Tool Wear Prediction Modelling for Sheet Metal Stamping Die in Automotive Manufacture

Transcription

1 Tool Wear Prediction Modelling for Sheet Metal Stamping Die in Automotive Manufacture by Xuan Zhi WANG A thesis submitted for full fulfilment of the requirement for the degree of Doctor of Philosophy Faculty of Engineering and Industrial Sciences, Swinburne University of Technology, Hawthorn, Victoria 3122, Australia March 2011

2 DECLARATION This thesis contains no material which has been accepted for the award of any other degree or diploma at any university and to the best of my knowledge and belief contains no material previously published or written by another person or persons excepts where due reference is made. Xuan Zhi WANG 24 March 2011 i

3 ACKNOWLEDGMENTS I would like to express my sincere thanks for all who have contributed to this thesis. First of all, I wish to thank my supervisor Prof Syed Masood. This thesis would not have been possible without his great and valuable support and guidance. I hereby express my special thanks to my co-supervisor Dr Matthew Dingle from Deakin University. Here particular thanks to Dr Tim Hilditch and Dr Matthias Weiss from Deakin University for their helps in preparing the channel bending tests. I would like to show my gratitude to the Cooperative Research Centre for Advanced Automotive Technology (AutoCRC) for funding my research project, especially to Ms Kate Neely from AutoCRC for her kind support for the project. I would also be thankful for GM Holden for providing samples and technical documents. Mr Shane Christian from GM Holden deserves special thanks for his assistance and coordination for the project. I am grateful to my parents for their continuous support thorough my life. They always encourage me to achieve my goals in my life, especially in some tough time. Special thank to my uncle for his valuable support. I would like to thank my colleagues from Faculty of Engineering & Industrial Sciences, Swinburne University of Technology. I would also like to thank my friends in Melbourne who helped me during my study. ii

4 ABSTRACT Advanced high strength steels (AHSS) are increasingly used in sheet metal stamping in the automotive industry. In comparison with conventional steels, AHSS stampings produce higher contact pressures at the interface between draw die and sheet metal blank, resulting in more severe wear conditions, particularly at the draw die radius. Developing the ability to accurately predict and reduce the potential tool wear during the tool design stage is vital for shortening lead times and reducing production costs. This thesis investigates the influence of draw die geometry on the wear distribution over the draw die radius for AHSS and develops a methodology for optimising the draw die geometry to reduce wear using numerical and experimental methods. Tool wear predictions on automotive sheet metal forming die and recommended protections of the tool surface under the initial production conditions were obtained from AutoForm simulation software. Effects of lubrication coefficients, binder pressure loads and die coating on tool wear distributions were investigated as well. It is concluded that the areas that are most sensitive to tool wear occurs at the locations corresponding to the large gradient of drawing depth. To study the tool wear distributions for more common stamping parts, a numerical tool wear model was developed and applied using the commercial software package Abaqus. Channel tests are carried out using an Erichsen sheet metal tester with high pressure prescale films to verify the numerical model results. Comparing the results obtained from the prescale film with the results from the simulation, it is concluded that the contact pressure distributions indicated by the prescale film are consistent with those from the simulation. iii

5 Various geometries of radius arc profiles, including standard circular profiles, high elliptical profiles, and flat elliptical profiles, were numerically investigated using the tool wear model developed, and the contact pressure distribution and tool wear work along the radii were determined. The following conclusions were reached from the investigations: (1) The colour contour of the high contact pressure on the die radius can be divided into three distinct zones of high pressure and tool wear; (2) The dominant zone leading to maximum contact pressure and tool wear severity depends on the geometry of die radius profile under the same material and process conditions; (3) The geometry of draw die radius has a significant influence on the tool wear and standard circular and elliptical curves can lead to the achievement of reduced and uniform contact pressure distribution (wear distribution) along most of zones of the draw die radius arc. The results suggest that to minimise contact pressure and tool wear using this approach it would be necessary to optimise the shape for a particular combination of material type, thickness and forming process. Effects of control parameters, such as blank geometry, punch geometry, deep-drawing process parameters and tool material, on wear behaviour in deep-drawing for various shape of die radius were then investigated to provide guidelines for impacts of these parameters. A specialised software routine was then compiled for optimisation of die radius profiles to minimise and achieve uniform contact pressure (wear distribution) iv

6 using Python programming language. The routine was fully integrated with Abaqus software and has the following functions: (1) To provide a user-friendly Graphical User Interface for pre-processing data input for users who have less experience and skill; (2) To optimise a die radius profile according to the control parameters that users input. The results obtained are relevant to the issue of reducing the high tool wear in automotive stamping tools by predicting the causes of such tool wear related to tool geometry and process parameters. They provide useful guidelines for enhancing the tool life of sheet metal processing in automotive industry. v

7 TABLE OF CONTENTS DECLARATION... i ACKNOWLEDGMENTS... ii ABSTRACT... iii TABLE OF CONTENTS... vi LIST OF FIGURES... xii LIST OF TABLES... xvii CHAPTER 1 INTRODUCTION Background and Significance of Research Objectives and Scope of Research Outlines of Thesis... 5 CHAPTER 2 LITERATURE REVIEW Overview Sheet Metal Stamping Introduction Contact zones in sheet metal forming Advanced High Strength Steel Dual phase (DP) steel Tool wear in stamping of AHSS Tool Wear Mechanism Introduction vi

8 2.4.2 Adhesive wear Abrasive wear Tool wear model for conventional deep-drawing Test Methods for Tool Wear Prediction Pin-on-disk test Modified bending-under-tension-test Bending-under-tension test Deep-drawing process-simulator Slider-on-flat-surface tribometer Twist compression test U-bending test Strip-drawing test Draw bead test Slider test system Acoustic emission technique Research and Development in Tool Wear Coating Lubrication Alternative die materials Tool wear modelling Die radius geometry Summary vii

9 CHAPTER 3 TOOL WEAR PREDICTION USING AUTOFORM SOFTWARE Introduction AutoForm software Simulation Setup Results and Discussion Identification of critical tool worn areas Relationship between tool worn area, contact pressure and drawing depth Comparison of contact pressure distribution for various lubrication coefficients Comparison of contact pressure distribution upon various binder pressure loads Comparison of tool wear distribution upon various die coating Advantages and Limitations of AutoForm Software Summary CHAPTER 4 NUMERICAL TOOL WEAR PREDICTION MODELLING Introduction Wear Work Calculation Along Die Radius Profile Finite Element Modelling Geometry Discretisation Material properties viii

10 4.3.4 Contact interaction Analysis steps with constraints and loadings Deformed and undeformed model Summary CHAPTER 5 EXPERIMENTAL VALIDATION OF TOOL WEAR PREDICTON MODEL Introduction Experimental Constraints Fuji Prescale Film Working principle of prescale film Momentary pressure measurement Experimental Equipment Erichsen sheet metal tester Fuji mono-sheet type prescale film Mild steel strip Experimental Sequences Experimental Results and Discussion Summary CHAPTER 6 INVESTIGATION OF DIE RADIUS ARC PROFILE ON WEAR BEHAVIOUR Introduction Variation of Die Radius Profiles Results and Discussion ix

11 6.3.1 Standard circular profiles High elliptical profiles Flat elliptical profiles Summary CHAPTER 7 INVESTIGATION OF CONTROL PARAMETERS ON WEAR BEHAVIOUR Introduction Variation of Control Parameters Results and Discussion Lubrication coefficient Binder holder force Young's modulus of die Clearance between die and punch Punch radius Punch diameter Blank thickness Summary CHAPTER 8 OPTIMISATION OF DIE RADIUS GEOMETRY Introduction Graphical User Interface Algorithm for Die Radius Optimisation Case Study Optimisation parameters settings x

12 8.4.2 Results and discussion Summary CHAPTER 9 CONCLUSIONS AND FURTHER RESEARCH Overview Major Research Outcomes Recommendations for Future Work REFERENCES APPENDIX A LIST OF PUBLICATIONS APPENDIX B MOMENTARY PRESSURE CHART xi

13 LIST OF FIGURES Figure 1.1 Automotive Stamping Die... 2 Figure 1.2 Body side components formed by stamping process... 2 Figure 2.1 Cross-sectional view of a simple sheet metal stamping die... 8 Figure 2.2 Contact zones in deep drawing... 9 Figure 2.3 Microstructure of DP steel Figure 2.4 Five principal types of tool failure Figure 2.5 Formation of an adhesive junction Figure 2.6 Schematic of a hypothetical model of generation of a hemispherical wear particle during a sliding contact Figure 2.7 A hard conical asperity in sliding contact with a softer surface in an abrasive wear model Figure 2.8 Pin-on-disk test Figure 2.9 A modified bending under tension test Figure 2.10 Schematic of bending-under-tension test Figure 2.11 Schematic of deep-drawing process-simulator Figure 2.12 Schematic presentation of the SOFS tribometer Figure 2.13 Schematic of twist compression test Figure 2.14 Temperature measurement using a thermocouple Figure 2.15 Schematic view of U-bending test Figure 2.16 U-bending equipment showing die-holder with inserts Figure 2.17 Principle for U-bending test xii

14 Figure 2.18 Strip-drawing test Figure 2.19 Strip-drawing test Figure 2.20 Draw bead test Figure 2.21 Slider test system Figure 2.22 Die sample dimensions and its actual photo on wear tracks Figure 2.23 Robot-based die wear test system Figure 2.24 Simulation testing machine for hot stamping Figure 2.25 Scheme of strip drawing test Figure 2.26 Algorithm applied in UGS Figure 3.1 Sheet metal forming process chain in AutoForm Software Figure 3.2 Reinforced rear suspension support Figure 3.3 Forming limit curve Figure 3.4 Simulation sequences in AutoForm Figure 3.5 Blank, binder, punch and die in AutoForm software Figure 3.6 Potential tool worn area location on die surface obtained from initial simulation Figure 3.7 Photos of the worn areas on the corresponding surface of actual sheet metal part Figure 3.8 Cross-section 1 on Areas 2 and Figure 3.9 Contact pressure distribution and drawing depth at Cross section Figure 3.10 Contact pressure distributions upon various lubrication coefficients along Cross-section Figure 3.11 Contact pressure distributions upon various binder pressure xiii

15 loads along Cross-section Figure 3.12 Tool wear distributions upon various die coating method with lubrication coefficient Figure 3.13 Maximum production volume until the occurrence of local wear along Cross-section Figure 3.14 Comparison of major strain results from AutoForm and Abaqus software with experimental results Figure 4.1 Meshed finite element model in the deep-drawing simulations Figure 4.2 Dimensions and parameters of finite element model Figure 4.3 Undeformed model after Step Figure 4.4 Deformed model after Step Figure 4.5 Deformed model during Step 5 in early stage Figure 4.6 Deformed model during Step 5 in middle stage Figure 4.7 Deformed model during Step 5 in late stage Figure 4.8 Fully-deformed model after Step Figure 5.1 Fuji prescale film types and corresponding pressure range Figure 5.2 Erichsen sheet metal tester Figure 5.3 Schematic of Erichsen sheet metal tester Figure 5.4 Steps in channel test (black: mild steel strip, red: prescale films).. 99 Figure 5.5 Placement of prescale film Figure 5.6 Comparison of contact pressure distributions obtained from tests and simulations Figure 6.1 Three regular types of die radius profile Figure 6.2 Contact pressure over die radius with CR5 profile xiv

16 Figure 6.3 Contact pressure over die radius with CR10 profile Figure 6.4 Contact pressure over die radius with CR15 profile Figure 6.5 Cause of high contact pressure of standard circular profiles Figure 6.6 Wear work over die radius with standard circular profiles Figure 6.7 Contact pressure over die radius with HER5r10 profile Figure 6.8 Contact pressure over die radius with HER5r15 profile Figure 6.9 Cause of high contact pressure of high elliptical profile Figure 6.10 Wear work over die radius with high elliptical profile Figure 6.11 Contact pressure over die radius with FER10r5 profile Figure 6.12 Contact pressure over die radius with FER15r5 profile Figure 6.13 Cause of high contact pressure of flat elliptical profile Figure 6.14 Wear work over die radius with flat elliptical profiles Figure 7.1 Wear work over die radius with various lubrication coefficients for three die radius arc profiles Figure 7.2 Wear work over die radius with various binder holder forces for three die radius arc profiles Figure 7.3 Wear work over die radius with various Young s modulus of die for three die radius arc profiles Figure 7.4 Wear work over die radius with various clearances between die and punch for three die radius arc profiles Figure 7.5 Wear work over die radius with various punch radius for three die radius arc profiles Figure 7.6 Wear work over die radius with various punch diameters for three die radius arc profiles xv

17 Figure 7.7 Wear work over die radius with various blank thicknesses for three die radius arc profiles Figure 8.1 GUI for Geometry created using Python programming language Figure 8.2 GUI for Process Parameters created using Python programming language Figure 8.3 GUI for Simulation Setting created using Python programming language Figure 8.4 Die radius profile Figure 8.5 Accumulated wear work along die radius Figure 8.6 Flow chart of proposed algorithm Figure 8.7 Divisions of die radius profile Figure 8.8 Positions of division points of optimised curve Figure 8.9 Wear work over die radius for CR5 and optimised curves xvi

18 LIST OF TABLES Table 3.1 Material properties of reinforce rear suspension support Table 4.1 Material properties of mild steel blank and die Table 5.1 Material properties and fitted values of K, e, n of mild steel Table 5.2 Comparison of locations of contact pressure peaks from Table 6.1 Material properties of DP780 blank and die Table 6.2 Various die radius profiles used in simulations Table 7.1 Material properties of DP780 blank and die Table 7.2 Die radius profiles in simulations Table 7.3 Control parameters in simulations Table 7.4 Impacts of control parameters on wear work Table 8.1 Material properties of DP780 blank and die Table 8.2 Effective radius R for CR5 and optimised curves Table 8.3 Comparison of wear work of un-optimised circular profile with optimised one xvii

19 CHAPTER 1 INTRODUCTION 1.1 Background and Significance of Research Sheet metal stamping is a process to stretch a part over a punch of complicated shape in a draw die [1]. Due to its efficiency in bulk forming operations, sheet metal stamping is widely implemented in automobile industries to convert sheet metal into exterior and interior parts, such as auto-body panels and a variety of appliance parts, with prescribed sizes and shapes (Figures 1.1 and 1.2). A rapidly changing automobile market demands high precision and perfect appearance of finished parts, soft flexibility of new materials as well as shortened lead time and decreased production costs in whole production-cycles. Compared with other parts produced by bulk forming operations, the automobile parts with complex three-dimensional shapes are desired to meet (i) high dimensional accuracy to ensure the compatibility and interchangeability in subsequent welding and painting operations, and (ii) perfect surface appearance, especially for exterior auto-body panel, to eliminate wrinkle, corrugation, indentation and scratching. As the material flow in sheet metal forming mainly depends on the sliding and bending friction between a workpiece and corresponding die/punch, wear of tools caused by high normal contact pressure and sliding distances could seriously influence dimensional accuracy and surface appearance of finished parts, which results in high scrap rate of workpieces. 1

20 Figure 1.1 Automotive Stamping Die [2] Figure 1.2 Body side components formed by stamping process [3] Emerging new materials, such as advanced high strength steels (AHSS), are used in sheet metal stamping, which involves higher contact pressure and temperature at the tool-workpiece interface than conventional materials. It leads to increased possibility of potential tool wear and decreased maximum production volumes without occurrence of tool wear if suitable protection measurements, such as 2

21 coatings and hardness treatments, are not applied under efficient investigation of the mechanism of tool wear. It is impossible to shorten lead time and decrease production cost without determination of the extent of tool wear. Unexpected tool changes caused by wear usually result in unacceptable down times and increased die maintenance cost with extra budgets. Thus, tool wear of sheet metal stamping dies is becoming a major obstacle for industries to meet the above demands from automotive markets. Due to the complicated geometric, material and nonlinear contact characteristics in the deformation of automotive parts, it is rather time-consuming and costly to research the mechanism of tool wear and predict the extent of tool wear by means of try-out techniques based solely on conventional trial and error and engineers experiences. To overcome limitations of traditional methods, a prediction model is required to be established based on numerical simulations and validated by wear tests. 1.2 Objectives and Scope of Research A typical die assembly of sheet metal stamping consists of a punch, a draw die and a binder. During a stamping operation, several contact pairs are established between one of the components in the die assembly and sheet metal blank. The contact pair of draw die and sheet metal blank is the most critical pair, because both stretching and bending occur in the contact zone formed by the pair. The tool surface in the contact zone is exposed to severe wear conditions with high contact 3

