of the weld was observed using the 2-D DIC. The same rotational effects still occur in the sectioned spot-welded sample. However, because symmetrical cuts were taken from both sides of the nugget, the 104

125 observed section remained planar during deformation, facilitating the strain mapping using 2-D DIC. It is noted that the surface in which the DIC data is analyzed is not the center plane of the weld, therefore the nugget sizes and heat affected zone regions do not match the true cross-sectional profile measurements described previously. In Figure 58, the sectioned side of the tensile shear specimen can be seen at the onset of deformation and immediately before ultimate failure of a weld made with nominal welding conditions. The weld fails through the softened SCHAZ and accumulates a maximum engineering strain of prior to failure. This value is lower than the strain accumulation prior to failure in the tension test. Surface and geometrical effects can have a play in the strain values recorded, but strain accumulation is seen away from the notch-like feature and in a region associated with the soften properties of the SCHAZ. Although not analyzed further in the present research, one future use of the reduced-section tensile shear test is that the quantitative strain maps can be used to validate the finite element model. Such data provides the deformation and failure behavior when the spot weld is subjected to mode-ii loading and is complementary to those generated in the single-sided wedge testing (mode-i) discussed in the next section. 105

126 Figure 58: Images and strain accumulation of sectioned weld samples subjected to tensile shear loading at (a) onset of deformation and (b) immediately before ultimate failure 106

127 5.4 Conclusions Mechanical testing is an integral part of validating resistance spot welds for structural applications. Methods currently being used give some information, but can be better analyzed and altered to give more information on the weldment properties. With higher strength materials, it is important to see how the inhomogeneous microstructural profile alters the failure and deformation associated with a spot weld. The geometry of a RSW poses problems in simplistically testing a weldment in a single loading direction. Typically, rotational effects create a multi-facet stress state on the weldment. Tension testing, which applies a load to a single sheet that is spot welded to a backing plate, creates a simple tension loading condition on the weldment. In this case, the base material surrounding the weldment plays a large role in failure and ultimate strength, but the strain accumulation can be observed to better understand where deformation is occurring. Tensile shear testing is used predominantly in the automotive industry to determine the peak loads and failure modes of resistance spot welds. While this test is useful, it allows for no in situ observation of the deformation behaviors leading up to failure. The weld is hidden at the interface and until ultimate failure, the weldment cannot be viewed. The heterogeneous profile of 2T Usibor 1500 weldments shows a more complex failure behavior that includes competing factors between the notch-like feature and softened SCHAZ. Different notch geometries are seen during welding and the variation adds another complex factor in determining failure mode of weldments. Computational models incorporating the unique properties of the different regions of the 107

128 spot weld can give more insight on where peak stresses and strain accumulation is being developed during testing. Modeling different notch features shows the variation of strain accumulation is very much dependent on the geometrical aspects of the weldment, which can only be discovered by destructive testing. By sectioning the weld during experimental testing, strain mapping can be used to determine where failure is initiating and propagating. Considering the different mechanical properties of the varying heat affected zone is essential in modeling the correct strain distribution seen during experimental testing. With unique constitutive behaviors, the localized hot spot of plastic deformation can be observed in the simulated tests to match the experimental observed results. When resistance welding ultra-high strength steel, it is essential to consider the microstructurespecific mechanical properties to accurately predict deformation and failure behaviors. 108

129 Chapter 6: Single-Sided Wedge Testing Most testing procedures for resistance spot welds are described in standards, but standardized testing procedures are limited in the quantitative feedback that can be achieved. Due to this, some developmental testing procedures are described in this section to give more detailed feedback on weldment behavior during loading. 6.1 Materials Two sheet stack-ups of Usibor 1500 were used for development and validation of different testing methods. The hot stamped boron steel sheets were 1.5 mm thick and had an aluminum silicon coating. 6.2 Approach Wedge testing is completed on cross-sectioned samples to allow for in situ observation of the deformation behaviors. The testing methods were inspired by Payen et al. and their use of similar testing methods in an SEM chamber [66]. From their results, an innovative testing method was developed to observe deformation behaviors of resistance spot welds. The details of the single-sided wedge testing procedure are explained below Experimental Development Welded samples are made in the center of 30 mm wide coupons, with beveled interfacial edges, in the 2T configuration (Figure 59). Welds are cross-sectioned to allow 109

130 the center-point of the nugget to be seen with some base material, but not excessive, at the rear of the weld. The cross-section of the weld is prepared for DIC analysis by speckling the surface exposing the weld nugget with a random black and white pattern. A fine coating of white spray paint is applied to cover the entire cross-sectioned surface. Once the white spray paint is dry, a random black speckle is applied to the crosssectioned surface and allowed to dry before testing. The cross-sectioned surface must not move or deform excessively out of the plane during testing to allow for accurate DIC analysis in 2-D. A Nikon D7100 camera was used to take a video during testing. Images were extracted from the video and analyzed using Ncorr and Post_Ncorr, open-source Matlab software programs [69]. To perform the single-sided wedge test, wedges were constructed of O2 tool steel in the geometry seen in Figure 60. The angle of the wedge was developed by using experimental trial and error to determine a consistent force and loading mode. A single wedge was plunged toward the weld nugget through the interface systematically with an Instron 5985 system while the other side of the joint was clamped to mitigate rotation, as seen in Figure 61. Figure 59: Sample geometry for wedge testing [mm] 110

131 Figure 60: Two different geometries used for wedge testing [mm] Figure 61: Wedges are plunged at the interface of the weld during testing The wedge and clamping device were placed in the grips of the testing machine. The wedge was inserted between the two sheets of the cross-sectioned samples and then gradually pushed inward (i.e., toward the center of the weld) using a compressive displacement rate of 3.00 mm/min. A Nikon D7100 camera was used to record changes in the speckle pattern as the wedges were pushed in, deforming the sample. Images were extracted from the recorded test and analyzed using Ncorr, an open source 2D DIC processing software based on Matlab [69]. The camera set-up for the wedge testing can be seen in Figure