22 pressure and long sliding distance compared with the tool surface in other contact pairs. Early work has shown that modification of the geometry of the draw die profile could improve the contact condition between the draw die and sheet metal blank and reduce the tool wear. Several researchers have investigated relationship between various draw die profiles and tool wear distribution. Due to the limitation and diversity in their experiments and numerical simulations in early years, some results obtained from these researches are not consistent with each other [4, 5]. Moreover, these previous researches were mainly limited to sheet metal stamping with conventional materials. To overcome these limitations, this research presents a comprehensive investigation, employing the latest experimental equipments and numerical simulation technologies, to study the influence of draw die geometry on the wear distribution over the draw die radius for advanced high strength steels (AHSS). The work presents a methodology for optimising the draw die geometry to reduce wear using numerical and experimental methods. Specifically, the research aims to achieve the following objectives: (1) To predict and identify critical tool worn area on GM Holden s sheet metal forming die using AutoForm simulation software; (2) To establish a numerical tool wear prediction model of deep-drawing process using Abaqus simulation software for a common part and perform experimental validation by a series of channel bending test; 4

23 (3) To determine the relationship between different die profile shape and tool wear distribution for deep-drawing process; (4) To determine the relationship between different control parameters (with the same type shape, e.g. elliptical, circular) and tool wear distribution for deep-drawing process; (5) To develop a specialised algorithm for achieving minimised and uniform wear distribution by changing the die profile shape for deep-drawing process using Python programming language. 1.3 Outlines of Thesis The thesis is composed of nine chapters, eight of which follow on from this introduction. Chapter 2 conducts a literature review of the current status of researches and developments in the area of tool wear prediction for sheet metal stamping. Chapter 3 presents tool wear predictions on a particular automotive sheet metal forming die and recommended protections of the tool surface under the initial production conditions as obtained from AutoForm simulation software. Effects of lubrication coefficients, binder pressure loads and die coating on tool wear distributions were investigated as well. It is concluded that the areas that are most sensitive to tool wear occur at the locations corresponding to the large gradient of drawing depth. Chapter 4 describes a numerical tool wear prediction model developed using the commercial software package Abaqus simulation software to study the tool wear 5

24 distributions for more common stamping parts. Chapter 5 outlines a series of channel bending tests to validate the prediction model presented in Chapter 4. The experimental equipments, procedures and validation results for testing are detailed. Chapter 6 investigates various geometries of radius arc profiles, including standard circular profiles, high elliptical profiles, and flat elliptical profiles using the tool wear prediction model developed in Chapter 4, and the contact pressure distribution and tool wear work along the radii were determined. Several significant suggestions were concluded from the investigation. Chapter 7 presents effects of control parameters, such as blank geometry, punch geometry, deep-drawing process parameters and tool material, on wear behaviour in deep-drawing for various shape of die radius, which provides guidelines for impacts of these parameters. Chapter 8 develops a specialised algorithm for optimising the die profile shape for deep-drawing process using Python programming language, which leads to a minimised and uniform tool wear distribution. Chapter 9 draws conclusions from the outcomes of the research program and details recommendations for further work to supplement the techniques outlined in this thesis. 6

25 CHAPTER 2 LITERATURE REVIEW 2.1 Overview In this Chapter, the background of sheet metal stamping is introduced in Section 2.2. Characteristics and wear behaviours of advanced high strength steel are described in Section 2.3. Then, in Section 2.4, tool wear mechanisms are briefly described. The various tool wear experimental methods are presented in Section 2.5. Section 2.6 introduces recent research and developments in tool wear prediction for sheet metal forming process using various coatings, lubricants, alternative materials, tool wear models and die radius geometries. Section 2.7 summarises the finding in the literature review and identifies the areas of research, which form the basis of the present research. 2.2 Sheet Metal Stamping Introduction Sheet metal stamping is a process of stretching a sheet metal blank over a punch of more complicated shape in a draw die [6]. A typical assembly of sheet metal stamping consists of a punch, a die and a binder. Figure 2.1 shows a simple sheet metal stamping die. A blank is clamped at the edges by the binder using one action of the press. Drawbeads on the binder surface optimise strain distributions in the 7

26 subsequent operations. The punch then travels down through the binder into the die cavity and presses the blank until the required shape of the part is formed. Figure 2.1 Cross-sectional view of a simple sheet metal stamping die [6] Contact zones in sheet metal forming One of the most common sheet metal forming operations is deep drawing as shown in Figure 2.2. In a deep drawing operation, there are five contact zones with different properties. Contact zones between the punch and the blank, as labelled as 1, 2, 3, are characterised by a low relative sliding velocity, in the order of 10-4 m/s, which means that the punch and the blank are moving at almost the same velocity. However, in Contact Zone 4 between the die and the blank and Contact Zone 5 between the blank holder and the blank, the sliding velocity is of the order from 10-3 m/s to 10-1 m/s, which is relatively high. 8

27 At contact Zone 4, i.e. the radius of the die, a combination of stretching and bending occurs and the contact pressure exceeds 100 MPa. Both boundary lubrication and mixed lubrication occur in Contact Zone 4. Boundary lubrication is a condition of lubrication in which the friction and wear behaviour are determined by the properties of the surfaces and by the properties of fluid lubricants other than their bulk viscosity, while mixed lubrication is a condition of lubrication in which the friction and wear behaviour are determined by the properties of the surfaces and by the viscous and non-viscous properties of fluid lubricants [7]. The contact condition in Contact Zone 4 is most severe in all contact zones as its predominant lubrication type is a combination of boundary lubrication and mixed lubrication [8]. Therefore, tool wear mechanism at radius portion of a die is important for tribological study of sheet metal forming. Figure 2.2 Contact zones in deep drawing [8] 2.3 Advanced High Strength Steel Dual phase (DP) steel Advanced high-strength steels (AHSS) are used extensively in the automotive 9

28 industry to help improve crash safety and reduce weight [9]. Dual Phase (DP) steel is a main type of AHSS. DP steels are low-carbon steels that contain a large amount of manganese and silicon as well as small amounts of microalloying elements, such as vanadium, titanium, molybdenum, and nickel [10]. A DP steel is created by heating a low-carbon micro-alloyed steel into the intercritical region of the Fe-C phase diagram between the A 1 and A 3 temperatures, soaking it so that austenite forms, slowly cooling it to the quench temperature, and then rapidly cooling it to transform the austenite into martensite [10, 11]. A 1 is the eutectoid temperature, which is the minimum temperature for austenite. A 3 is the lower-temperature boundary of the austenite region at low carbon contents. Upon quenching, the austenite is mostly converted to martensite, but will also partially be converted into ferrite if the cooling rate is not sufficiently high [12, 13]. Also, depending on the cooling rate, the austenite may be converted at least partially into bainite [14]. The ferrite that forms from austenite is referred to as epitaxial ferrite. The microstructure of DP steel, consisting of ferrite and martensite, is as shown in Figure 2.3 [13, 15]. Figure 2.3 Microstructure of DP steel [15] 10

29 DP steels have a bake hardening effect, which is an important benefit compared to conventional higher strength steels. The bake hardening effect is the increase in yield strength resulting from elevated temperature aging (created by the curing temperature of paint bake ovens) after prestraining (generated by the work hardening due to deformation during stamping or other manufacturing process) Tool wear in stamping of AHSS AHSS can result in severe loading, and therefore contact pressure, to traditional die structures with more than double tensile strengths [16-18] at radii and draw wall features. Such high local stresses have resulted in severe local die wear. Five principal types of tool failure related to tool wear (Figure 2.4) were reported as follows [9]: (1) Wear is damage to a solid surface involving loss or displacement of material. Wear is caused by sliding contact between the workpiece and tool. Two main types of wear are abrasive, caused by hard particles forced against and moving along a solid surface, and adhesive, caused by localized bonding between contacting solid surfaces and leading to material transfer between these surfaces. (2) Plastic deformation is caused by contact pressure exceeding the compression yield stress of the tool material. (3) Chipping is a result of stresses exceeding the fatigue strength of the tool material. 11

30 (4) Cracking is caused by stresses exceeding the fracture toughness of the tool steel. (5) Galling is a form of damage caused by sliding of two solids. It often includes plastic flow, material transfer, or both. Figure 2.4 Five principal types of tool failure [9] Billur [9] also summarised the four main factors that have an effect on these failures: (1) Contact pressure: Local contact pressure between the sheet and tool affects all types of tool failure. As stamping of AHSS requires increased contact pressure, the probability to observe tool failures increases significantly compared to stamping milder steel grades. For a given sheet material, contact pressure can be reduced by die design, such as using larger radii or reducing the sheet thickness. 12

31 (2) Surface quality: Although the surface of the tool is much smoother than the surface of the sheet, the tool s surface quality affects galling. Polishing the tool surfaces before and after coating helps to reduce galling. The sheet s roughness has little influence on tool failure. (3) Tool coating: The proper coating with a low coefficient of friction is crucial to reduce galling and tool wear. (4) Lubrication: Forming AHSS requires better-performing lubricants, possibly with extreme-pressure (EP) additives, because of the high contact pressure and temperature that occur during the process. 2.4 Tool Wear Mechanism Introduction Wear is the surface damage or removal of material from one or both of two solid surfaces in a sliding, rolling, or impact motion relative to one another. Scientific studies of wear developed little until the mid-twentieth century. In sheet metal stamping, adhesive wear and abrasive wear are two primary types of wear [19]. Raymond [20] recognised the following characteristics of wear: (1) Wear is a system property, not a material property; (2) Materials can wear by a variety of mechanisms and combinations of mechanisms, depending on the tribosystem in which it is used; (3) Wear behaviour is frequently nonlinear; 13

32 (4) Transitions can occur in wear behaviour as a function of a wide variety of parameters Adhesive wear Adhesive wear is a type of wear due to localised bonding between contacting solid surfaces leading to material transfer between two surfaces or loss from either surface [7]. Contact surfaces between a sheet metal blank and its die always exhibit some degree of roughness instead of being completely smooth. During the sliding contact between a die and a blank in a sheet metal stamping, fracture of the die usually occurs if internal stresses are so high that the fracture criterion of the material of the die is satisfied at some contact points. Figure 2.5 Formation of an adhesive junction [21] Asperities on the contact surfaces form contact spots at the interface of the blank and die. Deformations appear firstly at the contact spots characterised by high normal and tangential stresses. A series of adhesive junctions is created as a result of two contact surfaces being pressed together (Figure 2.5). Bonding occurs at these junctions and the tips of the softer asperities are sheared and adhered to the harder surface. These tips may subsequently be detached and become wear 14

33 particles or fragments. Severe types of adhesive wear are often referred to as galling, scuffing, welding or smearing. Holm [22] and Archard [23] concluded that wear volume w is generally proportional to the applied load F and sliding distance s but inversely proportional to the hardness H of the surface being worn away, so that, kfs w (2.1) H where k is the non-dimensional wear coefficient dependent on the materials in contact and their cleanliness. It is assumed that during an asperity interaction, the asperities deform plastically under the applied load and that only a wear particle will be produced at each unit. If asperities at the contact points are assumed to have an average radius of a, then, df 2 a H (2.2) Figure 2.6 Schematic of a hypothetical model of generation of a hemispherical wear particle during a sliding contact [24] 15

34 If a particle is assumed to be hemispherical in shape with radius equal to the contact radius (Figure 2.6), then, dw a (2.3) Finally, contact is assumed to remain in existence for a sliding distance ds equal to 2a, after which it is broken and the load is taken up by a new contact, so that, dw 1 df (2.4) ds 3 H 1 Fs kfs w 3 H H (2.5) Archard s equation is valid for dry contacts only. In the case of lubricated contacts, where wear is a real possibility, certain modification to Archard s equation is required [21] Abrasive wear Abrasive wear on a die surface is a common phenomenon in sheet metal stamping because the hardness of a die is larger than that of a sheet metal blank. Generally, abrasive wear is divided into two types: two-body abrasive wear and three-body abrasive wear [24]. In two-body abrasive wear, abrasive grits are embedded into one of the contact surfaces to scratch the other one, or asperities of the harder 16

35 surface slide on the softer one to damage the interface. In three-body abrasive wear, some small particles of abrasive are trapped between two surfaces but are free to move with respect to both surfaces, and are sufficiently hard to abrade one or both of the contact surfaces. In many cases, the wear mechanism starts with adhesive wear, which generates wear particles that are trapped at the interface, resulting in a three-body abrasive wear [25]. A simplified model for abrasive wear was developed by Rabinowicz [26], in which one surface consists of an array of hard conical asperities sliding on a softer and flat surface and ploughs a groove of uniform depth. Figure 2.7 shows a single conical asperity, with roughness angle of θ, creating a track through the softer surface with a depth of d and width of 2a. It is assumed that the material has yielded under the normal load df, so that, df a H (2.6) where H is the hardness of the softer surface. The wear volume w displaced in a distance s is dw a 2 s(tan ) (2.7) 2Fstan w (2.8) H 17

36 where tan is a weighted average of the tanθ values of all the individual conical asperities, called the roughness factor. Under same normal load and sliding distance, for a certain material, the larger the roughness factor is, the severer the abrasive wear occurs. df s Figure 2.7 A hard conical asperity in sliding contact with a softer surface in an abrasive wear model [24] Tool wear model for conventional deep-drawing Jensen et al [19] presented a tool wear model for conventional deep-drawing. In a deep drawing process, the blank slides over the die, resulting in tool wear mainly on the draw die profile. The normal force at a particular state in the process varies with the sliding distance on the die profile. Both the adhesive wear and the abrasive wear can be expressed as w Fs (2.9) To simplify the problem, the sliding distance is divided into small segments in which the normal force is assumed to be constant for the treated state in the process. Similarly, the process time t is also divided into small intervals in which 18

37 the normal force can be assumed constant. Thus, w n F s (2.10) x tx, tx, t 1 Because the wear depth h is more significant than the wear volume, Equation 2.10 can be expressed as below by dividing both sides by the area of each division, then, h n P s (2.11) x tx, tx, t 1 where P is the contact pressure. 2.5 Test Methods for Tool Wear Prediction Pin-on-disk test Pin-on-disk test (Figure 2.8) is a widely-used simple wear test to investigate the wear resistance of tool surfaces and surface coatings. A test ball is drawn over a disk surface with several revolutions in the same track at a pre-defined normal force and velocity. The test set up allows for the direct measurement of the normal and tangential (friction) forces during the test and by measuring the wear volume as a function of sliding distance the wear rate and the wear coefficient can be 19

38 determined [16, 26, 27]. SRV (Schwingung Reibung Verschlei reciprocating friction and wear) tester is one of several configurations of pin-on-disk test systems, and same surfaces of die and sheet materials of interest are in contact during the whole test [28]. Figure 2.8 Pin-on-disk test [16] Although the sliding speeds and normal forces can be adjusted to a level that is similar to sheet metal forming processes, the effect of plastic deformation is ignored in these tests. Therefore, the progression of tool wear in sheet metal stamping may not be presented by this test [16]. Analysis of the pin-on-disk test is standardized in ASTM G99-05, Standard Test Method for Wear Testing with a Pin-on-disk Apparatus with respect to volume loss [29]. The volume loss can be measured directly from the specimen dimensions before and after the test, or it can be calculated from mass loss. If galling is present, volume loss may not reflect the tool wear, so this test method should not be used [30]. 20

39 2.5.2 Modified bending-under-tension-test Eriksen [4] utilised a modified bending-under-tension test to investigate the influences of die edge geometry in a standard deep drawing process on the maximum wear and the wear distribution over the die edge (Figure 2.9). Figure 2.9 A modified bending under tension test [4] The test material (St 1403) 7 was wound in a coil 1. The material was drawn into the lubrication system 2 and then into the wedge dies 3. After the wedge die, the strip was bent 90 over a cylindrical die 4. The strip was pulled by a hydraulic cylinder 6 which had a clamping system 5 that held the strip. After the clamping system, the strip was transferred to a cutting machine 8, which cut the strip into small pieces. 21