132 Figure 62: A Nikon D7100 camera is used to record the deformation behaviors during wedge testing Metallographic Analysis The tested samples were sectioned transversely, through the center of the weld nugget, and mounted in bakelite powder with a Leco PR-32 mounting press. Mounted samples were polished using 240 to 800 grit SiC metallographic paper lubricated with water. Between each polishing step, samples were sprayed with ethanol, cleaned in an ultrasonic bath for 2-3 minutes, sprayed again with ethanol, and dried with a hot air source. The polished sample were etched with 2% Nital, for seconds depending on 112

133 the desired degree of etching. Optical microscopy was conducted using an Olympus GX51 microscope. The magnification of the optical microscope was great enough to measure crack distances and behaviors Computational Modeling Similar to those used in Chapter 5, stress models were developed using Abaqus, a commercial FEA software package, to aid in the development of the single-sided wedge test. Figure 63 shows the geometry and mesh near the spot weld of the model. Variations were developed using the cross-sectional profiles of experimental results to incorporate disparities in the weld geometries. The results of micro-hardness mapping (Figure 40) were used to partition and assign different constitutive behaviors to the specific regions of the heat affected zone. As discussed in Chapter 5, two annuluses were partitioned from the spot weld and assigned to CGHAZ and SCHAZ stress-strain properties, respectively. The rest of the spot weld is assigned to the base metal properties for simplicity. For single-sided wedge testing, the constrained side of the weld was fixed to minimize any rotational effects. The wedge was modeled as a rigid part and mechanical contact interactions were defined between each sheet and the wedge surface. For simplicity, residual stresses from the resistance spot welding process were not considered. Abaqus Standard, an implicit solver, was used to understand the effect of the inhomogeneous microstructure on deformation behaviors of the weldment. Models using 113

134 uniform base metal properties for the entire weldment were also ran for comparison purposes. Figure 63: Single-sided wedge test geometry and mesh for simulation 6.3 Results and Discussion Previously a current range was developed for the 2T Usibor 1500 stack-up (Figure 35). For the mechanical testing evaluation, welds were made to evaluate the behavior of welds of assorted sizes. Resistance spot welding was conducted to create samples at the two extremes, below the minimum nugget diameter and at expulsion, as well as multiple points within the acceptable range, respectively small, medium, and large. These points are placed in the current range in Figure

135 Figure 64: 2T Usibor 1500 current range with markers indicating tested weld sizes Effect of Weld Nugget Size The main goal of the new testing procedure was to be able to track where deformation occurs and how the weld size can attribute to localized deformation. As the weld size increases, the failure mode changes from interfacial, through the weld metal, to failure in the SCHAZ, adjacent to the unaffected base metal as seen in Figure 65. The different weld sizes are plotted versus the insertion depth, or how far the wedge was inserted during testing. Trials containing welds that did not fail interfacially were stopped prior to ultimate failure. 115

136 Figure 65: Wedge testing results showing weld size versus insertion depth for 3 pulse welding schedule At weld sizes below 5.8 mm the failure is predominantly abrupt interfacial failure along the centerline of the stack-up. In between 5.8 mm and the expulsion point, failure mode deviates from weld metal, CGHAZ, and SCHAZ failure. Figure 66 plots deformed welds of varied sizes in FEA models that account for the unique properties of the weld profile. The wedge insertion for the modeled joints is ideal and the force is shared perfectly between the two legs of the weldment. As the weld size increases, the maximum plastic deformation shifts from the notch-like feature to the SCHAZ. A bending moment is created in the SCHAZ allowing the sheets welded to distort away from the notch-like feature where a high geometrical stress intensity is already present. 116

137 Figure 66: Simulated single-sided wedge testing results showing maximum plastic strain accumulation for three different weld sizes Interfacial failure occurs abruptly and is not an ideal failure mode for the weldment (Figure 67a). The fracture path is very straight along the interface which allows for no warning before ultimate failure. CGHAZ failure occurs directly adjacent to the weld metal (Figure 67b). The fracture path starts from the interface of the two materials, but travels along the outside periphery of the weld nugget, in the CGHAZ. Weld metal failure begins from the same notch-like feature as interfacial failure, but does not follow a straight path at the centerline in the weld nugget (Figure 67c). Instead, the path is diverted upward into one of the work pieces. SCHAZ failure occurs adjacent to the 117

138 unaffected base metal (Figure 67d). The strain locally accumulates in the softened HAZ, which is more ductile than the other fully martensitic regions. Figure 67: Failure modes of single-sided wedge testing (a) interfacial (b) CGHAZ (c) weld metal (d) SCHAZ The load versus displacement curves can be seen in Figure 68. The data for weld metal, CGHAZ, and SCHAZ failure is not extended to failure because testing was stopped prior to ultimate failure. Interfacial failure records the lowest peak force and insertion value of the failure modes. CGHAZ failure undergoes the highest peak load. The load quickly reaches a peak value of roughly 2475 N with very little deformation. SCHAZ failure sees more deformation before the peak load, 1577 N, is reached. These failure modes see deformation that is not in line with the loading path. Since the initial propagation does not coincide with the notch-like stress concentration, higher loads can be applied. Weld metal and interfacial failure begin at the notch-like stress concentration at the interface, inherently needing a smaller load for fracture initiation to occur. 118

139 Figure 68: Load versus displacement curves for the different failure modes of singlesided wedge testing (NOTE: tests were stopped prior to ultimate failure, except in case of interfacial failure) The DIC axial strain mapping of interfacial, weld metal, and SCHAZ failure can be seen in Figure 69. The scale of strain goes from to 0.14 to show compressive and tensile strains accrued in the vertical direction during testing. In interfacial failure (Figure 69a) very little strain is accrued in the sample. Slight build-up can be seen at the notchlike feature prior to the end of testing, but no deformation is seen before ultimate failure which is the end of testing in this case. Weld metal fracture (Figure 69b) shows strain accumulation in the weld metal at the peak load and end of testing, which is not ultimate failure. SCHAZ fracture (Figure 69c) has indications of tensile and compressive strain on the inside and outside of the joint respectively. More strain is built up in the left leg of the weldment due to misalignment of the test. The region of maximum axial strain correlates 119

140 to the SCHAZ measured in hardness mapping. The strain accumulation and plastic deformation seen in the sample is due to the bending being localized in the softened SCHAZ. Figure 69: Normal strain map in the Y direction measured by DIC for (a) interfacial (b) weld metal and (c) SCHAZ fracture paths 120