40 2.5.3 Bending-under-tension test Alinger and Van Tyne [31] evaluated five die materials during repeated stretch-bend sheet steel deformation using the bending-under-tension test with each of three sheet steel surfaces. Figure 2.10 shows the schematic of the bending-under-tension test. Approximately 140 tests have been performed on a fresh surface of each die using each sheet material. The dies, made from a number of alternative materials, are 25.3 mm diameter cylinders, with 360 of testing surface. It was concluded that the tungsten carbide die material performed the best in the wear study. Figure 2.10 Schematic of bending-under-tension test [31] Deep-drawing process-simulator Boher et al [32] developed an experimental device, named the deep-drawing process-simulator (DDPS), to study the tribological interaction between the metal 22

41 strip and the tool in the radius portion of a die in deep drawing (Figure 2.11). A steel strip, unrolled directly from a coil, was in contact with a portion of the radius tool. The flat blankholder and the die radius constituted the working system of DDPS. A rolling up engine pulled the strip through the working system. The loading of the die radius was a result of the restraining forces H and the pulling forces T. The blank holder forces were controlled by a hydraulic cylinder. The sliding of the steel strip over the die radius varied in accordance with a defined angle α which simulated the running of the strip steel on the tool. The strip exit angle α was fixed in relation to the angular position of the reversing cylinder. A low-carbon steel sheet and an X160CrMoV12 steel die radius were used in the experiments. Figure 2.11 Schematic of deep-drawing process-simulator [32] Two mechanisms of surface degradation were determined on the die radius portion through micrographs: adhesion and ploughing. It was found that the tool wear on the die radius was localised in two areas but varied in intensity depending on the exit angle between the sheet and the die radius, which was in accordance with the high contact pressure areas obtained from the numerical simulation. For a strip exit angle of 70 and 80, the main damage at the surface of the die radius was adhesion, while for a strip exit angle of 90, ploughing dominated the main 23

42 damage. The degradation evolution reveals that the adhesion occurred after the first cycle and ploughing was observed after 500 or 700 cycles Slider-on-flat-surface tribometer Gaard et al [33] designed a slider-on-flat-surface (SOFS) tribometer (See Figure 2.12) to investigate the tool wear mechanism in sheet metal forming. In the test, a tool was pushed against a sheet material placed on a solid table with a normal load, applied with a servo engine and slid with a velocity v in the y-direction. A double-curved tool geometry with radii of 5 and 25 mm was utilised. At the end of a track, the tool was lifted and returned to the starting position and moved a selected distance in the x-direction, after which the process was reiterated. During testing, the normal and friction force was measured with a sampling frequency of 1 khz using two separate force transducers, A and B, respectively. Transducer B, used for monitoring the friction force, was mounted as close as possible to the sheet surface to minimise torque due to friction. To indicate the presence of wear, the coefficient of friction was monitored and continuously plotted during the experiments, along with the observation of the sheets. The worn surface morphologies and mechanisms of a deep drawing die were compared with worn surfaces obtained by the SOFS tribometer. It identified abrasion and adhesion as the main surface damage mechanisms on the investigated dies. Transfer of sheet material and abrasive scratching were found as the main damage mechanism of the dies. 24

43 Figure 2.12 Schematic presentation of the SOFS tribometer [33] Twist compression test Kim et al [34] utilised a twist compression test (TCT) to investigate galling, a form of adhesive wear, in forming galvanised advanced high strength steel (AHSS) in automotive stamping. Figure 2.13 shows the schematic of TCT. In the TCT, a rotating annular tool was pressed against a fixed sheet metal specimen while the pressure and torque are measured. To determine the effect of interface temperature upon lubricant effectiveness and galling, the temperature near the tool-workpiece interface was measured as shown in Figure A dummy sheet of 1 mm thickness with a slot for the thermocouple was used. Thus, the temperature was measured at the bottom surface of the sheet specimen used in the test. The specimen and the dummy sheet were held in position with two fixture wings. 25

44 Figure 2.13 Schematic of twist compression test [34] Figure 2.14 Temperature measurement using a thermocouple [34] 26

45 2.5.7 U-Bending test Sato and Besshi [35] carried out a U-bending test is carried out for the evaluation of anti-galling performance of the tools in aluminium sheet forming (Figure 2.15). Bending tests were carried out with a high-speed hydraulic press, the working speed used in the test being mainly 10 mm s -1, but for comparison, a high speed of 100 mm s -1 was used also. Lubricant was applied to the surface of sheet by brushing. All tools were cleaned with acetone before each series of tests. Figure 2.15 Schematic view of U-bending test [35] Nilsson, Gabrielson and Ståhl [36] also utilised U-bending test to evaluate the wear resistance for three different zinc-alloys with different primary phase as die-tool material in forming process equipment. Wear tests were conducted in an excenter press, which was equipped with a die-holder for the die-tools (Figure 2.16). The die-holder was equipped with a monitoring system that allows measurements during the forming process. Measurements during pressing 27

46 operation were performed regarding press force and drawing height. Outside the excenter press, measurements for every 1000 strokes were performed on the loss of weight, surface roughness and radii alteration of the die-tools. The principle for the U-bending process is shown in Figure Two different sheet-metal materials, aluminium AA6016-T4 and steel 220RP, with different wear characteristics have been investigated. Figure 2.16 U-bending equipment showing die-holder with inserts [36] Figure 2.17 Principle for U-bending test [36] 28

47 2.5.8 Strip-drawing test Jonasson et al [37] conducted a strip-drawing test to study shotblasted and electrical-discharge-textured rolls with regard to frictional behavior of the rolled steel sheet surfaces. In this test, originally developed by Wojtowicz [38], a steel strip is pulled between a pair of flat tools while a normal force is applied. In the strip-drawing test all deformation occured in the asperities by replacing one of the tools with a cylinder. A lower contact force gave a larger spread on friction levels. Figure 2.18 strip-drawing test [37] Hortig and Schmoeckel [39] also performed a strip-drawing test to analyse of local loads on the draw die profile with regard to wear (Figure 2.19). The intermitting strip-drawing test with bending was a wear-test, modelling the loads in the flange-region of a deep-drawing die. A sheet metal strip was drawn through a model-tool consisting of blank holder and draw die. During the test the blank holder force was kept on a constant value and the friction force on the blank holder and the total drawing force are measured continuously. In addition to these 29

48 global measurements, the local wear marks on the tool surface were examined in long-time tests to check the plausibility of the calculations. In the experiments with steel sheet material, TiC/TiN coatings on steel were used, because the coating shown visible change of colour by means of tribo-oxidation according to local tribological load. Experiments with aluminium sheet material showed significant influence of the local tribological load on local galling. For localisation of highly loaded areas, a minimal lubrication was used in the tests with aluminium. The WCC coating used for the experiments showed beginning contamination with aluminium in the highly loaded regions. Figure 2.19 strip-drawing test [39] Draw bead test Sanchez [40] carried out draw bead test to measure friction on sheet metal under plane strain. The test method follows Nine s original work in draw bead simulation (DBS) [41]. The sheet metal is pulled to flow between three cylindrical pins of equal radii (Figure 2.20). To determine a coefficient of friction, two test 30

49 specimens are required as a minimum. One specimen is pulled between cylindrical pins supported by ball or roller bearings. Friction on the bearings is considered small enough to be neglected. The pulling (FR) and clamping (FCR) forces measure the bending and unbending resistance of the sheet under frictionless conditions. A second specimen is pulled between pins of radii equal to the rollers, but firmly secured to the tools. Friction opposes the sliding of the sheet over the fixed tools. The pulling (FP) and clamping (FC) forces measure the combined loads required to slide, and to bend and unbend the sheet as it flows over the fixed pins. Figure 2.20 Draw bead test [40] 31

50 Slider test system Cora, Namiki and Koc [42] developed a slider test system to assess Wear performance of alternative stamping die materials. This test system is based on the use of precise and controlled motion of a vertical machining centre (HAAS VF-3 CNC) s x-, y- and z-axes and spindle (no rotation).a load sensor was mounted on the spindle through a holder which also houses the die sample of interest. AHSS sheet blanks are laid on the x y table with clamps at four corners as can be seen in Figure CNC was programmed for the precise pressing of die sample and one-way scratching/sweeping on the AHSS sheet blank. Normal force occurring at the die and blank interface was recorded during the tests. Figure 2.21 Slider test system [40, 42] 32

51 Figure 2.22 Die sample dimensions and its actual photo on wear tracks [42] Figure 2.22 shows the die sample dimension and an actual picture with the wear tracks on the sheet blank. Performance evaluation of die samples was based on the following measurements (1) mass loss, (2) surface profile (roughness) and (3) microscopic evaluations Acoustic emission technique Skåre and Krantz [43] monitored wear and frictional behaviour of high strength steel in stamping by acoustic emission (AE) technique. AE from a forming operation contains measurable data from events such as galling, tool wear, lubricant penetration, stick slip, wrinkling, necking in the sheet material and cracking in the tool or the sheet material. The detected AE is directly proportional to the energy (mechanical) consumed between the contacting surfaces and can therefore be used to estimate the forces acting on these surfaces. A change in the tribological parameters, such as materials in contact, the efficiency of lubricants, the roughness of the contacting surfaces, relative velocity between the contacting materials and contact pressure can be monitored by AE technique. Wear tests have 33

52 been made using flat dies and a U-bending tool. The results indicate that the U-bending tool can be used to study wear behaviour and it simulates forming over the linear portion of a stamping tool. AE, punch force and tool temperature are shown to be essential in the evaluation and understanding of the wear process. The result shows that the surface treatment and surface quality of the tool are important for the wear behaviour. These results indicate that it is possible to use uncoated hardened tools provided that a minimum tool surface quality is maintained. These results also show that hot-dip galvanised high strength steel (HSS) wears the tool out less than uncoated HSS. 2.6 Research and Development in Tool Wear Coating Nowadays, several types of commercial film coatings prepared by chemical and physical deposition process are commonly used to increase the tool life and reduce the requirement for high performance lubricant in sheet metal forming process. Sresomroeng et al [44] evaluated the anti-adhesion performance of commercial nitride and DLC films coated on cold work tool steel against HSS in forming operation. The friction coefficient and wear rate of the non-coated ball (SKD11; hardness 60±2 HRC), balls coated with TiN-PVD, TiCN-PVD, AlTiN-PVD, Nitride+CrN and DLC have been evaluated in sliding contact against SPFH 590 (JIS) disk. The scratch and nano-indentation tests were done on each type of coated tools to characterise the adhesive strength between the film and the 34

53 substrate, and the hardness and the elastic modulus, respectively. The anti-adhesion performance of various film-coated tools in metal stamping process was also investigated by performing U-bending experiment. The cold roll carbon steel (JIS: SPCC) was also used to compare a material transfer problem to the case of using HSS (JIS: SPFH590). As the results, for HSS sheet, the adhesion of workpiece material on a non-coated die surface was detected after 49 strokes whereas adhesion could not be found in case of stamping SPCC sheet up to 500 strokes. The TiCN, AlTiN, and Nitride+CrN films showed good anti-adhesion performance when forming HSS, while the TiN and DLC films did not provide the satisfied results. Fox-Rabinovich et al [45] analysed the wear behaviour for cutting tools with nitride PVD coatings. The chemical and phase composition as well as the structural characteristics of TiN-based PVD versus the nitrogen pressure used during deposition coatings were analysed using AES and XRD methods. Also the friction and wear properties of the coatings were established under different wear conditions. Using these results a relation between the TiN PVD coating s wear resistance and its ability to dissipate the energy of plastic deformation as well as to accumulate the energy of elastic deformation were obtained by using a nano-indentation method. Based on this work, a microhardness dissipation parameter (MDP) was developed to serve as an indicator of a coating s durability. Straffelini, Bizzotto and Zanon [46] improved the wear resistance of tools for stamping using coating by physical vapour deposition with a AlCrN layer. In the first stage of the investigation, the progression of tool wear during a precision stamping operation was investigated. Punches and dies wear made by a heat-treated HSS and each operation took place in a boundary lubrication 35

54 condition. Observed wear was due to adhesion (with some transfer) and after 160,000 strokes micro cracking damage was also shown to start in the punch. A commercial AlCrN (Alcrona) coating was thus selected as the PVD AlCrN coating was reported to give optimal behaviour for a variety of tools [47-49]. The coating was deposited on the S390 HSS tools in the mirror polished condition. The results show that the AlCrN coating gave rise to a significant increase in the wear resistance. Wang et al [50] investigated material transfer phenomena and failure mechanisms of a nanostructured Cr-Al-N coating in laboratory wear tests and an industrial punch tool application. CrAlN and TiN coatings were deposited on AISI M2 tool steel substrate test coupons and on industrial punch tools by electron beam plasma-assisted physical vapour deposition (EB-PAPVD). The microstructure and morphology of the coatings were investigated by XRD, XPS, TEM, and SEM with EDX. Pin-on-disc tribotests were conducted on the coatings against AISI steel counterface material in order to investigate their wear performance, with particular emphasis on the material transfer phenomena during the sliding tests. After industrial trials on piercing high strength steels, the worn uncoated as well as CrAlN- and TiN-coated punches were also studied. The results showed that the nanostructured CrAlN coating exhibited less material transfer and thus better adhesive wear protection than the TiN coating under both laboratory pin-on-disc tribotests and industrial trial conditions. It was also found that the coating morphologies replicated the surface finish of the punch substrates, and that local coating spallation appeared to be initiated at machining grooves on the punches, which were detrimental to the coating lifetime. Aizawa, Iwamura & Itoh [51] explored the effect of a number of layers and 36

55 bi-layer thickness on the mechanical properties by the nano-indentation technique. Nano-lamination is a new way to make full use of multi-layered structure for coating instead of the monolayered coating system. Different from the conventional nano-lamination approach, where two different kinds of material system are deposited in layers, the amorphous carbon layer, a-c:h, is alternatively deposited with graphite-like cluster layer, resulting in an amorphous carbon base nano-laminated coating. Higher hardness and Young's modulus are attained with reduction of bi-layer thickness. The scratching test of this nano-laminated coating is made to demonstrate that it has sufficient scratch load above 100 N. Furthermore, a dry micro stamping test is performed to prove that this nano-laminated coating has sufficient wear-toughness to make dry stamping 10,000 times in practice even if it has nearly the same Young's modulus and hardness as the mono-layered coating. No delamination or break-away occurs on the ironed surface of coated tools while severe delamination is observed in the conventional mono-layered coating Silva, Dias and Cavaleiro [52] assessed the tribological behaviour of W-Ti-(N) thin films by pin-on-disk testing with contact geometry of uncoated and coated 100Cr6 balls sliding against uncoated different disk materials used as stamping sheet. Different types and amounts of lubricants were used in the tests. In non-lubricated tests, friction coefficients as high as 0.8 were achieved. For the more ductile sheet materials (Al alloy and Zn-coated steel) strong adhesion was observed. The best compromise between low wear rate and low friction coefficient was achieved for N-containing coatings deposited without ion gun assistance. In lubricated conditions, a significant decrease of the friction coefficient down to 0.05 and a reduction of the wear coefficient in more than one order of magnitude down to < m 2 N 1 were reached in relation to non-lubricated tests. Very good tribological results were achieved using the 37

56 corrosion protection oil as lubricant, with amounts usually applied for protection of sheet materials (2 g/m 2 ). It was found that the wear coefficient of the coated ball decreased linearly with increasing hardness of the coating, being the best that deposited with N contents in the range from 35% to 40%. The tribological performance of the coated samples was approximately constant even when the amount of used lubricant was reduced to only 25% of the initial value (0.5 g/m2). Schramm et al [53] presented the tribological properties and dry machining characteristics of PVD-coated carbide inserts. The mechanical properties and the dry machining characteristics show that chromium-based cutting tools might have sufficient potential to become a machining alternative to the state-of-the-art TiAlN coating. It could be shown that the deposited CrxN and CrxAlyN coatings have a poor machining performance, which could be explained by the brittle coating structure and/or high coefficient of friction. The high hardness of both CrN and CrAlN could not yet be completely utilized for dry machining, which can be seen in the increased abrasive wear. In contrast, the good surface quality during machining of SGI-50 and 42CrMo4 are encouraging for further investigations. It is possible to improve the coating systems by changing, pre- and post-treatment of the cemented carbide tools [54-56]. Van der Heide et al [57] conducted the wear resistance of alternative tooling materials by a combination of forming tests at a high speed stamping line and model wear tests using the TNO slider-on-sheet tribometer. With this tribometer, volume loss of alternative tooling materials can be determined as a function of the sliding distance, using sheet materials from automotive practise. Results show that the wear rate of a soft tool material can change two orders of magnitude as a result of the zinc layer type used. Furthermore, it is shown that the relative performance 38