141 6.3.2 Effect of Weld Leg Size Depending on how the samples were cut, there was a noticeable variation in the sample length, especially the distance from the cut edge to the periphery of weld nugget. It was later recognized that such distance, referenced to as the weld leg size, actually played a large role in the deformation and failure behavior of the joint. The leg varies the degree of bending or rotation that the weldment can experience. In Figure 70, the leg size is plotted versus the insertion depth. For interfacial failure, the leg size had very little effect as failure was seen with from 8 mm and 12 mm legs. For the more complex fracture paths, there seems to be a correlation between leg size and failure mode. At smaller leg sizes, failure occurs in the SCHAZ which is located furthest away from the weld nugget. As the leg size increases, the failure mode shifts toward the weld nugget and weld metal and CGHAZ failure occurs. With smaller leg sizes, the SCHAZ can act as a hinge allowing more deformation and fracture initiation. At larger leg sizes, the strong base material mitigates the hinge effect of the SCHAZ and causes more force to be applied at the notch-like feature at the interface. 121

142 Figure 70: Single-sided wedge testing results showing leg size versus insertion depth for 3 pulse welding schedule Effect of Welding Schedule The above results are all using a three-pulse welding schedule. In addition, a onepulse, 20 cycle weld schedule was used for comparison (Figure 71). Eftekharimilani discusses how adding multiple pulses to a resistance spot welding schedule of ultra-high strength steels can alter the edge of the weld, changing its properties and typical failure mode. A second pulse (or third) of equal current to the initial pulse fully anneals the primary weld nugget edge, thus homogenizing the alloying elements and increasing the strength of the weldment [86]. The last pulse weld has a very distinct dendritic structure where the grain orientation is in the direction of the cooling electrodes. Conversely, the periphery or the previous weld nugget is annealed and shows equiaxed grains of martensite along the weld nugget edge. 122

143 The minimum weld size to ensure that interfacial failure does not occur is not as defined, as seen in Figure 72. Interfacial failure and WM/CGHAZ failure coincide around 6.2 mm weld size. Without a clearly defined minimum nugget diameter to remove interfacial failure, it becomes difficult to ensure the failure mode is suitable for enduring loads and deformation in which it may be subjected. Figure 71: Schematic comparison of the welding schedules for the three-pulse (solid line) versus one pulse (dashed line) welds Figure 72: Single-sided wedge testing results showing weld size for 1- and 3-pulse welds versus insertion depth 123

144 6.3.4 Effect of Testing Configuration The single-sided wedge test is still being developed for optimized results and repeatability. One of the major issues that arises is misalignment of the wedge and sample. Simulated results show ideal conditions where the load is evenly distributed between the two sheets. Experimentally, the alignment is not perfect and one of the two sheets typically endures the bulk of the load. Using the FEA model developed, the effect of varying degrees of wedge rotation was analyzed and displayed in Figure 73. At 10 rotation, there is a preference to the top sheet as opposed to the bottom. Increasing the amount of rotation, there is severe deformation in one sheet and at 35 rotation there is no strain accumulated and thus no load being applied to the bottom sheet. Figure 73: Simulated maximum plastic strain accumulation for different degrees of misalignment in single-sided wedge testing 124

145 6.4 Conclusions Single-sided wedge testing is a novel mechanical testing method that can give much more insight on the deformation behaviors of the joint compared to conventional peel and cross-sectional testing. As the weld size increases, the failure mode shifts from abrupt interfacial to weld metal, CGHAZ, and SCHAZ failure. The shift of failure mode is at a 5.8 mm nugget diameter for a three-pulsed welding schedule, which is nearly 1 mm larger than the recommended weld size per 4 t standards. Without a pulsed welding schedule, and an enhanced grain structure at the weld nugget edge the nugget diameter becomes much be less clear. Interfacial failure is seen intermittently with other failure modes. Once the weld size is large enough to mitigate interfacial failure, the leg size influences the failure mode. At larger leg lengths, the base metal strength hinders the hinging action of the softened and ductile SCHAZ, thus forcing failure to initiate near the notch-like feature. At smaller leg lengths, the bending moment around the SCHAZ is greater and deformation can accumulate in the softened region, where crack initiation is seen. The single-sided wedge test needs to be further developed to become a standardized testing method. The effect of misalignment needs to be further characterized to determine its effect on failure mode and the robustness of the procedure. 125

146 Chapter 7: Complex Stack-Up Analysis The work discussed in the earlier chapters is composed of 2T stack-up joints of the same steel. While many resistance spot welds can be related to this geometry, there are many situations where more complex configurations are needed. With light-weighting initiatives, heterogeneous stack-ups are being studied more heavily. In this chapter, dissimilar materials are combined in 2T and 3T combinations to determine the robustness of these materials to unique stack-up scenarios. 7.1 Materials The joints of interest consist of dissimilar sheets, as seen in Figure 74, which make selecting standard welding procedures more difficult. The 1.5 mm-thick Usibor 1500, coated in Al-Si, has been used for the work conducted in the previous chapters. For the other steels, two sheets are each 1.4 mm-thick uncoated advanced high strength steels, JSC 980 YL and JSC 590 R respectively. The last steel of interest is a 0.75 mm-thick mild steel with a galvanneal coating. The thickness of this sheet is roughly half, or less than half of the thickness of the other three materials. 126

147 Figure 74: Complex stack-up configurations for 2T and 3T welds 7.2 Approach Complex stack-ups is a term used for any resistance joint that contains two or more different materials. Material variation could be in terms of composition, thickness, or coating and can imply steels, non-ferrous alloys, and adhesives. Some automotive OEMs and suppliers also use the term regarding single-sided joining methods and difficult to reach locations on an automotive body [12]. A common joint combination that has received much attention due to its plethora of applications is the thin/thick/thick combination. Gould et al. discuss its application in an automotive body where two thick frames are joined to a floor panel or skin side panel Resistance Spot Welding Preliminary RSW tests were conducted on a Taylor-Winfield single-phase 60 Hz alternating-current (AC) resistance spot welding machine with a Medar control unit. The electrode tips used were dome-shaped RWMA class II with 6 mm diameter face. The 127