57 of alternative tool materials is strongly related to the hardness of the (tooling and sheet) materials. Industrial forming tests with a selection of alternative tooling materials confirmed the model wear test results. The same ranking of the tooling materials with respect to volume loss is obtained per sheet material. Bressan et al [58] concluded Wear on tool steel AISI M2, D6 and coated with Al 2 O 3 by the MOCVD process. The wear tests by sliding and abrasion were performed in a pin-on-disk and ball-on-disk apparatus, whose pin and ball substrates were steels fabricated from AISI M2, D6 and From the plotted graphs of lost volume versus sliding distance, it was observed a greater wear rate of AISI D6 pins without coating, and this is possibly due to more severe adhesion and delamination mechanisms. The AISI M2 and D6 pin coated with Al 2 O 3 showed similar wear resistance and higher resistance than the uncoated D6 pin. However, the tested sphere of AISI showed different behaviour under 20N normal load. For both the spheres, coated with Al 2 O 3 and uncoated, the wear rate was similar. Nitrided M2 and D6 tool steels coated with Al 2 O 3 showed superior wear resistance characteristics for cold working tooling. The spheres of AISI coated with Al 2 O 3 presented poor wear resistance due to surface defects. Cora and Koc [59] investigated the wear performances of seven different, uncoated die materials (AISI D2, Vanadis 4, Vancron 40, K340 ISODUR, Caldie, Carmo, 0050A) using a robot-based die wear test system (Figure 2.23). DP600 AHSS (advanced high-strength steel) sheets were used in these tests. For the same force levels, similar wear scars and depths were observed for all tests except for 0050A and K340 Isodur. In some part of the K340 Isodur tests, depth of wear tracks on sheet blank was shallower and the sheet surface was shiny. It is concluded that this material is more prone to material stacking on the surface and 39

58 coating might be necessary for some cases. For the Vancron 40 specimen, the wear pattern was almost uniform along the contact surface. Figure 2.23 Robot-based die wear test system [59] Lubrication Lubrication plays a critical role in sheet metal stamping process as it reduces friction between the tool and blank and enhances the ability to produce a good quality part. The lubrication fluid acts as a barrier to separate the tool surface from the sheet material and then decreases the interface strength between the contacting surface asperities [60]. It is important to understand the influence of the lubrication on the tool wear distribution in sheet metal stamping, especially in forming complicated automotive parts using AHSS. There has been extensive research carried out to evaluate the influences of lubrication behaviour in sheet metal processing of various materials. Kim et al [61] presented a practical methodology that uses the deep drawing test and finite element (FE) analysis to evaluate stamping lubricants under near 40

59 production conditions. In this study, five stamping lubricants (four dry film lubes and one wet lube) were evaluated using the deep drawing test. The performance of the lubricants were evaluated based on: (a) maximum punch force measured, (b) the maximum applicable blank holder force (BHF), (c) the draw-in length, (d) the perimeter of flange after test, (e) the change of surface roughness, and (f) the inspection of surface topography. The coefficient of friction for each lubricant tested was determined through the FE-based inverse analysis by matching the predicted and measured values of the load-stroke curve and the draw-in length. This study showed that one of the tested lubricants was most effective, regardless of test speed and the magnitude of BHF. The methodology used was shown to be effective in evaluating various lubricants for sheet metal forming and accurately differentiating their performances. Chandrasekharan et al [62] developed a laboratory ironing tribo test to evaluate stamping lubricants at various temperature levels (Figure 2.24). Lubricants were evaluated and ranked based on (a) ironing load, (b) surface quality of the ironed cup and (c) apparent shear friction factor. Five lubricants, namely a dry film, a zinc phosphate coating+sodium soap, a pre-emulsified with solid lubricant and two emulsions were tested using the ironing test at both room temperature and elevated temperature (100 C) conditions. It is concluded that at both room temperature and elevated temperature, Lub B (the zinc phosphate coatingcsodium soap), performed best followed by Lub A (the dry film) and Lub E (pre-emulsified with solid lubricant), while the emulsions failed (scratching and galling) due to the high interface pressures. However, Lub A and Lub B are ideal lubricants for sheet metal forming operations that generate contact pressures in the range of 650 MPa and interface temperatures in the range of C; however, they cannot realistically be used in a high speed progressive die or transfer die operation because they are costly to apply and remove. Lub E is a pre-emulsified lubricant 41

60 with solid lubricant and cannot be sprayed. Thus, it requires special application equipment similar to brushing at each stage in stamping operation. Figure 2.24 Simulation testing machine for hot stamping [62] Yanagida and Azushima [63] discovered that the obtained coefficients of friction under lubricated conditions for steels were lower than those under dry conditions in hot stamping. The coefficient of friction in hot stamping was measured using a tribosimulator. Simulative experiments were carried out using SPHC steel and 22MnB5 steel under dry conditions. The coefficient of friction of 22MnB5 steel was higher than that of SPHC steel. It was shown that the use of lubricants was effective for decreasing the stamping load and die wear in hot stamping. Deshmukh et al [64] carried an extended duration pin-on-disk experiments to determine the relative performance of a wide range of lubricant combinations in a commercial brake valve assembly. In the experiments, the lubricants were initially applied to the disk surface but were not replenished over a sliding distance of more than 6000 m. The experimental results revealed that the environmentally friendly lubricant, boric acid, was highly ineffective for reducing the wear in the surfaces tested. When combined with a commercial transmission fluid, however, the boric acid mixture proved to be highly effective in terms of both friction and 42

61 wear performance. Based on the success of the combined lubricant experiments, the boric acid was then mixed with canola oil to form a completely natural lubricant combination. Based on further pin-on-disk experiments, this lubricant combination yielded the best wear performance of all the lubricants tested Alternative die materials Recently, automotive industry shifts its focus on customised production, facing an increasing demand for medium and small batch production, where cost-effective manufacturability of sheet metal forming dies with improved tool wear behaviours comes into the foreground. Some alternative materials, such as filled polymers, offer possibilities to fulfil such requirements. Work has also been carried out to prolong tool-life through utilisation of alternative die materials. Rück, Boos and Brown [65] conducted an investigation of the effect of metal ion implantation into high speed steel dies, using high current metal ion beams from a repetitively pulsed vacuum arc ion source. The testing method used was the upsetting process, which is comparable to actual forming processes and stimulates the wear strain of the tools used in the metal forming industry. Narojczyk, Werner and Piekoszewski [66] utilised nitrogen ion implantation for stamping die to form the cross-recessed heads of screws. It was revealed that the effect of nitrogen ion implantation on the wear rate of stamping dies resulted in an improvement by a factor of about 2.5. And tool chipping rate was reduced by a similar factor as well. However, no effect of ion implantation upon the force exerted by the tool on the workpiece was found. 43

62 de Souza and Liewald [67] investigated the tribological and tool design aspects of using polymeric materials for sheet metal forming purposes. In the study, wear behaviour of two polymer composites on sheet metal counterface materials have been investigated. A new testing method for wear evaluation of polymeric materials for sheet metal forming using a Strip Drawing Test facility was developed as shown in Figure A method to predict lifetime of polymeric stamping dies using the linear wear-distance relation measured with the new testing method was also proposed. Significant improvements in wear performance of polymer composites have been observed using sheet materials with structured surfaces. Figure 2.25 Scheme of strip drawing test [67] Myint et al [68] compared the tool wear mechanism of tetragonal zirconia polycrystal (TZP) punch with that of commercially available WC (tungsten carbide) punch during stamping. The tool life for the TZP punch was found to be over 2.5 times higher than that of commercial tungsten carbide. The worn-out tools were analysed using scanning electron microscope and optical microscope for studying the tool wear mechanisms. Tool wear and chemical action possibly 44

63 cause the failure of the tungsten carbide punch, whereas wear of TZP punch is predominantly caused by mechanical shearing of asperity and plastic deformation. Due to their inherent high melting point and the absence of the second-phase binder, ceramics materials do not soften at higher temperature unlike the carbide tools. Hence, they can be used at high cutting speeds without initiating deformation/diffusion wear. This assists in improving the tool life significantly. In addition, TZP ceramics is inert, corrosion resistant and non-wetting when contacting metals. Exposed carbide grains act as a site for increased wear and metal pickup during precision, high-speed metal stamping and forming. Moreover, cobalt-depleted carbide tools can create burring of the strip being stamped, leading to poor part quality. A few rapid tooling technologies have been recently proposed and among them Selective Laser Sintering is probably one of the most relevant and promising. Levy et al [69] reported some results of a wide experimental research on the application of SLS tools in sheet metal forming. A wear test was carried out to investigate the progressive degradation of laser-sintered materials in comparison with traditional cold-work steels. In conclusion, SLS may represent an effective rapid tooling technique in the field of tool manufacturing for sheet metal stamping. Pinto et al [70] studied the usability and robustness of polymer and wood materials for tooling in sheet metal forming. A target production volume has been defined and both tool materials were submitted to stampings in the press shop and the evolution of tool wear, roughness and geometrical changes in punch and die radius were measured throughout production. In spite of that the tooling costs of presented alternative materials are very similar, results have shown that a good 45

64 compromise for this particular presented part should be the use of the polyurethane based material once this material does not suffer an excessive wear like the densified wood material, and therefore, the stamped part accuracy is preserved. Nevertheless, according to the experimental results, both materials have shown that they can be a practical alternative in the production of tools for sheet metal forming, both by their aptitude and robustness as well as their economic feasibility, in the low volume production series Tool wear modelling Although a number of test methods have been developed in recent years, some numerical tool wear models for sheet metal forming process were also introduced. Ersoy-Nürnberg et al [71] have studied the simulation of tool wear in sheet metal forming tools using the modified Archard s model in which wear coefficient is a function of accumulated wear work and is proportional to the dissipated energy. In order to determine these wear coefficient values as well as their gradients along the life cycle, deep drawing experiments with a cylindrical cup geometry were carried out. The prediction of tool wear is accomplished by REDSY, a wear simulation software developed at the Institute of Metal Forming and Casting, TU München. The wear predictions made by this software are based on the results of a finite element deep drawing simulation. The results obtained using the proposed model are in a good agreement with the experiments. Hambli [72] has developed a wear model in sheet metal blanking/punching process using finite element analysis with tool wear as a function of normal 46

65 pressure and material properties. A wear model has been implemented in a finite element code, in which the tool wear is a function of the normal pressure and some material parameters. A damage model is used in order to describe crack initiation and propagation into the sheet. The distribution of the tool wear on the tool profile is obtained and compared to industrial observations. In general, the need for regrinding of the shearing tool is determined on the basis of allowable burr height on the final product. This wear analysis is very helpful to improve the reliability of the shearing tool and to determine the tool repair or change Die radius geometry It has also been observed in some studies that tool geometry also plays an important role in affecting the tool life of sheet metal stamping of conventional steels. Boher et al [32] studied the tool wear behaviour by investigating the degradations of the radius portion of a die in deep drawing process of low-carbon steel sheets. It is concluded that strip particle transfer is the main wear damage and it is located on two specific areas of the die radius. The tribological behaviour of the die radius is quite different in function of the strip exit angle. For low strip exit angles, particle transfer on the die radius is important and for high strip exit angle, the main damage is abrasion. The friction coefficient may also give information about the contact evolution. Hortig and Schmoeckel [39] have described an approach to identify the characteristic distribution of local loads on the draw die surface and have analysed 47

66 the influence of various parameters such as sheet thickness, draw die radius, coefficient of friction, and material parameters on load amount and positions. Hoffmann, Hwang and Ersoy [73] developed an advanced simulation scheme for tool wear modelling that considers the geometry changes caused by wear using interactive iterations. REDSY is tool wear simulation software developed by Hoffmann et al, which implemented Geometry-Update-Scheme (UGS) to consider the change of tool geometry by the increase of the number of punch strokes. Figure 2.26 describes an iterative algorithm applied in UGS. REDSY imports the results of forming simulation, calculates elemental wear, and exports the worn grometry to a file for the next iterative. Figure 2.26 Algorithm applied in UGS [73] Jensen et al [19] applied a finite element method to determine the distribution of tool wear in deep drawing of austenitic stainless steel using a quantitative semi-empirical wear model and compared it to industrial observations. In the tool wear prediction model, it was assumed that the equations for the adhesive wear and abrasive wear were identical, and the constants and the hardness of the die 48

67 remained constant through the operation time. The simulation results showed that the tool wear was concentrated in two areas at around 20 and 70, which agrees with the experimental results obtained at Grundfos A/S and the Technical University of Denmark [74]. Through the simulations, it was observed that an increase in the n-value led to a significant reduction of the tool wear. The blank thickness, the ratio of the blank thickness to the die radius and draw ratio, resulted in a large increase of tool wear when these parameters were increased. However, it was found that tool wear does not depend significantly on the die radius. More recently Pereira et al [75] have provided a qualitative description of the evolution and distribution of contact pressure at the die radius in sheet metal stamping process and have identified three distinct phases of contact pressure distribution to better understand the wear phenomenon. It was found that a typical-peak steady-state contact pressure response existing for the majority of the process proceeded by a transient response. It was revealed that the peak transient contact pressure was more than double the steady-state peak, which may have a significant influence on the tool wear distributions. In a numerical study on the circular cup drawing test, Shahani and Salehinia [5] have used finite element method to study the wear depth on draw die arc segment and revealed that the contact stress peaks can be reduced by simply increasing the die radius. The wear model was developed by considering the abrasive wear only. It was concluded that the wear profile contains two peaks: one near the inlet of die arc and the second at some distance from the outlet. The second relative maximum wear point moved toward the end point (90 ) of the draw die arc by decreasing the clearance between the punch and the die. The influence of the blank holder force on the first peak of the relative wear depth was much more than 49

68 that on the second peak. For blank holding force, there was a certain value before which increasing its magnitude increased the peak values of the relative wear profile and beyond that the peak values decreased as the blank holder force increased. In opposition to the results obtained by Jensen et al [19], the results of the simulations showed that increasing the radius of the die causes the relative wear depth to be decreased and to be more distributed. 2.7 Summary Rising fuel prices and the increasing customer demand for safety have led to the greater usage of new AHSS in the automobile industry. Compared to conventional mild steels, AHSS show higher strength levels as well as improved hardening characteristics, which makes them suitable for applications where low weight and improved passenger safety are major design targets. During sheet metal stamping, however, the higher surface hardness and high material strength of AHSS lead to higher contact stresses between the tooling and the work piece, which results in increased tool wear compared to conventional steel grades [18, 76]. There has been extensive research carried out to study and predict the tool wear behaviour in sheet metal processing of various materials both numerically and experimentally. As many control parameters can affect the severity of tool wear, research work has also been carried out to prolong tool-life through a combination of surface coatings and alternative die materials. 50

69 In predicting tool wear, it has also been observed in various studies that tool geometry also plays an important role in affecting the tool life of sheet metal stamping of conventional steels. However, very little work seems to have been done on determination of exact die geometry to reduce tool wear in sheet metal stamping of AHSS. Even though some of the previous studies have shown that tool wear can be reduced by modifying the die shape, they mainly focused on dies with standard circular profiles and the forming of a conventional steel sheet. Automotive industries are now adopting increasing use of AHSS for their body panels due to their superior strength, light weight and crashworthiness capabilities. However, as AHSS can show strength levels three to four times higher than conventional steel sheet, these previous studies on tool wear do not give an accurate indication of the die contact stress distribution and the effect of the die shape on tool wear when forming AHSS. 51

70 CHAPTER 3 TOOL WEAR PREDICTION USING AUTOFORM SOFTWARE 3.1 Introduction During sheet metal forming, various process control parameters usually display their effect leading to a degree of uncertainty in the tool wear severity. To identify the complex tool wear problems in stamping as early as possible, simulation software is used to study the forming process during tool development. As the rapid growth in the research and development of finite element simulation for sheet metal forming application continues, a number of commercial software is available in the markets, for example, AutoForm, Pam-stamp, dynaform, etc. In this chapter, AutoForm 4.1 software was employed as pre-processing, simulation and post-processing tool to conduct a finite element analysis of the tool wear prediction for a particular stamping part. For the simulation study, an assembly tool and a part model for a reinforced rear suspension support of a GM Holden s vehicle was created in the Unigraphics NX2 software. The model consists of a die, punch, binder and part. The model was exported to AutoForm software. After defining material properties of each part in the model, the model was automatically meshed to produce nodes and elements by selecting corresponding parameters in AutoForm, followed by determining the parameters of boundary conditions and contact characteristics between each 52