148 welding variables altered consisted of weld current, weld pulse length, number of weld pulses, and electrode force. A 2T configuration composed of a thin/thick stack-up and two 3T configurations composed of both a thin/thick/thick and thick/thick/thick stack-up were investigated. These different stack-ups gave insight on the effect of current and weld time at the different interfaces. Literature standards and research investigating 3T stack-ups were used to develop starting parameters to weld the stack-ups presented in Figure Metallurgical Analysis The resistance spot welded samples were sectioned transversely, through the center of the weld nugget, and mounted in bakelite powder with a Leco PR-32 mounting press. Mounted samples were polished using 240 to 800 grit SiC metallographic paper lubricated with water. Between each polishing step, samples were sprayed with ethanol, cleaned in an ultrasonic bath for 2-3 minutes, sprayed again with ethanol, and dried with a hot air source. The polished sample were etched with 2% Nital, for seconds depending on the desired degree of etching. Optical microscopy was conducted using an Olympus GX51 microscope. This type of analysis could be used for visual determination of nugget size and broad comparison between samples. The minimum acceptable nugget sizes for the thickest material in the stack-up can be seen in Table 6, where the 1.5 mm thickness is highlighted. While the AWS minimum nugget size is 5.0 mm, the thinnest material in the stack-up is only half as thick. Further studies will be needed for standardizing acceptable nugget sizes for the complex stack-up. 128

149 7.3 Results and Discussion The time and current were used to alter the welding parameters of the complex stack-ups. In a production setting the electrode geometry and weld force are constant values of a weld gun. The current and time can be exclusive for a specific weld configuration. It is important to see how the parameters effect of weld nugget growth at the geometrical center point of the joint and each individual interface Thin/Thick 2T Stack-Up The thin/thick stack-up was primarily easy to weld using a single pulse welding schedule. The weld nugget begins to grow at the geometrical center point, not at the interface between the two sheets. With low cycles times, the weld nugget does not have sufficient time to grow and create a nugget diameter sufficient to meet the minimum requirement. In Figure 75, the 4-cycle weld nugget is 3.34 mm at the geometrical center and 2.99 mm at the interface. The 6-cycle weld nugget is 4.60 mm at the geometrical center and 4.45 at the interface. At 8-cycle weld time the nugget met the minimum size requirement with a diameter of 4.95 mm at the geometrical center and 4.59 mm at the interface. By doubling the weld time, from 4 to 8 cycles, the nugget can grow outward and form a more oblong nugget that can encompass the interface with a larger horizontal nugget presence. At shorter times, the nugget is unable to grow outward and thus the interface has a much smaller diameter than the geometrical center. 129

150 Figure 75: Effect of time on thin/thick stack-up (a) 4 cycles, 7.8 ka (b) 6 cycles, 8.0 ka and (b) 8 cycles, 8.0 ka It should be noted in Figure 75(a) there is an excess indentation on the top sheet. This can be due to misalignment of the stack-up to the electrode during fabrication. There is also warping of the top sheet in Figure 75(c), which could have been caused by preexisting welds placed on the same coupon of material (as some of the coupons were reused for testing). In other words, these features are likely due to process inconsistencies and the inherent scatter of the resistance welding process. The presence of a thin and soft steel sheet coated with Zn coating may amplify the process variables and further work is necessary to develop a robust process window for production. 130

151 In the thin/thick 2T stack-up penetration of the weld nugget into the thin sheet is not an apparent problem. There is sufficient nugget growth into the think sheet to meet and exceed the 50% threshold to signify an acceptable weld [12]. The equation used for calculating the penetration is below. The height of the nugget from the thin/thick interface is divided by the thickness of the thin sheet, which is 0.75 mm nominally but needs to account for indentation during welding. Per AWS Standards, the indentation value should be no more than 25% of the sheet thickness [17]. % Penetration = h nugget t thin sheet 100% (7) In the 4-cycle weld (Figure 75a), the weld nugget almost reaches the bottom of the indentation crevice at the thin sheet. The indentation value is approximately 52%, which is considered excessive by AWS specifications. The nugget penetration into the thin sheet is 87% when the indentation is accounted for and 40% compared to the nominal thickness. In the 8-cycle weld (Figure 75b) the indentation is approximately 10%, which is sufficient for an acceptable weld. On the left side of the nugget, there is excessive deformation though which could be due to misalignment of the sample during welding. The nugget penetration into the thin sheet is 60% for the nominal thickness of 0.75 mm, which exceeds the penetration requirement stated by Gould et al. By increasing the current, the weldment grew even until the expulsion limit. In Figure 76 there are two welds made at 8.0 ka and 8.5 ka with an 8-cycle weld time. As the current increases, the weld size grows due to increased heat being generated. With a 500 A rise in current, the nugget grew in diameter by 0.23 mm. While the nugget 131

152 diameter increases, the height of the weld nugget does not increase. The penetration levels remain the same, showing that a vertical growth threshold was reached in the nugget. The expulsion point was reached at 9.0 ka, and the nugget size also fell due to the molten material leaving the nugget. Figure 76: Effect of current on thin/thick stack-up (a) 8 cycles, 8.0 ka and (b) 8 cycles, 8.5 ka Thin/Thick/Thick 3T Stack-Up The thin/thick/thick 3T stack-up is like the previous 2T stack-up but adds an additional thick sheet to the bottom of the joint. This increases the offset of the thin/thick interface from the geometrical center point of the weldment. The effect of current is similar in the 3T configuration as seen in the 2T previously. The nugget diameters in Figure 77 increase with respect to the current increases. 132

153 Figure 77: Effect of current on thin/thick/thick stack-up (a) 8 cycles, 7.0 ka (b) 8 cycles, 7.7 ka and (c) 8 cycles, 8.6 ka Since the thin sheet is further away from the geometrical center point of the weld, the penetration is much less into the thin sheet. Until 8.6 ka is applied to the 3T stack-up, there is less than 10% nugget penetration into the thin sheet. At 8.6 ka, as seen in Figure 133