71 component and loads. These procedures are referred to as the pre-processing procedures. After specifying parameters of the solution, AutoForm software would solve the specified problem. 3.2 AutoForm software AutoForm software package version 4.1 was developed by AutoForm Engineering GmbH [77]. The software offers solutions for the die-making and sheet metal forming industries. The software can improve reliability in planning, reduce the number of die tryouts and tryout time, and lead to higher quality part and tool designs that can be produced with maximum confidence. In addition, press downtime and reject rates in production are substantially reduced. Based on practical, industrial know-how and sheet metal forming expertise, AutoForm s solutions result in a complete, integrated and specialised system to analyse, review and optimise every phase of the process chain. AutoForm provides solutions all along the sheet metal forming process chain (Figure 3.1). It ranges from stand-alone modules for small and mid-size companies to complete, integrated multi-module systems for large companies. 53

72 Figure 3.1 Sheet metal forming process chain in AutoForm Software [78] The software provides accurate simulations for sheet metal forming based on the static implicit approach, which can be expressed as: T u dv t u da (3.1) v ij i, j A i i where V is the volume, A is the surface area, T ij is the Cauchy stress tensor, u i,j is the gradient of the displacements, t i is the traction vector and δ is the variational operator [79]. In sheet metal forming, for a certain part with a fixed drawing depth, the contact pressure distribution of the work-piece provides a reference to predict the tool wear of the die. The contact pressure shows the normal stress imposed on a 54

73 work-piece by the action of the die and punch. By examining the reaction stresses of a die, it can be used to assess the danger of the tool wear during the forming process. AutoForm incremental module produces the contact pressure distributions of a work-piece at the die-workpiece interface to indicate the wear of the corresponding die, under various binder pressure loads and lubrication coefficients [77]. AutoForm die advisor module was used for the prediction of the tool wear location and the extent of wear, and determination of the optimal coating method of the tool. Various coating methods, such as physical vapour deposition (PVD), chemical vapour deposition (CVD) and protective coatings, including TiN, TiCN, TiAlN, hard-chrome and a-c:h, are supported by the module. This module utilised the finite element model to calculate friction work generated at contact regions between the die and workpiece [77]. Friction work is the work of friction per unit area and can be expressed as the integral of frictional shear stresses over an element: A F Fds (3.2) where A F is friction work, τ F is the frictional shear stress at the nodes in an element and s is the sliding distance. Wear volume w can be expressed as: k w AF (3.3) H 55

74 where H is the hardness of tool material and wear coefficient k was measured by experiments performed by VST Keller, a partner of AutoForm Engineering GmbH. 3.3 Simulation Setup A reinforced rear suspension support of a vehicle was used as a case study (Figure 3.2). The material of the part is cold rolled hot dip galvanized high strength steel. The production rate is 8 strokes per minute and production volume is 100,000. The thickness of the part is 2.5 mm. Table 3.1 shows the material properties of the part. Figure 3.3 illustrates the forming limit curve (FLC) of the part obtained from a test from the material supplier, in which the minor principal strain is along the x axis and the major principal strain is along the y axis. The FLC is used in sheet metal forming for predicting forming behaviour of sheet metal [1, 80]. The diagram attempts to provide a graphical description of material failure tests, such as a punched dome test. Table 3.1 Material properties of reinforce rear suspension support Young's Specific Strain Initial yield Strength Poisson's module weight hardening stress coefficient ratio coefficient (MPa) (Nm -3 ) (MPa) (MPa) Normal anisotropy

75 Figure 3.2 Reinforced rear suspension support Figure 3.3 Forming limit curve Figure 3.4 shows the sequences of simulation steps used in AutoForm. The CAD data of finished part was imported to AutoForm incremental module and then meshed automatically. The blank, binder, punch and die were then imported to the module by AutoForm-UG interface, respectively, and placed at their specified 57

76 locations according to the information obtained from the plant-site (Figure 3.5). Process parameters, including lubrication coefficient and binder pressure load, as well as material parameters were defined in the incremental module. Parameters concerning the die, including tool surface protection method, production volume and production rate were then set in the die advisor module. Initial simulation was performed to find critical tool worn areas of the die. Simulations were then run for varying lubrication coefficients, binder pressure loads and tool surface protection methods to determine the influences of lubrication coefficients and binder pressure loads on the contact pressure distribution of the workpiece and the influence of coating method on the tool wear distribution of die in the critical tool worn area. The contact pressure distributions of the workpiece and tool wear distributions were obtained through the incremental module and die advisor module, respectively. CAD data of finished part Incremental module Die advisor module Tool wear and contact pressure distribution Blank Material Tool surface protection method Production volume Die/punch and process condition Production rate Figure 3.4 Simulation sequences in AutoForm 58

77 Binder Punch Die Blank Figure 3.5 Blank, binder, punch and die in AutoForm software 3.4 Results and Discussion Identification of critical tool worn areas In the initial simulation used to determine locations of the critical tool worn area, the binder pressure load and lubrication coefficient were set as 4.5 MPa and 0.15, respectively. The initial coating method was selected as PVD Steel. Figure 3.6 plots the tool worn areas distribution obtained from the initial simulation. In Figure 3.6, the area with the red colour presents an area of sensitivity to tool wear, and it is a tool worn failure area with high probability. However, the area with the green colour in Figure 3.6 means an area of insensitivity to tool wear, and it is unlikely the tool wear would occur in the area. The area with the yellow colour is marginal area to suffer tool wear. It was concluded that Areas 1, 2, 3 and 4 were highly sensitive to tool wear, and tool wear would occur in the very early stage of 59

78 the production. These areas were compared with the worn out areas of the actual surfaces of the parts produced. Area 1 Area 2 Area 4 Area 3 Safe Marginal Failure Figure 3.6 Potential tool worn area location on die surface obtained from initial simulation Figure 3.7 illustrates the photos of the worn areas, named Areas 1, 2, 3 and 4, located on the surface of the actual sheet metal part, which contacted to the corresponding Areas 1, 2, 3 and 4, respectively, in the sliding movement during the sheet metal forming process. The initial predicted result is found to be in accordance with the result obtained from the plant-site. Areas 1, 2, 3, 4 were identified as the critical tool worn area on the die surface. As the gradient in both Areas 2 and 3 was extremely large in both longitudinal and latitudinal directions, this resulted in highly increased sliding movement between the tool surface and the part surface, and accelerated the formation of worn areas in both the die surface and sheet metal blank surface. In the following discussion of contact pressure distributions, a cross-section of Areas 2 and 3, named 60

79 Cross-section 1, was selected as a sample for further investigation (See Figure 3.8). Area 1 Area 2 Area 3 Area 4 Figure 3.7 Photos of the worn areas on the corresponding surface of actual sheet metal part X Cross-section 1 Figure 3.8 Cross-section 1 on Areas 2 and 3 61

80 3.4.2 Relationship between tool worn area, contact pressure and drawing depth Figure 3.9 Contact pressure distribution and drawing depth at Cross section 1 To study the relationship between tool worn area, contact pressure and drawing depth, Areas 2 and 3 are selected as a sample. The worn area in the tool surface is corresponding to the surfaces of sheet metal part surface with high contact pressure. Figure 3.9 shows the contact pressure distribution and the drawing depth at Cross section 1 on Areas 2 and 3 of the sheet metal part, which are the corresponding areas of Areas 2 and 3. The result shows the drawing depth changes rapidly in these areas. The gradient in these areas is extremely large in both longitudinal and latitudinal directions, which results in highly increased 62

81 sliding movement between the tool surface and the part surface, and accelerates the formation of the worn area in both tool surface and sheet metal part surface. From the simulation, it is concluded that the areas that are most sensitive to the tool wear occur at the locations corresponding to the large gradient of drawing depth, and these locations are also the areas with high contact pressure Comparison of contact pressure distribution for various lubrication coefficients Figure 3.10 plots various contact pressure distributions under different lubrication coefficients based on the Coulomb Model. The lubrication coefficient is the dynamic friction coefficient, which indicates that the frictional force is proportional to the normal load. Two positive extrema of the contact pressure along the section increased from 50 MPa to 90 MPa and from 67 MPa to 118 MP as the lubrication coefficient rose from 0.05 to The positive extrema were located at Areas 2 and 3, which validates that Areas 2 and 3 were critical areas sensitive to the tool wear. The variation of negative positive extrema is not as significant as positive ones. From Figure 3.10, it is noticed that the contact pressures at Area 3 decreased abnormally as the lubrication coefficient increased from 0.15 to This was caused by the split of Area 3. Area 3 began gradually splitting while the lubrication coefficient was rising from 0.20, as the dry lubrication condition blocked the smooth movement of the material flow. Considering the formability 63

82 of the workpiece, lower contact pressure and qualified formability could be reached by selecting 0.10 as the lubrication coefficient. Area 3 Area 2 Figure 3.10 Contact pressure distributions upon various lubrication coefficients along Cross-section 1 (Binder pressure load is 4.5 MPa) Comparison of contact pressure distribution upon various binder pressure loads To study the contact pressure distribution under various binder pressures, 3 MPa, 4.5 MPa and 6 MPa binder pressure loads were applied in the simulation, respectively. Figure 3.11 shows that two positive extrema of the contact pressure along the section rose from 55 MPa to 82 MPa and from 88 MPa to 115 MPa as the pressure loads increased from 3 MPa to 6 MPa, which shows again that Areas 64

83 2 and 3 were extremely sensitive to the tool wear. The negative maximum of the contact pressure remained at approximately -90 MPa to show it was not sensitive to the variation of the pressure loads. Under the condition of the same drawing depth, the large contact pressure indicates the increased potential tool wear. However, the lower contact pressure, i.e. lower binder pressure load, results in insufficient stretch of the workpiece. To balance the tool wear and formability of the workpiece, 4.5 MPa was selected as the binder pressure load. Area 3 Area 2 Figure 3.11 Contact pressure distributions upon various binder pressure loads along Cross-section 1 (Lubrication coefficient is 0.10) 65

84 3.4.5 Comparison of tool wear distribution upon various die coating Figure 3.12 illustrates sensitive tool worn areas on the die surface by the colour red using various coating methods. As Areas 2 and 3 were most sensitive to the tool wear, the maximum production volume until the occurrence of local wear along cross-section 1 under various die coating methods is shown in Figure The binder pressure load used is 4.5 MPa and the lubrication coefficient is 0.10 in these simulations. It is noted from Figures 3.12 and 3.13 that PVD steel coating provides the least protection of the die and the local wear would appear in a short time at Areas 2 and 3 when the production volume arrived at 40K. The die surface was found to obtain high-quality protection using a CVD TiC 3D steel coating and the maximum production volume without the wear being increased to 120K. From the results of simulation, it is observed that a CVD TiC 3D steel coating was highly recommended, as it postponed the tool wear appearance to the utmost extent and extended the die-life, which could reduce the frequency of die maintenance. 66

85 (a) PVD steel (b) Advanced PVD steel (c) CVD TiCTiN steel (d) CVD Tic 3D steel Safe Marginal Failure Figure 3.12 Tool wear distributions upon various die coating method with lubrication coefficient 0.10 Area 3 Area 2 Figure 3.13 Maximum production volume until the occurrence of local wear along Cross-section 1 (Binder pressure load is 4.5 MPa and lubrication coefficient is 0.10) 67

86 3.5 Advantages and Limitations of AutoForm Software Based on the case study of the GM Holden s part, it was seen that AutoForm 4.1 software is able to highlight the location and relative severity of tool wear using numerical simulations and analysing the contact pressure on the tool surfaces during forming. Based on the case study, it is possible to use wear prediction tools to locate tool worn areas. In terms of the simulations and tool wear predictions of sheet metal stamping, the AutoForm 4.1 software has the following advantages: (1) The software is specialised in the simulations of sheet metal stamping; (2) Pre-processing of data is fully designed towards to the requirements of the sheet metal stamping; (3) Entire sequence is fully automatic and user-friendly; (4) Creation of elements is automatic and fast; (5) Computation time is reduced and acceptable. However, it is found that the AutoForm software has a number of noticeable limitations. One of them is the problem of accuracy. Figure 3.14 shows the comparison of major strain results from AutoForm 4.1 and Abaqus 6.8 software with experimental results for channel bending test under binder holder forces 12 KN and 36 KN [81]. The comparison done by other team members in this project concluded that the AutoForm software does not predict strains with the same accuracy as the Abaqus model [81]. In general AutoForm under predicts all values. The likely reason for this is that the Abaqus model is significantly more detailed with many small elements to increase the accuracy of the result, whereas the AutoForm software does not support flexible elements creation. However, this 68

87 level of detail is not currently feasible for industrial sheet forming simulations due to the time-constraints in setting up and running the simulation, but will be extremely valuable during the development of tool wear prediction methods in this work. Figure 3.14 Comparison of major strain results from AutoForm and Abaqus software with experimental results [81] Besides the lesser accuracy of AutoForm software, there are a few other limitations of the software: (1) The pre-processing is less flexible and, especially, in the creation of elements. (2) Although the post-processing of the simulation results is suitable for the industrial simulations, it is significantly insufficient for the research work as 69

88 the simulation results cannot be easily retrieved and tabulated for external data processing. 3.6 Summary In this Chapter, the influences of the binder pressure load, lubrication coefficient and coating on the tool wear distribution for a certain sheet metal stamping die were investigated based on numerical simulations using AutoForm software. The areas that were sensitive to the tool wear were identified in the initial simulation, which were found to be in accordance with the phenomena observed from the on-site production of the actual parts. From results obtained from simulations, the lower binder pressure load, improved lubrication coefficient and coating were selected, which could reduce the likelihood of too wear. Results have shown that numerical simulation method using AutoForm can be used effectively in reduction of lead-time in the tool wear prediction for automobile manufacturers. However, due to lesser accuracy and limited support for the post-processing of AutoForm software for the research purpose, Abaqus software is selected as the development tool for the tool wear prediction modelling in the following chapters. 70

89 CHAPTER 4 NUMERICAL TOOL WEAR PREDICTION MODELLING 4.1 Introduction As the commercial specialised software AutoForm has limitations for the tool wear prediction modelling as described in Section 3.5, in order to study tool wear behaviours of a common stamping part, Abaqus 6.8 was used as pre-processing, solution and post-processing tools to establish a numerical tool wear prediction model. Abaqus 6.8 is a software package for the finite element analysis and design developed by SIMULIA. Abaqus software is an extremely popular, class-leading modular suite of FEA software used across a broad spectrum of industries. Its open and flexible simulation solutions provide a common platform for fast, efficient and cost-effective product development, from design concept to final-stage testing and performance validation. It provides the most complete and flexible solution to help researchers and engineers understand the detailed behaviour of a complex assembly, refine concepts for a new design, investigate the behaviour of new materials, and simulate a discrete manufacturing process. The software suite delivers accurate, robust, high-performance solutions for challenging nonlinear problems, large-scale linear dynamics applications, and routine design simulations. Its user programmable features, scripting and GUI customization features allow proven methods to be captured and deployed to an enterprise, enabling more design alternatives to be analysed in less time. 71

90 The deep drawing process for a common stamping part has large deformation characteristics, thus a reliable nonlinear analysis tool is required for the simulation of its drawing process. Abaqus is a pioneer in the discipline of nonlinear analysis. Its nonlinear capabilities have evolved according to emerging analysis needs, maturity of analysis methods and increased computing power. Abaqus includes a full complement of nonlinear elements, material laws ranging from metal to rubber, and the most comprehensive set of solvers available. It can handle even the most complex assemblies especially those involving nonlinear contact and large deformation. In this chapter, pre-processing and solution procedures for establishment of tool wear prediction model are presented in detail, while the post-processing, i.e. results and discussion, will be described in a later chapter. An assembly model for Abaqus according to a typical environment and structure of a deep drawing process was created in Abaqus/CAE module. The model consists of a die, a punch, a binder holder and a part. After defining material properties of each part in the model, the model was meshed to produce nodes and elements for the FEA followed by determined boundary conditions. The contact characteristics between each component need to be defined as well. A series of steps with loading was applied on the punch to drive the deep-drawing process to stretch the part. These procedures are referred to as the pre-processing procedures. After specifying parameters of the solution, Abaqus would solve the specified problem. 72

91 4.2 Wear Work Calculation Along Die Radius Profile In sheet metal stamping, adhesive wear and abrasive wear are two primary types of wear. Adhesive wear is a type of wear due to localised bonding between contacting solid surfaces leading to material transfer between two surfaces or loss from either surface [7]. Holm [22] and Archard [23] concluded that adhesive wear volume V Adhesive is generally proportional to the applied load F and sliding distance s but inversely proportional to the hardness H of the surface being worn away, so that, V Adhesive kfs H (4.1) where k is the non-dimensional wear coefficient dependent on the materials in contact and their cleanliness. Abrasive wear on a die surface is a common phenomenon in sheet metal stamping because the hardness of a die is larger than that of a sheet metal blank. In many cases, the wear mechanism starts with adhesive wear, which generates wear particles that are trapped at the interface, resulting in an abrasive wear [25]. A simplified model for abrasive wear volume V Abrasive was developed by Rabinowicz [26] as Eq. (4.2), 2 tan V Fs Abrasive H (4.2) 73

92 where θ is roughness angle, H is the hardness of the softer surface and F is the normal load. Assuming k and tan are constant during sheet metal stamping, the total wear volume V can be express as Eq. (4.5) V V V (4.3) Adhesive Abrasive V 2tan Fs c H (4.4) V KFs (4.5) To apply Eq. (4.5) in the finite element analysis in the deep-drawing simulation, an integral form of the equation is introduced. The wear work after a stroke with T as the total time of stroke is expressed as Eq. (4.6) and (4.7): W T F vdt t KA 0 A (4.6) W T pvdt (4.7) 0 t t where v is the sliding velocity at the time point t and A is the area. 74

93 The finite element analysis for deep-drawing simulation is an incremental process in Abaqus 6.8. The total time for a stroke is divided into n incremental segments. Eq. (4.7) needs to be discretised in an incremental form, then the wear work W θ on the angle θ of the die radius after a stroke is given by the following equations: n W p v ( t t ) (4.8) i, i, i i 1 i 1 W n p s (4.9) i, i, i 1 Where p t is the normal contact pressure at time point t. In the Abaqus simulations, the contact pressure on each element of the die radius surface is obtained for each incremental segment. As the sliding movement of the blank relative to the die radius surface of each element is not identical, due to the stretch of the blank during the deep-drawing, the distance of the sliding movement of the blank is not assumed to be same in each element. Thus, the distance is also obtained for each incremental segment. 75

94 4.3 Finite Element Modelling Geometry To analyse the contact pressure and wear work distribution along the die radius profile, a finite element model of the deep-drawing was established using Abaqus 6.8/CAE module. With Abaqus/CAE, the geometries of each component of the model and its assembly can be quickly and efficiently created, edited, monitored, diagnosed, and visualised before the execution of Abaqus analyses. Abaqus/CAE supports familiar interactive computer-aided engineering concepts such as feature-based, parametric modelling, interactive and scripted operation, and GUI customisation. In deep-drawing process, the tool (die, punch, blank holder) and blank were symmetric about a plane along the centre of the channel, respectively, in all aspects, including the geometry, materials, loadings and boundary conditions. To take advantage of these symmetric characteristics and to reduce the size, scope and processing time of the model, only a half of the model was created in the analysis (Figure 4.1). A two-dimensional, plane strain model was used under the assumption that there is no strain in the out-of-plane direction of the model. The model consists of four distinct components: die, punch, blank holder and blank. The blank was squeezed between the blank holder and the die. These four components were assembled as a whole model. The dimensions and parameters of the model are listed in Figure 4.2. Although the tool components are much stiffer than the blank, they were modelled 76

95 as deformable materials as the contact pressures and tool wear distributions have to be measured at the surface of die, especially at the surface of die radius. Die 90 Blank y Blank holder Punch 0 x Figure 4.1 Meshed finite element model in the deep-drawing simulations 77

96 Punch diameter, D p Punch radius, r p Punch displacement, l d Die radius, r D Die to punch clearance, c Blank holder force, F B Draw depth, d Blank width, w Blank Length, l Blank Thickness, t Lubrication coefficient, f 30 mm 5 mm 50 mm 5 mm 2.1 mm Various 50 mm 19 mm 150 mm Various 0.15 (Mild oil) Figure 4.2 Dimensions and parameters of finite element model 78

97 4.3.2 Discretisation Based on the geometry of the model, manual meshing was applied instead of automatic meshing. This was because manual meshing gave more advanced parameters to control the element shape, size and accuracy than automatic meshing. Before the model was meshed, the type of element was needed to be selected according to several aspects of the model, such as the model's geometry, the type of deformation, the loads being applied, etc. Due to the contact simulation between the surfaces of each component, first-order elements or modified second-order tetrahedral elements should be used for contact simulations. In addition, significant bending of the blank is expected under the applied loading. Fully integrated first-order elements exhibit shear locking when subjected to bending deformation. Therefore, either reduced-integration or incompatible mode elements should be used. In this analysis, reduced-integration element with enhanced hourglass control CPE4R was mainly selected to mesh the model. The reduced-integration element helps decrease the analysis time, and enhanced hourglass control reduces the possibility of hourglassing in the model. This type of element is suitable for the large non-linear distortion in the finite element analysis. The die, punch, blank holder and blank are set as two dimensional deformable parts. The whole model was meshed mainly with four node plane strain reduced integration element CPE4R. A few linear triangular plane strain elements CPE3 are used for the tooling parts. 79

98 To ensure adequate accuracy and to save computational simulation time, the elements of the die radius and top blank region are meshed more finely than those of the rest of the model. The mesh sizes of the die radius and top blank are 0.03 mm 0.03 mm and 0.05 mm 0.1 mm, respectively Material properties Various material properties were assigned to each component in the model, respectively. For the purpose of the experimental validation, the material properties as shown in Table 4.1 were used in the analysis. Table 4.1 Material properties of mild steel blank and die [5] Blank (Mild steel) Die Material definition Elastic-plastic Elastic Young s Modulus, E 205 GPa 210 GPa Poisson s ratio, v K MPa e n The blank would undergo significant rotation as it is deformed. Reporting the values of stress and strain in a coordinate system that rotates with the blank's motion will make it much easier to interpret the results. Therefore, a local material 80

99 coordinate system that was aligned initially with the global coordinate system, but moves with the elements as they deform, was created for the analysis Contact interaction Contact interactions were defined between the top of the blank and the punch, the top of the blank and the blank holder, and the bottom of the blank and the die. Three contact pairs based on the surface-to-surface contact (standard) were created between the top of the blank and the punch, the top of the blank and the blank holder, and the bottom of the blank and the die, respectively, because large deformation and relative sliding appeared between them, and exact locations of contacting areas in interfaces were not known in advance. Abaqus makes a distinction between analyses where the magnitude of sliding is small and those where the magnitude of sliding may be finite. The finite sliding contact behaviour was selected for sliding formulation in the definition of the contact pairs as the magnitude of sliding between the blank and other components should be considered as finite instead of small. The master surface and slave surface were defined in terms of the stiffness of two corresponding components. The softer surface was selected as the slave surface, while the stiffer surface was selected as the master surface. 81

100 4.3.5 Analysis steps with constraints and loadings Rigid body motion of the components before contact conditions constrains them and sudden changes in contact conditions lead to severe discontinuity iterations as Abaqus tries to establish the exact condition of all contact surfaces. To remove rigid body motion, adequate constraints have to be applied to prevent all rigid body motions of all the components in the model. This may mean using boundary conditions initially to get the components into contact, instead of applying loads directly. Using this approach may require more steps than originally anticipated, but the solution of the problem can proceed more smoothly. To eliminate sudden changes in contact conditions, therefore, though the deep-drawing process is a continuous operation, the simulation run by Abaqus need to be divided into several analysis steps to establish contact between components in a reasonably smooth manner, avoiding large overclosures and rapid changes in contact pressure. The simulation consisted of five steps. As the material, geometric, and boundary nonlinearities were involved in the simulation, general steps have to be used. Step 1 This step was intended to establish firm contact between the blank and the blank holder. In this step the endpoints of the midplane of the blank was fixed in the vertical direction to prevent the blank from moving initially, and the blank holder was pushed down onto the blank using a displacement boundary condition. Given 82

101 the quasi-static nature and nonlinear response of the analysis, a static, general Step 1 was created. Step 2 Since contact was established between the blank and the blank holder and die in the previous step, the constraint on the right end of the blank midplane is no longer necessary and had to be removed in Step 2. Since the previous step considered the effects of geometric nonlinearity, these effects would be included automatically in this and all subsequent steps. Step 3 The magnitude of the blank holder force needed to be introduced in the analysis. In this step the boundary condition used to move the blank holder down would be replaced with a force in Step 3. Step 4 At the beginning of the analysis, the punch and the blank are separated to avoid any interference while contact was established between the blank and the die and blank holder. In this step the punch was moved up in the y direction just enough to achieve contact with the blank. In addition, the vertical constraint on the left end of the blank midplane was removed; and a small pressure was applied to the top surface of the blank to pull it onto the surface of the punch. 83

102 Step 5 In the fifth and final step the pressure load applied to the blank was removed, and the punch was moved up to complete the forming operation. Because of the frictional sliding, the changing contact conditions, and the inelastic material behavior, there was significant nonlinearity in this step. Therefore, the maximum number of increments needed to be set as a large value The initial time increment needed to be input as , the total time period have to be determined to 1.0, and the minimum time increment should be decreased to 1e 06. With these settings Abaqus can take smaller time increments during the highly nonlinear parts of the response without terminating the analysis Deformed and undeformed model Figure 4.3 illustrates the undeformed model, the die, punch, binder holder and blank are illustrated in yellow, blue, green and red, respectively. The punch keeps at the initial position until Step 3 and the blank remains undeformed. 84

103 Die Blank Blank holder Punch Figure 4.3 Undeformed model after Step 3 Figure 4.4 illustrates the deformed model after Step 4. The punch moves up to establish the contact with the blank, but the blank still remains undeformed. 85

104 Figure 4.4 Deformed model after Step 4 Figure 4.5 illustrates the deformed model during Step 5 in early stage. The punch continues moving up to drive the blank begin covering the die radius portion from 0. 86

105 Figure 4.5 Deformed model during Step 5 in early stage Figure 4.6 illustrates the deformed model during Step 5 in middle stage. The punch continues moving up to drive the blank covering whole die radius portion from 0 to

106 Figure 4.6 Deformed model during Step 5 in middle stage Figure 4.7 illustrates the deformed model during Step 5 in late stage. The punch continues moving up to drive the blank covering whole die radius portion and the wall of the die. 88

107 Figure 4.7 Deformed model during Step 5 in late stage Figure 4.8 illustrates the fully-deformed model after Step 5. The punch moves up to 50mm and blank is fully-deformed. 89

108 Figure 4.8 Fully-deformed model after Step Summary In this chapter, Abaqus 6.8 was used as pre-processing, solution and post-processing tools to establish a numerical tool wear prediction model for study of tool wear behaviours of a common stamping part. A customised wear work calculation equation was developed based on Archard equation to be utilised in the finite element analysis. In Abaqus/CAE module, an assembly model which simulated a typical environment and structure of a deep drawing process was 90

109 created. The model consists of a die, a punch, a binder holder and a part. After defining material properties of each part in the model, the model was meshed to produce nodes and elements for the FEA followed by determined boundary constraints and contact conditions. Five steps were applied in the analysis to remove rigid body motion of the components before contact conditions constrain them and sudden changes in contact conditions. After specifying parameters of the solution, Abaqus would solve the specified problem. 91

110 CHAPTER 5 EXPERIMENTAL VALIDATION OF TOOL WEAR PREDICTON MODEL 5.1 Introduction To validate the numerical tool wear prediction model presented in Chapter 4, a series of channel bending tests were conducted with a prescale film. Section 5.2 describes the experimental constraints. Section 5.3 introduces the working principle of Fuji prescale film used in the experiments. Then, experimental equipments are presented in Section 5.4, followed by experimental sequences in Section 5.5. In Section 5.6, experimental results are discussed and compared with the results obtained from the numerical simulations. 5.2 Experimental Constraints The experiment was constrained by a number of key considerations defined by the experimental environment, equipments availability and the research boundaries. These constraints included: (1) A mm 2 Fuji mono-sheet type prescale film was used in the experimental set up as no pressure sensor, which could measure the real-time contact pressures between two surfaces. 92

111 (2) A mild steel strip was used in the experiment instead of a DP steel strip. When forming the DP steel, the high contact pressures led to the failure of the prescale film. So it was necessary to use mild steel to reduce the contact pressures. (3) Only maximum pressure was obtained from the experiment, as the prescale film was only able to record the value of maximum contact pressures. 5.3 Fuji Prescale Film Working principle of prescale film Six types of prescale film are available to cover a wide range of pressure. Figure 5.1 summarises the film types and their corresponding pressure range [82]. A Fuji Pre-scale Film consists of microcapsules filled with colour forming material. When pressure is applied on the film, microcapsules are broken and distribution and density of magenta colour is determined by true pressure distribution and magnitude. When microcapsules are broken, their material is released and it reacts with the colour developing material and this process will cause magenta colour forming. The pre-scale films are designed with Particle Size Control (PSC) Technology. Through PSC technology, microcapsules are designed to react to various degrees of pressures, releasing the colour forming material at a density corresponding to specific levels of applied pressure. 93

112 Figure 5.1 Fuji prescale film types and corresponding pressure range [82] The films are available in two structures: two sheets film and mono sheet film [82]. The film types MS (Medium pressure) and HS (High pressure) are mono sheet types and the other Fuji film types are two sheets types. For two sheets film, prescale is composed of an A film and a C film. The A film is coated with a micro encapsulated colour forming material, and the C film is coated with a colour developing material. The two films should be placed with the coated surfaces facing each other. For mono sheet film, the colour forming layer is coated on the polyester base of film. Micro encapsulated colour forming material is layered on the top of film. According to the pressure range of the experiment, single sheet high pressure films were used Momentary pressure measurement Both extended pressure measurement and momentary pressure measurement can be applied by mean of prescale film [82]. For extended pressure measurement, applied pressure is increased gradually up to the given level, and it will be maintained continuously at that level. In order to get the best and accurate results, the pressure should be applied gradually up to its highest value by a 2 minutes time basis and it should be maintained at the highest level for other 2 minutes. For momentary pressure measurement, application time can be dependent on the application itself. When possible, the given pressure should be applied gradually 94

113 up to its highest magnitude by a 5 seconds time basis, and it should be maintained at the highest level for other 5 seconds. As the momentary pressure measurement simulates the contact between the blank and die radius more realistic than the extended pressure measurement, all measurements in this experimental channel test are momentary pressure measurement. To convert the obtained colour densities from the Fuji Pre-scale films to a pressure, a momentary pressure chart as shown in Appendix B should be used. The momentary pressure chart consists of several curves under various determined temperatures and relative humidity. The proper curve can be selected by determining the temperature and relative humidity first. With these values known, the proper area corresponding to the proper pressure chart curve can be selected from the chart. 5.4 Experimental Equipment Erichsen sheet metal tester A series of channel bend test was conducted in an Erichsen sheet metal tester (Figure 5.2). Erichsen sheet metal tester is widely utilised at research, development and in-process testing. Figure 5.3 shows the schematic of the tester. It can increase drawing speed of the drawing punch, which, in addition to the normal drawing speed range of 0-1,200 mm/min, can be adjusted in an infinitely variable manner and independent of load, up to 3,000 mm/min. This is achieved by using a separate oil circuit, fed by a pump with high volumetric displacement. Contrary to the high speed attachment based on a nitrogen accumulator, here a 95

114 constant drawing speed behaviour is guaranteed over the total displacement of 150 mm. Various deep-drawing process parameters can be monitored from a PC linked to the tester. Figure 5.2 Erichsen sheet metal tester Figure 5.3 Schematic of Erichsen sheet metal tester 96

115 5.4.2 Fuji mono-sheet type prescale film Fuji mono-sheet type high pressure (HS) prescale films were used to measure the maximum contact pressure between the die radius and blank by momentary pressure measurement method Mild steel strip The stress-strain curve of the mild steel was obtained from the tensile test performed on an Instron 5567 Material Test Machine. A stress-strain curve was then fitted using Eq. (5.1) that was subsequently used in the material setup for the corresponding numerical simulation. Eq. (5.1) thus is the material equation for the fitting of the strain-strain curve obtained by the experimental data. Table 5.1 summarises the material properties and fitted values of K, e, n of the mild steel utilised in the channel bending tests. n K e (5.1) where σ is stress, ε is strain, and K, e, n are constants obtained from curve fitting. 97