154 78, there is about 15% indentation caused by the electrode. Considering the indentation, the nugget penetration is 55%, but when using the nominal thickness of the thin sheet the nugget penetration is 45%. At higher welding current, the expulsion limit is reached. Figure 78: Close-up of the penetration measured in the thin sheet of the stack-up With the use of a more complex welding schedule, the penetration could potentially be increased in the joint, without increasing the current beyond the point of expulsion. The addition of another thick sheet to the bottom of the stack-up alters the vertical growth needed to create ample penetration into the thin sheet. The nugget growth in the thin sheet can also be related back to the Joule heating equation discussed in Equation Q = I 2 Rt (1. The heat generated is a function of the resistance, which can be altered by the material and contact. The thin material has a galvanneal coating, which is much more conductive, thus reducing the resistance level at the electrode and top interface of the stack-up. The thin sheet is also weaker than the bottom thicker sheets. The weaker, softer material is able to deform and reduce the asperity levels, resulting in an increase in effective contact area and thus a decrease in the current density entering the top sheet and that flowing from the 134

155 top sheet to the middle sheet. In other words, this means that the current at both the top and bottom surfaces of the top sheet is spread over a wider area, decreasing the current density and thus the Joule heat generation at a given point. A schematic depiction of this phenomenon is shown in Figure 79. Moreover, as the heat is quickly conducted into the water-cooled top electrode, a less intense heat input would have a challenging time in heating the top sheet. Figure 79: Schematic view of the current density mismatch between the softer, more conductive material and the harder, more resistive bottom sheet seen in the 3T stack-up Thick/Thick/Thick 3T Stack-Up The last stack-up is made of three different sheet materials with roughly the same thicknesses. With similar thicknesses of all three sheets, it is not difficult to grow the nugget to encompass both the top and bottom interface. The weld time was varied from 10 to 30 cycles to determine the amount of time needed to generate a nugget that exceeds the minimum requirement. 135

156 Figure 80: Effect of time on thick/thick/thick stack-up (a)15 cycles, 6.4 ka (b) 20 cycles, 6.8 ka and (c) 30 cycles, 7.2 ka In Figure 80, as the weld time increases, the nugget grows further in the horizontal direction and creates more of an elliptical weld nugget commonly seen in the other welding stack-ups. Each of the welds shown in Figure 80 meet the minimum size requirements with nuggets of 4.82, 5.54, and 7.14 mm respectively. The nugget diameters 136

157 were taken as an average of the diameters at the geometrical center point and each interface. It can be seen from the shrinkage pores and dendrite growth direction that the nugget center point is located at the geometrical center of the nugget. In this stack-up, this is in the center of the middle sheet. In the largest nugget, which was made with the longest weld time, a notch-like feature is created at the interface of the JSC 590 and Usibor A similar phenomenon was seen in the 2T Usibor welds. 7.4 Conclusions Welding complex stack-ups cannot simply follow many of the guidelines written for common welding configurations. With dissimilar materials and different thicknesses, more knowledge must be generated on how the weld nugget grows and is effected by asymmetrical stack-ups. By looking at the three stack-ups of interest, information was gathered on how time and current effect the diameter and penetration of the weld nugget. Ample time is needed to grow the weld nugget in the horizontal direction. The diameter is largely effected by the amount of time that current is being passed through the stack-up. When a thin sheet is present on the outside of the stack-up, sufficient current is needed to penetrate the weld nugget into the thin sheet. Another key factor that may affect the growth of nugget into the thin sheet is the low contact resistance, and thus low Joules heating at the thin/thick sheet interface due to the galvanneal coating on the thin sheet. In the literature, a penetration depth of 50% has been deemed acceptable in unequal thickness spot welds. By adding extra sheets and shifting the thin/thick interface further 137

158 away from the geometrical center point, more current is required to reach the minimum penetration value. This work provided pertinent information regarding the effect of two common welding variables on the nugget formation of complex stack-ups. More work can be completed to determine how complex welding schedules can better accommodate complex stack-up configurations for use in the automotive industry. 138

159 Chapter 8: Summary and Conclusions In summary, the deformation behavior of the resistance spot welds of Usibor 1500 comprising inhomogeneous microstructure (especially weld metal, CGHAZ, and SCHAZ) was studied using a combination of experimental and computational tools including Gleeble heat treatment, microstructural characterization, conventional and innovative mechanical testing methods, and FEA simulation. DIC was used for mechanical testing to develop surface strain maps to directly compare to simulated tests. 8.1 Conclusions The following conclusions can be drawn from the work presented: Ultra-High Strength Steel Characterization The CGHAZ and FGHAZ encompass the solidified weld nugget. The microhardness of these regions is like that of the base metal and weld nugget due to the martensitic microstructure. There is a significant prior austenite grain growth in the CGHAZ due to extended time above the AC3 temperature, while the FGHAZ has a finer grain size. The SCHAZ is located as a 0.7-mm-wide annulus immediately adjacent to the unaffected base metal. The micro-hardness for SCHAZ is about 300 HVN, significantly lower than that (500 HVN) for the base metal. The microstructure in 139

160 this region is composed of over-tempered martensite that has decomposed into a mixture of ferrite and cementite. The simulated welding heat treatment in Gleeble successfully reproduces CGHAZ and SCHAZ in the tensile samples, with microstructure and microhardness results consistent with those of actual welds. Tensile testing of Gleeble simulated SCHAZ shows its yield strength is 618 MPa, and ultimate tensile strength 866 MPa. The CGHAZ shows its yield strength as 1342 MPa, and ultimate tensile strength of 1811 MPa. The softened SCHAZ is markedly lower and the CGHAZ is slightly higher than the Usibor 1500 base metal which yield and tensile strength is 1179 and 1464 MPa, respectively Conventional Mechanical Testing Methods In pure tension, the softened SCHAZ accumulates the bulk of deformation and is the location of crack initiation. The inferior properties of this region cause failure at a stress of 1348 MPa, which is slightly lower than the base metal ultimate strength. In tensile shear testing, there are two different locations competing against each other for failure: the notch-like stress concentration at the joint interface and the softened SCHAZ. In tensile shear testing, all welds with a minimum nugget diameter greater than 4 t failed well above the AWS required strength, regardless of the failure mode. A sweet spot was found below the expulsion point that created button pull-out 140

161 failure through the SCHAZ. This failure mode allowed for the highest energy absorption with an average maximum load of 30.4 kn and elongation of 1.62%. The tensile shear test can be modified by sectioning the weld to allow for in-situ observation of the deformation occurring during testing. Considering the microstructure-specific mechanical properties in FEA model is essential to the accurate simulation of deformation behavior of resistance spot welds. By just using a homogeneous property, strain accumulation in the SCHAZ cannot be correctly predicted by the models. The strain accumulation in a weldment is highly dependent on the geometry at the interface. The notch feature geometry is difficult to predict but plays an important role in failure mode Single-Sided Wedge Testing The single-sided wedge testing of half spot weld samples reveals the plastic deformation and fracture on the transverse section. Once the weld size grows to greater than 5.8 mm, the failure shifts from abrupt interfacial to more complex, energy absorbing modes. For the 1.5-mm thick Usibor 1500 material, this nugget diameter is much greater than typical industry standards of 4 t. As the leg size increases in single-sided wedge samples with a weld nugget greater than 5.8 mm, the failure mode shifts toward the weld nugget. Longer legs allow the strong base metal to differ a bending moment in the softened SCHAZ and apply the load primarily to the notch-like feature at the interface. 141

162 By eliminating the pulsing and changing the grain structure at the edge of the weld nugget, the maximum weld size needed to ensure interfacial failure becomes undistinguishable. Further testing needs to be conducted to see the applicability and repeatability of single-sided wedge testing for various material combinations Complex Stack-Up Analysis The weld time dictates the horizontal growth of the weld nugget. The diameter of the nugget is largely effected by the amount of time in which welding current is applied. When a thin sheet is present on the outside of a stack-up, additional current is needed to grow the weld nugget into the thin sheet. The difficulty is exacerbated by the soft and conductive coating on the weaker thin sheet steel. By increasing the current, there is vertical growth of the nugget to a threshold limit. When the overall thickness of the joint is increased and the thin/thick interface is shifter further from the geometrical center point of the weldment, it is more difficult to create ample penetration of the nugget to the thin sheet. More current is required to increase the horizontal growth of the nugget. 8.2 Recommendations for Future Work While this work determines many of the deformation and failure behaviors of resistance spot welded ultra-high strength steel more work can be completed in the following areas: 142

163 8.2.1 Failure Criterion for Computational Modeling Develop a set of failure parameters that can be extrapolated to more complex models but better represent the failure modes seen in experimental testing. Preform more mechanical testing to verify the failure parameters are consistent in more complex loading conditions Single-Sided Wedge Testing Further testing optimization can be completed to determine how to account for misalignment of the spot weld and wedge device. Different wedge geometries that have a blunted tip for insertion may help mitigate the loading of one leg. A lubricant can be used to create a more uniform loading between both legs of the weldment. Testing can be completed on different stack-ups of interest to determine the applicability of the testing procedure to a wide range of material combinations Complex Stack-Up Analysis An MFDC welding machine can be used with better control for welding complex stack-ups. A similar study can be completed using advanced weld schedule solutions like pulsation, current sloping, and cool time to determine the effects on nugget growth, much like the work done by Gould et al. in Figure 28. The use of alternative welding electrodes can be used to develop a more robust welding schedule. A smaller electrode face on the thin sheet can allow for more heat generation at the thin/thick interface. 143

164 8.3 Impact of Research Ultra-high strength steels, especially boron steels, are integral structural materials need to further light-weight vehicles while still maintaining crash safety requirements. A better understanding of the deformation and fracture behavior of spot-welded joints is needed to develop more accurate methods for predicting failure. Such accurate modeling techniques are essential for the automotive industry to implement CAE-driven design and manufacturing to shorten design cycles and reduce the amount of crash test vehicles. The knowledge generated in this study constitutes a first step toward an improved understanding of the deformation and failure behavior of resistance spot welds of ultrahigh strength steels and complex stack-ups. 144

165 References [1] WorldAutoSteel, "Advanced High-Strength Steels Application Guidelines," unpublished. [2] S. Dancette et al, "HAZ microstructures and local mechanical properties of high strength steels resistance spot welds," ISIJ Int, vol. 51, pp , [3] M. Khan, M. Kuntz and Y. Zhou, "Effects of weld microstructure on static and impact performance of resistance spot welded joints in advanced high strength steels," Science and Technology of Welding & Joining, [4] E. Biro et al, "Predicting transient softening in the sub-critical heat-affected zone of dual-phase and martensitic steel welds," ISIJ Int, vol. 53, pp , [5] B. V. Hernandez et al, "Influence of microstructure and weld size on the mechanical behaviour of dissimilar AHSS resistance spot welds," Science and Technology of Welding and Joining, vol. 13, pp , [6] Y. Jong et al, "Microstructural evolution and mechanical properties of resistance spot welded ultra high strength steel containing boron," Materials Transactions, vol. 52, pp , [7] S. Vignier, E. Biro and M. Hervé, "Predicting the hardness profile across resistance spot welds in martensitic steels," Weld. World, vol. 58, pp , [8] Crashworthiness. NHTSA. N.p Web. [9] S. Burget and S. Sommer, "Modeling of deformation and failure behavior of dissimilar resistance spot welded joints under shear, axial and combined loading conditions," in Icf13, 2013,. [10] T. Coon et al, "Resistance spot weldability of three metal stack dual phase 600 hotdipped galvanized steel," SAE Technical Paper Series, vol. 01, [11] H. Eizadi and S. Marashi, "On the resistance spot welding of four-sheet stack of unequal sheet thickness," Science and Technology of Welding and Joining, pp. 1-6,