116 Table 5.1 Material properties and fitted values of K, e, n of mild steel Blank (Mild steel) Material definition Young s Modulus, E Elastic-plastic 205 GPa Poisson s ratio, v 0.3 K 569 MPa e n Experimental Sequences The temperature of the laboratory is 23 and the humidity is 58%. The experimental channel bend tests are divided into six steps (Figure 5.4): Step 1 A flat steel strip was continually formed into a channel section until 5 mm depth; Step 2 Prescale film is placed on the flange surface of the steel strip, which was contacted with the die radius portion of the die insert (Figure 5.5). To avoid movement of the film, Vaseline was spread on the flange surface of the steel strip; Step 3 The strip was then formed from 5 mm to 5.2 mm depth; Step 4 The prescale film was removed, and the strip then continually formed to a channel depth of 20 mm; 98

117 Step 5 Another unused prescale film was placed on the flange surface of the steel strip as for Step 2; Step 6 The strip was then formed from 20 mm to 20.2 mm depth. (a) Step 1 (b) Step 2 (c) Step 3 (d) Step 4 (e) Step 5 (f) Step 6 Figure 5.4 Steps in channel test (black: mild steel strip, red: prescale films) Figure 5.5 Placement of prescale film 99

118 The contact pressure distributions of forming 5.2 mm and 20.2 mm depth channels were obtained from the various colour densities on the two prescale films, respectively. 5.6 Experimental Results and Discussion A comparison of the contact pressure distributions obtained from the experimental channel bend test and the simulation is presented in Figure 5.6. Table 5.2 summaries the comparison of locations of contact pressure peaks from 0 angle onward. The distinct strips identified from the prescale film indicate contact pressure peaks. The density of red colour illustrates the values of the contact pressure peak, however, the maximum value which can be determined is limited to the maximum pressure that the prescale film can measure. From the prescale film for the forming of the 5.2 mm depth channel, two distinct pressure regions or stripes can be identified. The left stripe appears between 0 mm and 1 mm from the start of the die radius (0 ) and the right one is at 2 mm from the start of the die radius. Compared with the sample chart of the colour densities, the contact pressure indicated by the left right stripe exceeds 130 MPa, the highest threshold of the measuring capability of the prescale film, and the contact pressure of the next stripe is between 82 MPa and 98 MPa. The results from simulations indicate that the first peak appears between 0 mm to 1 mm from 0 and its value is 205 MPa. The second peak is located at the position of 2 mm from 0 and its value is 91 MPa. 100

119 Prescale film for 20.2 mm drawing depth Prescale film for 5.2 mm drawing depth MPa Figure 5.6 Comparison of contact pressure distributions obtained from tests and simulations Two stripes can also be identified on the prescale film at a forming depth of 20.2 mm. The left stripe, between 0 mm and 1 mm from the start of the die radius, is distinct and the contact pressure indicated by this strip exceeds 130 MPa. The right one ranging from 3 mm to 6 mm is less distinct with a contact pressure from 50 MPa to 66 MPa. The results from the simulations reveal that the first peak of contact pressure of 223 MPa appears between 0 mm and 1 mm from 0, and the second peak of contact pressure of 61 MPa is located between 3 mm and 6 mm from 0. Table 5.2 Comparison of locations of contact pressure peaks from 0 Mild steel strip forming range 5 mm to 5.2 mm 20 mm to 20.2 mm Peak Number Location of contact pressure peaks from 0 On prescale film On simulation graph First peak 0 1 mm 0 1 mm Second peak 2 mm 2 mm First peak 0 1 mm 0 1 mm Second peak 3 6 mm 3 6 mm 101

120 Comparing the results obtained from the prescale film with the results from the simulation, it is concluded that the contact pressure distributions indicated by the prescale films are consistent with those from the simulation. 5.7 Summary In this chapter, a series of channel bending test were conducted with prescale film. The experimental constraints were presented. The working principle of Fuji prescale film used in the experiments was introduced. Experimental equipments and experimental sequences were discussed. The experiments have validated the numerical tool wear prediction model presented in Chapter 4 and shown that the contact pressure distributions indicated by the prescale films are consistent with those from the simulation.the numerical tool wear prediction model will be utilised in the investigation of die radius profile on wear behaviour, the investigation of control parameters on wear behaviour and the optimisation of die radius profile as described in subsequent chapters. 102

121 CHAPTER 6 INVESTIGATION OF DIE RADIUS ARC PROFILE ON WEAR BEHAVIOUR 6.1 Introduction Tool geometry plays an important role in affecting the tool life of sheet metal stamping of AHSS. However, very little work seems to have been done on determination of exact die geometry to reduce tool wear in sheet metal stamping of AHSS. This chapter presents the influences of die arc profile on wear behaviour for AHSS. Section 6.2 illustrates variation of die radius profiles. In Section 6.3, the effects of various geometries of radius arc profiles, including standard circular profiles, high elliptical profiles, and flat elliptical profiles, on the wear volume and contact pressure distribution along the radii are discussed. 6.2 Variation of Die Radius Profiles Cases with various die radius profile were studied using the finite element tool wear model as illustrated in Figure 6.1. The blank material is AHSS DP780 with a width of 25 mm. Table 6.1 summarises the material properties of the blank and the tools used in the model. 103

122 Table 6.1 Material properties of DP780 blank and die [5] Blank (DP780) Die Material definition Elastic-plastic Elastic Young s Modulus, E 205 GPa 210 GPa Poisson s ratio, v Yield strength, 480 MPa - Ultimate Tensile Strength, 780 MPa - Three regular types of the die radius profile (Table 6.2 and Figure 6.1) were investigated in the simulations, including standard circular curves, high elliptical curves and flat elliptical curves. CR15 CR10 CR5 HER5r15 HER5r10 HER5r5 (CR5) FER15r5 FER10r5 FER5r5 (CR5) (a) (b) (c) (a) Standard circular profile; (b) High elliptical profile; (c) Flat elliptical profile Figure 6.1 Three regular types of die radius profile 104

123 Table 6.2 Various die radius profiles used in simulations Case Number Shape of die radius Radius (mm) CR5 Circular curve 5 CR10 Circular curve 10 CR15 Circular curve 15 Case Number Shape of die radius Radius in x direction (mm) Radius in y direction (mm) HER5r10 High elliptical curve 5 10 HER5r15 High elliptical curve 5 15 FER10r5 Flat elliptical curve 10 5 FER15r5 Flat elliptical curve 15 5 An assembly tool wear prediction model for Abaqus developed in Chapter 4 is used for the investigation. The model consists of a die, a punch, a binder holder and a part. After defining material properties of each part in the model, the model was meshed to produce nodes and elements for the FEA followed by determined boundary conditions. However, to obtain higher accuracy, the die radius arc was meshed more finely. The contact characteristics between each component need to be defined as well. A series of steps with loading was applied on the punch to drive the deep-drawing process to stretch the part. These procedures are referred to as the pre-processing procedures. After specifying parameters of the solution, Abaqus would solve the specified problem. The tool wear work results are obtained from the post-processing of the simulations. 105

124 6.3 Results and Discussion Standard circular profiles From wear work model of Eq (4.9), it is concluded that the contact pressure plays a significant role in tool wear. Figures show the contact pressure over the die radius as a colour contour graph. Three cases, including CR5, CR10 and CR15, are studied. With the standard circular profiles, the maximum contact pressure was reduced as the radius increased. The maximum contact pressure applied on the die radius with CR5 profile was 1130 MPa, which is approximately twice that of CR15 profile. It is found that the colour contour for all three cases can be divided into three distinct zones, which are caused by different mechanisms. Zone 1 is located at the area near 0. In this area, the blank is constrained by the blank holder pressure and is restricted to slide over the die radius (Figure 6.5). The contact pressure in this zone decreases as the die radius increases. Zone 2 is a bold straight line from the bottom left corner to the upper right corner in the contour plots. The slope of the line is related to the radius of the circular profile. As the radius increased, so did the slope of the line. The high contact pressure revealed in this zone is caused by the relative tangential sliding movement between the blank and the die radius which results in the concentrated contact force at the tangent point (Figure 6.5). Because the tangent point is moving toward 90 during the punch travel instead of remaining fixed, the shape of Zone 2 is shown as a straight line from bottom left to the upper right corner. However, Zone 2 does not cross the entire die radius and entire punch travel as the 106

125 blank and the die radius completely overlap after a certain time point. The maximum contact pressure was located at Zone 2. Zone 1 Angular interval of Zone Zone 3 Zone 2 Figure 6.2 Contact pressure over die radius with CR5 profile Zone 1 Angular interval of Zone 3 Zone 3 Zone 2 Figure 6.3 Contact pressure over die radius with CR10 profile 107

126 Zone 1 Angular interval of Zone 3 Zone 3 Zone 2 Figure 6.4 Contact pressure over die radius with CR15 profile Zone 1 Zone 2 Figure 6.5 Cause of high contact pressure of standard circular profiles Zone 3 The angular interval of Zone 3 was from 20 to 40 with various beginning and ending angles depending on the radius of the profiles. The larger the radius, the wider the angular interval was. As the radius increased from 5 mm to 15 mm, the interval moved toward 90. With the CR5 profile, the beginning angle was 10 and the ending angle was 30. The beginning and ending angles increase to 35 and 60, respectively, for the CR10 profile. As the radius increased to 15 mm radii, the angular interval was located between 30 and 70. The critical high contact 108

127 pressure in Zone 3 results from the relative sliding movement between the die radius and the blank after the die radius and the blank completely overlap (Figure 6.5). Figure 6.6 shows the wear work over the die radius with the standard circular profiles. It is concluded that the maximum wear work is located at the zone near 0 and can be reduced by enlarging the radius of the circular profile. Besides the maximum wear work, another peak value of the wear work occurs at the area between 10 to 30, 35 to 60, and 30 to 70, with CR5, CR10 and CR15 profiles, respectively. This is mainly caused by the relative tangential sliding movement between the blank and the die radius. Figure 6.6 Wear work over die radius with standard circular profiles 109

128 6.3.2 High elliptical profiles Figures 6.7 and 6.8 show the contact pressure over die radius with high elliptical profiles with y-radius of 10 mm and 15 mm respectively. Compared with the standard circular profiles, the three zones are less distinct. Zone 1 with critical high contact pressure is narrower than that of the standard circular profiles. It suggests that the sharp curvature of the die radius near 0 leads to a concentration of force and restriction of the steel strip from sliding freely (Figure 6.9). The maximum contact pressure with HER5r15 and HER5r10 high elliptical profiles are twice and three times as large as that of CR5 circular profile, respectively. However, enlarging of the length of the y axis has no dominant influence on the maximum contact pressure. Figure 6.7 Contact pressure over die radius with HER5r10 profile 110

129 Figure 6.8 Contact pressure over die radius with HER5r15 profile Figure 6.9 Cause of high contact pressure of high elliptical profile As the blank cannot completely overlap the die radius with the high elliptical profiles during the punch travel, the zones with high contact pressure caused by the tangent and overlapped sliding movement (Zones 2 and 3) are relatively small compared with those of the circular profiles (Figure 6.9). 111

130 Figure 6.10 shows the wear work over the die radius with high elliptical profiles. The maximum wear work lies mainly at the zone near 0 with the high elliptical profiles, which is approximately twice that of CR5 profile. It is noted that this zone is the dominant wear location of the die radius with these profiles. However, the profiles show very low wear work for almost all zones of the die radius. Thus the application of high elliptical profiles seems to have a significant influence on reducing the wear work at the area between 3 and 90, compared with that of the standard circular profile, despite the location of the peak moving towards to 90. Figure 6.10 Wear work over die radius with high elliptical profile Flat elliptical profiles Figures 6.11 and 6.12 show the contact pressure over die radius with flat elliptical profiles with x-radius of 10 mm and 15 mm respectively. The critical locations of the high contact pressure of the flat elliptical profiles can also be divided into 112

131 three zones. Zones 1 and 2 are similar to those of the standard circular profiles. Figure 6.11 Contact pressure over die radius with FER10r5 profile Figure 6.12 Contact pressure over die radius with FER15r5 profile Because of the larger contact area between the die and blank, the contact pressure 113

132 of Zone 1 is reduced (Figure 6.13). Zone 3 is located at the area near 90 with a discrete character in values of the contact pressure. The maximum contact pressure is located at Zone 3 due to the shape curvature of the area near 90. However, the increase of the length of the x axis has no significant influence on the maximum contact pressure of the die radius with flat elliptical profiles. Figure 6.13 Cause of high contact pressure of flat elliptical profile Figure 6.14 Wear work over die radius with flat elliptical profiles 114

133 Figure 6.14 shows wear work over the die radius with the flat elliptical profiles. It is noted that the maximum wear work of the flat elliptical profiles is located at the zone near 90. The zone near 0 is still a high tool wear location for the flat elliptical profiles, but the values are lower than the high elliptical profiles and radius profile. It is seen that increase of the length of the x axis can decrease the tool wear near in the middle zones, but has no significant influence on another peak near Summary This study investigated the influence of various die radius profiles on the tool wear parameters, including contact pressure and accumulated wear work. The following conclusions are drawn from the results: (1) The colour contour of the high contact pressure on the die radius can be divided into three distinct zones in all cases. Each zone reveals the different characteristics of the cause and pattern of the high contact pressure as well as tool wear. The reaction force caused by the blank is constrained by the blank holder pressure and is restricted from sliding over the die radius. This results in a high contact pressure in Zone 1. Relative tangential sliding movement between the die and blank leads to high contact pressure in Zone 2. Critical contact pressure in Zone 3 is produced by the overlapped movement of the die and blank. (2) The dominant zone leading to maximum contact pressure and tool wear severity depends on the geometry of die radius profile under the same material and process conditions. Both Zones 1 and 2 are critical for the standard circular 115

134 profiles. However, Zones 1 and 3 play a significant role for high and flat elliptical profiles, respectively. (3) For standard circular profiles, the maximum contact pressure and tool work drops significantly when the radius increases. (4) For high elliptical profiles, the most critical high contact pressure and tool wear work is located at the area near 0. The high elliptical profiles produce low wear work during 90% of the die radius zones than the standard circular profile or flat elliptical. However, enlarging of the length of the y axis has no dominant influence on the maximum contact pressure. (5) For flat elliptical profiles, the area near 90 is critical for high contact pressure and tool wear work. However, the profile does not provide better wear work compared to other profiles. The increase of the length of the x axis has no significant influence on the maximum contact pressure. (6) There are two peaks of the accumulated tool wear work in all cases, but the locations vary. The value and location of the peaks depends on the various influences of the three zones with high contact pressure. The zone near 0 is the common zone for server tool wear in all cases. However, for the flat elliptical profile, the zone near 90 is the severest tool worn area. (7) The geometry of draw die radius has a significant influence on the tool wear, and standard circular and high elliptical curves can lead to the achievement of reduced and uniform wear distribution along most of the zones of the draw die radius arc. (8) The results suggest that to minimise tool wear using this approach it would be necessary to optimise the shape for a particular combination of circular and high elliptical profiles in relation to the material type, thickness and forming process. 116

135 CHAPTER 7 INVESTIGATION OF CONTROL PARAMETERS ON WEAR BEHAVIOUR 7.1 Introduction Control parameters in stamping process do affect the distribution of wear behaviour. The wear, especially adhesive wear, varies with the change of control parameters including lubrication coefficient, material strength, and blank thickness in deep-drawing process [5, 19, 64, 83]. However, previous work regarding the influences of these control parameters on tool wear distribution mainly focused on the circular die radius profile. To study the influences of various control parameters on tool wear distribution for various die radius geometries, this chapter investigates the effects of process control parameters on the severity of wear in deep-drawing process using numerical simulations. Section 7.2 illustrates the types of control parameters and material properties used in this study. In Section 7.3, the influence of these control parameters on tool wear work with different die radius profiles, including a circular profile, a flat elliptical profile and a high elliptical profile, is presented in detail. 7.2 Variation of Control Parameters Cases with various control parameters for different die radius profile are studied using the finite element tool wear model illustrated in Figure The material of the blank strip is AHSS DP780 and the width of the blank strip is 25 mm. Table 117