166 [12] J. Gould, W. Peterson and J. Cruz, "An examination of electric servo-guns for the resistance spot welding of complex stack-ups," Weld. World, vol. 57, pp , [13] C. V. Nielsen et al, "Three-sheet spot welding of advanced high-strength steels," Welding Journal, vol. 90, [14] T. Eller et al, "Identification of plasticity model parameters of the heat-affected zone in resistance spot welded martensitic boron steel," [15] M. Greitmann, "Welding through the Ages," unpublished. [16] ISO Standard, "Resistance welding - weldability, part 2 : Alternative procedures for the assessment of sheet steels for spot welding," Tech. Rep. ISO , [17] AWS Standards, "Specification for automotive weld quality - resistance spot welding of steel," Tech. Rep. AWS D8.1M, [18] Honda Research & Development, "Honda Engineering Standards," unpublished. [19] W. Savage, E. Nippes and F. Wassell, "Static contact resistance of series spot welds," Welding Journal, vol. 56, pp. 365s-370s, [20] H. Zhang and J. Senkara, Resistance Welding: Fundamentals and Applications. Taylor & Francis, [21] N. Williams and J. Parker, "Review of resistance spot welding of steel sheets Part 1 Modelling and control of weld nugget formation," International Materials Reviews, vol. 49, pp , [22] Resistance Welder Manufacturers' Association, Resistance Welding Manual. Resistance Welder Manufacturers' Association, [23] I. Khan et al, "Monitoring the Effect of RSW Pulsing on AHSS using FEA (SORPAS) Software," SAE Technical Paper Series, vol. 01, [24] W. Li, D. Cerjanec and G. A. Grzadzinski, "A comparative study of single-phase AC and multiphase DC resistance spot welding," Journal of Manufacturing Science and Engineering, vol. 127, pp , [25] K. Hofman et al, "AC or DC for resistance welding dual-phase 600?" Welding Journal, vol. 84, pp ,

167 [26] W. Li et al, "Energy consumption in AC and MFDC resistance spot welding," in Proceedings of the Sheet Metal Welding Conference XI, Sterling Heights, Michigan, 2004,. [27] Z. Mikno, S. Kowieski and A. Pilarczyk, "Analysis of liquid metal expulsion from the weld nugget," in 9th International Seminar & Conference on Advances in Resistance Welding, Miami, 2016,. [28] D. Dickinson, Welding in the Automotive Industry: State of the Art: A Report. Republic Steel Research Center, [29] W. Chuko and J. Gould, "Development of appropriate resistance spot welding practice for transformation-hardened steels," American Iron and Steel Institute, Pittsburgh, Tech. Rep. DE-FC07-97ID13554, [30] P. Barthelemy, "Servo weld gun present and future," in Sheet Metal Welding Conference XI, Detroit AWS Section, Detroit, Paper, 2004, pp [31] C. Anderson, C. Wiermaa and M. Morel, "Developments in resistance spot welding," Practical Welding Today, vol. 4, pp , [32] C. Tsai et al, "Modeling of resistance spot weld nugget growth," Welding Journal(USA), vol. 71, pp. 47, [33] S. Na and S. Park, "A theoretical study on electrical and thermal response in resistance spot welding," Welding Journal-Including Welding Research Supplement, vol. 75, pp , [34] H. Nied, "The finite element modeling of the resistance spot welding process," Weld.J., vol. 63, pp. 123, [35] Y. Cho and S. Rhee, "Experimental study of nugget formation in resistance spot welding," Welding Journal, vol. 82, pp , [36] M. Khan, "Spot Welding of Advanced HIgh Strength Steels (AHSS).", UWSpace, [37] N. J. Den Uijl, "Resistance Spot Welding of Advanced High Strength Steels.", TU Delft, Delft University of Technology, Netherlands, [38] J. Buckley and R. Servent, "Improvements in ultrasonic inspection of resistance spot welds," Insight-Non-Destructive Testing and Condition Monitoring, vol. 51, pp ,

168 [39] K. Yeung and P. Thornton, "Transient thermal analysis of spot welding electrodes," Welding Journal-New York-, vol. 78, pp. 1-s, [40] K. Easterling, Introduction to the Physical Metallurgy of Welding. Elsevier, [41] S. Kou, "Welding Metallurgy," New York, vol. 2, [42] M. Naderi et al, "A numerical and experimental investigation into hot stamping of boron alloyed heat treated steels," Steel Research International, vol. 79, pp. 77, [43] Honda Research & Development, "Hot Stamped Boron Steel (Usibor 1500) Compositional Data," unpublished. [44] H. Karbasian and A. E. Tekkaya, "A review on hot stamping," J. Mater. Process. Technol., vol. 210, pp , 11/19, [45] V. B. Hernandez, S. Nayak and Y. Zhou, "Tempering of martensite in dual-phase steels and its effects on softening behavior," Metallurgical and Materials Transactions A, vol. 42, pp , [46] Honda Research & Development, "Certificate of Conformance JSC980YL," unpublished. [47] K. Osawa, Y. Suzuki and S. Tanaka, "TS MPa grade low-carbon equivalent type galvannealed sheet steels with superior spot-weldability," Kawasaki Steel Technical Report, pp. 9-16, [48] R. Rocha et al, "Microstructural evolution at the initial stages of continuous annealing of cold rolled dual-phase steel," Materials Science and Engineering: A, vol. 391, pp , [49] S. Wei et al, "Effect of joint configuration on resistance spot weldability of galvanised DP780 steel sheets," Science and Technology of Welding and Joining, vol. 21, pp , [50] S. Wei et al, "Weldability and mechanical properties of similar and dissimilar resistance spot welds of three-layer advanced high strength steels," Science and Technology of Welding and Joining, vol. 20, pp , [51] P. Ghosh et al, "Weldability of intercritical annealed dual-phase steel with the resistance spot welding process," Welding Journal, vol. 70, pp. 7,

169 [52] H. Choi et al, "Evaluation of weldability for resistance spot welded single-lap joint between GA780DP and hot-stamped 22MnB5 steel sheets," Journal of Mechanical Science and Technology, vol. 25, pp , [53] Y. Omiya and M. Kamura, "Characteristics of 590MPa Grade Low YP Type Hot Dip Galvannealed Steel Sheet," Research and Development-Kobe-, vol. 52, pp , [54] T. B. Hilditch et al, "Experimental evaluation of curl and tensile properties of advanced high strength sheet steels," SAE Technical Paper Series, vol. 113, [55] I. Precision Strip, "Chemistry Report by Heat-Lot," unpublished. [56] Honda Research & Development, "Certificate of Conformance JAC270D 45/45," unpublished. [57] Galvanneal - Differences from Galvanize. GalvInfoNote. N.p. Rev 1.2 Apr [58] M. Khan et al, "Welding behaviour, microstructure and mechanical properties of dissimilar resistance spot welds between galvannealed HSLA350 and DP600 steels," Science and Technology of Welding and Joining, vol. 14, pp , [59] W. D. Callister and D. G. Rethwisch, Materials Science and Engineering. John Wiley & Sons NY, [60] Y. J. Chao, "Ultimate strength and failure mechanism of resistance spot weld subjected to tensile, shear, or combined tensile/shear loads," Journal of Engineering Materials and Technology, vol. 125, pp , [61] S. Dancette et al, "Experimental and modeling investigation of the failure resistance of Advanced High Strength Steels spot welds," Eng. Fract. Mech., vol. 78, pp , 7, [62] P. Wung et al, "Failure of spot welds under in-plane static loading," Exp. Mech., vol. 41, pp , [63] H. Ghassemi et al, "Quasi-static spot weld strength of advanced high strength sheet steels," in 9th International Seminar & Conference on Advances in Resistance Welding, Miami, 2016,. [64] D. Radakovic and M. Tumuluru, "An evaluation of the cross-tension test of resistance spot welds in high-strength dual-phase steels," Welding Journal, vol. 91, pp. 8-15,

170 [65] C. Tsai et al, "Mechanical Behaviors of Resistance Spot Welds During Pry- Checking Test," Supplement to Welding Journal, June2006, [66] G. Payen et al, "Design of an in situ mechanical test for spot-welded joints," Eng. Fract. Mech., vol. 96, pp , [67] S. Zuniga and S. D. Sheppard, "Resistance spot weld failure loads and modes in overload conditions," in Fatigue and Fracture Mechanics: 27th VolumeAnonymous ASTM International, 1997,. [68] Y. Chao, "Failure mode of spot welds: interfacial versus pullout," Science and Technology of Welding and Joining, [69] J. Blaber, B. Adair and A. Antoniou, "Ncorr: open-source 2D digital image correlation matlab software," Exp. Mech., vol. 55, pp , [70] A. Kammers and S. Daly, "Small-scale patterning methods for digital image correlation under scanning electron microscopy," Measurement Science and Technology, vol. 22, pp , [71] R. Harilal and M. Ramji, "Adaptation of open source 2D DIC software Ncorr for solid mechanics applications," 9th International Symposium on Advanced Science and Technology in Experimental Mechanics, [72] J. Fang et al, "Weld modeling with MSC/nastran," in Second MSC Worldwide Automotive User Conference, Dearborn, MI, 2000,. [73] A. Zaouk, D. Marzougui and C. Kan, "Development of a detailed vehicle finite element model Part II: Material characterization and component testing," International Journal of Crashworthiness, vol. 5, pp , [74] E. H. Lamouroux et al, "Detailed model of spot-welded joints to simulate the failure of car assemblies," International Journal on Interactive Design and Manufacturing (IJIDeM), vol. 1, pp , [75] P. Wung, "A force-based failure criterion for spot weld design," Exp. Mech., vol. 41, pp , [76] F. Seeger et al, "An investigation on spot weld modelling for crash simulation with ls-dyna," in 4th LS-DYNA User Forum, Bamberg, 2005,. [77] M. Palmonella et al, "Guidelines for the implementation of the CWELD and ACM2 spot weld models in structural dynamics," Finite Elements Anal. Des., vol. 41, pp ,

171 [78] W. W. Jung, Y. Dal Kwon and S. S. Kang, "Selecting the spot welding condition of multi-layer vehicle structure," SAE Technical Paper , [79] J. Shen et al, "Modeling of resistance spot welding of multiple stacks of steel sheets," Mater Des, vol. 32, pp , [80] M. Pouranvari and S. Marashi, "Critical sheet thickness for weld nugget growth during resistance spot welding of three steel sheets," Science and Technology of Welding and Joining, vol. 16, pp , [81] F. Schreyer and S. Weihe, "Identification of resistance spot welding parameters for ultra-high-strength steels by use of finite element calculations," in METEC & 2nd ESTAD, Dusseldorf, 2015,. [82] R. Hertzberg, R. Vinci and J. Hertzberg, Deformation and Fracture Mechanics of Engineering Materials. John Wiley & Sons, Inc., [83] G. Huang, H. Zhu and B. Yan, "A Novel Approach for Generating a Full-Range Tensile Stress-Strain Curve," vol. 01, [84] S. Zhang, "Fracture mechanics solutions to spot welds," Int. J. Fract., vol. 112, pp , [85] S. Burget and S. Sommer, "Characterization and modeling of fracture behavior of spot welded joints in hot-stamped ultra-high strength steels," in LS-DYNA Forum, Ulm, 2012,. [86] P. Eftekharimilani et al, "The microstructural evolution and elemental distribution of a 3rd generation 1 GPa advanced high strength steel during double pulse resistance spot welding," unpublished. 151

172 Appendix A: Calculations Load vs. Stroke Curves for MTS System To perform a tension test with the MTS testing system, the final displacement, testing time, and recording frequency is input into the programming software. After the test, an output file generates an index and load (Figure 81). To determine the stroke of the testing machine, it is assumed that the sample is pulled at a constant loading rate. The stroke can be calculated by using the equation below. Stroke = Index Final Displacement Frequency Testing Period (8) The index is a unit-less number that give reference to the specific output line. The frequency tells how many output values should be recorded every second and is entered in hertz. A hertz is defined as one cycle per second, so a value of 50 hertz would mean output is being collected 50 times every second. The final displacement is input by the user and is assumed to be reached at the ending of the test. The testing period is entered by the user and coincides with the index and frequency to cancel out the time increment of the calculation. The value of the displacement can be represented by any unit of distance. In most cases, the displacement of the test is small and can be described best by inches or millimeters. 152

173 Figure 81: Example of the output file from the MTS system Constitutive Behavior Development This section gives more details on the procedure of developing constitutive behaviors using the Gleeble simulated microstructural samples. The images displayed are for the CGHAZ, but the same procedure was used for base metal and SCHAZ samples. The data is exported from Post_Ncorr and graphed using Excel (Figure 82). The y- displacement of the sample for the given extensometer size is measured against the image number. For all testing, images were taken at a second interval. Therefore, the image number coincides with the time increment in seconds. A polynomial line is fit to the data to explain the relationship between displacement and time. It is important to ensure that the line is well suited for the data by checking the R-squared value. 153