136 7.1 summarises the material properties of blank and tools used in this investigation. Table 7.1 Material properties of DP780 blank and die [5] Blank (DP780) Die Material definition Elastic-plastic Elastic Young s Modulus, E 205 GPa 210 GPa Poisson s ratio, v Yield strength, Ultimate Tensile Strength, 480 MPa 780 MPa Three types of the die radius profile (Table 7.2) and six control parameters (Table 7.3) were investigated in the simulations, including standard circular curves, high elliptical curves and flat elliptical curves. Table 7.2 Die radius profiles in simulations Case Number Shape of die radius Radius in x direction (mm) Radius in y direction (mm) CR5 Circular curve 5 5 F510 Flat elliptical curve 10 5 H510 High elliptical curve

137 Table 7.3 Control parameters in simulations Control parameters Notation Values Lubrication coefficient LC 0.10, 0.15, 0.20 Binder holder force (kn) BHF 10, 20, 30 Young s modulus of die (GPa) EX 190, 210, 230 Clearance between die and punch (mm) C 0.1, 1.1, 2.1 Punch radius (mm) P 2, 5, 8 Punch diameter (mm) PD 15, 30, 45 Blank thickness (mm) T 1.5, 2, 2.5 The lubrication fluid acts as a barrier to separate the tool surface from the sheet material and then decreases the interface strength between the contacting surface asperities [60]. It is important to understand the influence of the lubrication on the tool wear distribution in sheet metal stamping, especially in forming complicated automotive parts using AHSS. Material properties, such as Young s modulus of die and blank thickness may influence the tool wear distribution due to the enhancing effect of material strength and thickness on the contact stresses between the tool surface and the metal sheet [76, 83]. Parameters including binder holder force, clearance between die and punch are reported as having significant effects on the tool wear distribution in sheet metal stamping, while other parameters such as Poisson s ratio had limited influence on the tool wear distribution [5, 19]. An assembly tool wear prediction model for Abaqus developed in Chapter 4 is used for the investigation. The model consists of a die, a punch, a binder holder 119

138 and a part. After defining material properties of each part in the model, the model was meshed to produce nodes and elements for the FEA followed by determined boundary conditions. However, to minimise the computation time, the die radius arc was discretised into 30 segments. The contact characteristics between each component need to be defined as well. A series of steps with loading was applied on the punch to drive the deep-drawing process to stretch the part. These procedures are referred to as the pre-processing procedures. After specifying parameters of the solution, Abaqus would solve the specified problem. The tool wear work results are obtained from the post-processing of the simulations. 7.3 Results and Discussion Lubrication coefficient Figure 7.1(a) to Figure 7.1(c) show the variation of tool wear work over the various die radii angles with lubrication coefficients (LC) 0.10, 0.15 and 0.20 for three types of die radius profiles, i.e. circular profiles, flat elliptical profiles and high elliptical profiles. For the circular profiles, as shown in Figure 7.1(a), the wear work for all three cases peaks at the location near 0 with similar values. It reflects that the lubrication coefficient has less effect on the first peak of the wear work. However, in all three cases, though the positions of the second peak are similar and range from 30 to 50, the values vary. Higher lubrication coefficient leads to higher wear work. 120

139 For the flat elliptical profiles, as shown in Figure 7.1(b), in all cases, the values of the wear works are similar. The first peak appears at the location near 0. Then the values of the wear work reduce to near zero and climb dramatically to approximately 25 GPamm at the location of 80. The lubrication coefficient has insignificant effects on the tool wear distribution for the flat elliptical profiles. (a) Circular die radius (b) Flat elliptical die radius (c) High elliptical die radius Figure 7.1 Wear work over die radius with various lubrication coefficients for three die radius arc profiles For the high elliptical profiles, as shown in Figure 7.1(c), in all cases, the maximum tool wear work rises dramatically at the zone near 0 which is 121

140 approximately two times larger than that of the circular profile. Then the values of the wear work reduce sharply to near zero and the second peaks of all cases appear after the location of 50 with relatively smaller values. The dominant worn location of the flat elliptical profile is located near 90. It is concluded that the lubrication coefficient has less effect on the first peak of wear work but can vary the location of the second peak Binder holder force Figure 7.2(a) to Figure 7.2(c) show the variation of tool wear work over the various die radii angles with various binder holder forces (BH) 10 kn, 20 kn and 30 kn for the three types of die radius profiles. For the circular profiles, in all cases, high binder holder pressure force causes high wear work over the die radius, which results in the non-uniform tool wear pattern. The first peaks of the wear work in all cases appear near the locations between 5 and 10. And then the wear work reduces to lower values. The locations of second peaks in three cases vary between 20 and 40 and their values have small differences as well. The binder holder force also has no impact on the wear work for flat elliptical profiles. The first peak of the wear work appears at 5 and then the wear work decreases gradually to zero and climbs sharply again to the second peak near 80. The binder holder force also has no impact on the wear work for high elliptical profiles. The first peak of the wear work appears at 3 which is approximately two 122

141 times larger than that of the circular profile. And then the wear work decreases dramatically to zero and climbs again to the second peak at 70 with a smaller value compared with the value of the first peak. (a) Circular die radius (b) Flat elliptical die radius (c) High elliptical die radius Figure 7.2 Wear work over die radius with various binder holder forces for three die radius arc profiles Young's modulus of die Figure 7.3(a) to Figure 7.3(c) show the variation of tool wear work over the 123

142 various die radio angles with various die materials for the three types of die radius profiles. The variation in Young s modulus (EX) 190 GPa, 210 GPa and 230 GPa represents the variation in die materials. It is concluded that Young s modulus of the die material has no significant influence on wear work for all die radius profiles. (a) Circular die radius (b) Flat elliptical die radius (c) High elliptical die radius Figure 7.3 Wear work over die radius with various Young s modulus of die for three die radius arc profiles For the circular profiles, the first peak of the wear work in all cases appears near 124

143 the location of 10. And then the wear work reduces to a lower value. The locations of second peaks in three cases are near 30 with relatively low values. The Young s modulus of die material has no impact on the wear work for flat elliptical profiles. The first peak of the wear work appears at 5 and then the wear work decreases gradually to zero and climbs sharply again to the second peak near 80. The Young s modulus of die material has no impact on the wear work for high elliptical profiles as well. The first peak of the wear work is approximately two times larger than that of the circular profile, which appears at the 3. And then the wear work decreases dramatically to zero and climbs again to the second peak at 70 with a smaller value compared with the value of the first peak Clearance between die and punch Figure 7.4(a) to Figure 7.4(c) show the variation of tool wear work over the various die radii angles with various clearances between die and punch (C) 0.1 mm, 1.1 mm and 2.1 mm for the three types of die radius profiles. The clearance has less impact on the wear work of the circular profiles, though it can result in slight difference on the second peaks of wear works in the location between 30 and 50. However, it has no significant influence on the first peak of the wear work. 125

144 However, the clearance has significant influences on the wear work of flat elliptical profiles. Less clearance causes dramatically higher second peaks of the wear works of these profiles on the location between 60 and 90. However, it has no noticeable effects on the first peak of the wear work. (a) Circular die radius (b) Flat elliptical die radius (c) High elliptical die radius Figure 7.4 Wear work over die radius with various clearances between die and punch for three die radius arc profiles For the high elliptical profiles, the clearance has no remarkable influence on the first peak of wear work. However, it has limited influence on the second peak of the wear work. Larger clearance leads to the decreased second peak of the wear 126

145 work Punch radius Figure 7.5(a) to Figure 7.5(c) show the variation of tool wear work over the various die radii angles with various punch radius (P) 2 mm, 5 mm and 8 mm for the three types of die radius profiles. It is concluded that punch radius has no significant influence on wear work for all die radius profiles. For the circular profiles, in all cases, the wear work peaks near the location of 10. And then the wear work decreases gradually. At the location near 30, the wear work of all three cases reach their second peaks with relatively low values. The punch radius has no impact on the wear work for flat elliptical profiles. The wear work achieves its first peak at 5 and then the wear work decreases gradually to zero and climbs sharply again to the second peak near 80. The punch radius has no impact on the wear work for high elliptical profiles as well. At the location of 3, the wear work reaches its first peak, at which the value is two times larger than that of the circular profile. And then the wear work decreases significantly to zero and climbs again to the second peak at 70 with a smaller value compared with the value of the first peak. 127

146 (a) Circular die radius (b) Flat elliptical die radius (c) High elliptical die radius (b) Flat elliptical die radius (c) High elliptical die radius Figure 7.5 Wear work over die radius with various punch radius for three die radius arc profiles Punch diameter Figure 7.6(a) to Figure 7.6(c) shows the variation of tool wear work over the various die radii angles with various punch diameters (PD) 15 mm, 30 mm and 45 mm for the three types of die radius profiles. 128

147 For the circular profiles, the punch diameter has no significant effect on the wear work. In all cases, the wear work peaks near the location of 10. And then the wear work decreases gradually. At the location near 30, the wear work of all three cases reach their second peaks with relatively low values. (a) Circular die radius (b) Flat elliptical die radius (c) High elliptical die radius Figure 7.6 Wear work over die radius with various punch diameters for three die radius arc profiles 129

148 The punch diameter has limited impact on the wear work for flat elliptical profiles. The peak values of wear work with 30 mm punch diameter are slightly more than those with 15 mm and 45 mm punch diameters. The wear work achieves its first peak at 5 and then the wear work decreases gradually to zero and climbs sharply again to the second peak near 80. The punch diameter has very slight influence on the wear work for high elliptical profiles as well. The 30 mm punch diameter results in largest peak values in three cases. At the location of 3, the wear work reaches its first peak, at which the value is two times larger than that of the circular profile. And then the wear work decreases significantly to zero and climbs again to the second peak at 70 with a smaller value compared with the value of the first peak Blank thickness Figure 7.7(a) to Figure 7.7(c) show the variation of tool wear work over the various die radii angles with various blank thicknesses (T) 1.5 mm, 2.0 mm and 2.5 mm for the three types of die radius profiles. It is noted that punch radius has significant influences on the second peak value of wear work for all die radius profiles. For the circular profiles, in all cases, the wear work peaks near the location of 10. And then the wear work decreases gradually. The change of blank thickness has limited effects on the first peak value. However, the increase of blank thickness leads to higher second peak value of wear work. 130

149 For flat elliptical profile, the variation of thickness has less effect on the first peak value. However, the variation of the thickness results in the change of locations for the second peak values, at which the locations are ranging from 70 to 90. (a) Circular die radius (b) Flat elliptical die radius (c) High elliptical die radius Figure 7.7 Wear work over die radius with various blank thicknesses for three die radius arc profiles The variation of thickness also results in noticeable changes of the second peak, both location and value, for the high elliptical profile. However, the influence of the thickness on the first peak of the wear work is less significant compared with 131

150 that of the first peak. 7.4 Summary It is concluded that various control parameters have different impacts on the tool wear of die radius, depending on the specified die radius profile. Table 7.4 summarises the impact of these parameters. For the circular profile, lubrication coefficient, binder holder force and blank thickness play critical roles in the wear work distribution. Clearance between die and punch and blank thickness can significantly affect the wear work distribution for the flat elliptical profiles. Lubrication coefficient, clearance between die and punch and blank thickness are three major factors which control the wear work distribution for the high elliptical profiles. 132

151 Table 7.4 Impacts of control parameters on wear work (S: Significant impact; L: Less impact; N: No impact) Control parameters Circular profile Flat elliptical profile High elliptical profile Lubrication coefficient S N S Binder holder force S N N Young s modulus of die N N N Clearance between die and punch L S S Punch radius L N N Punch diameter N L L Blank thickness S S S 133

152 CHAPTER 8 OPTIMISATION OF DIE RADIUS GEOMETRY 8.1 Introduction In Chapter 6, it was concluded that to minimise tool wear using the approach of varying tool radius profile, it would be necessary to optimise the shape for a particular combination of circular and high elliptical profiles in relation to the material type, thickness and forming process. This chapter presents a methodology to optimise a die radius profile. For this, a specialised software routine is developed and compiled for optimisation of die radius profiles to minimise or achieve uniform contact pressure (wear distribution) using Python computer programming language. Python computer programming language is the programming tool supported by Abaqus. Section 8.2 presents the Graphical User Interface of the specialised software routine. In Section 8.3, a detailed algorithm for the optimisation is explained. A case study based on the algorithm is discussed in Section Graphical User Interface As Abaqus is commercial general finite element software, establishment of a tool wear prediction model for a unique combination of geometries and control parameters will require a time-consuming trial and error approach by designers. To simplify the task, first a graphical user interface (GUI) is developed by Python 134

153 programming language. The use of GUI provides the following advantages: (1) It dramatically decreases the time on the modelling to few minutes instead of hours; (2) It provides a convenient visualised GUI for designers to directly input their specified geometric and control parameters instead of considering details of the modelling; (3) It ensures the consistency of all cases, such as the coordinate system, datum surface and datum axis, and it guarantees the standard and accuracy of the FEA simulations. Figures illustrate the GUI created using the Python computer programming language. The GUI consisted of three tabs for geometry, process parameters and simulation settings. In the tab named Geometry, designers can input all geometric parameters of the model without considering detailed modelling procedures. In the tab called Process Parameters, various control parameters can be directly input into the corresponding spaces. And simulation settings can also be specified in the simulation setting tab. The whole simulation can then be automatically run by clicking the OK button. However, users can always preview the model in advance by clicking the Preview button. 135

154 Figure 8.1 GUI for Geometry created using Python programming language 136

155 Figure 8.2 GUI for Process Parameters created using Python programming language 137

156 Figure 8.3 GUI for Simulation Setting created using Python programming language 138

157 8.3 Algorithm for Die Radius Optimisation To optimise the die radius profile, a customised algorithm is developed and applied in the optimisation process. The main point of the algorithm is to optimise the die radius geometry by changing the effective radii along the die radius profile through the iteration process. B R i,j node i A Figure 8.4 Die radius profile W i - W S S Figure 8.5 Accumulated wear work along die radius 139

158 The initial geometry of the die radius is set as a standard circular profile as shown in Figure 8.4. The whole profile is divided into n equal segments. Each division point will be used as a node for meshing the profile in Abaqus. For example, point i is also node i in the finite element model, where i = 0 ~ n. In the proposed algorithm, the first point A and the last point B of the profile (See Figure 8.2) are assumed to be fixed. It is assumed that after m th simulation, the optimised profile is obtained, where j = 1 ~ m. The algorithm will generate an optimised die radius profile. The optimised die radius profile should provide a uniform wear work distribution at all points on die radius angles. Let us consider a case of an unoptimised wear work distribution profile as shown in Figure 8.5. Let us consider that the optimised wear work distribution profile is a uniform sinusoidal type of distribution as shown in Figure 8.5. This uniform distribution profile has a maximum variation of magnitude of S with respect to the mean line (red line) as shown in that figure. Let us define the following parameters for the development of the proposed algorithms. R i,j Effective radius on the node i in the simulation j W i.j Accumulated wear work on the node i in the simulation j W Nominated average accumulated wear work required (Red line in Figure 8.3) S Maximum variation of accumulated wear work allowed S i,j Variation of accumulated wear work on the node i in the simulation j, S i,j = W i,j W (Dash lines in Figure 8.5) 140

159 Figure 8.6 Flow chart of proposed algorithm f i,j Control coefficient for changing the effective radius f i,j = k R i -R i-1 k Constant defined by user to determine the control coefficient f n Number of segments along die radius profile 141

160 m Number of simulations (If the number of simulations exceeds m and the optimised result is still not found, the whole simulation is terminated) The flow chart of the functioning of the algorithm is as shown in Figure 8.6. According to this algorithm, the Abaqus simulation will take the input data and work out the wear work for each point selected for the die radius profile of Figure 8.4. Thus, after each simulation, the accumulated wear work is calculated for each node. Effective radius of each node is then adjusted according to the difference between the nominated average accumulated wear work and actual accumulated wear work. If the accumulated wear work is much larger (exceeding the variation allowed), the effective die radius would be decreased. If the accumulated wear work is much smaller (below the variation allowed), the effective die radius would be increased. If the accumulated wear work is averaged (within the variation allowed), then the effective die radius would remain unchanged. This can be written as follows: If S i,j > S, then R i,j+1 = R i,j f S i,j If S i,j < S, then R i,j+1 = R i,j The next simulation is run after the adjustment of effective die radii (the position of nodes). The simulation will finish if the variation between the nominated average accumulated wear work and actual accumulated wear work for all nodes is within the maximum variation allowed, i.e. the accumulated wear work result curve is within the space between two dash lines in Figure

161 8.4 Case Study Optimisation parameters settings To apply the GUI and the algorithm developed in the previous sections, a case study will be presented for the circular die profile. Circular curve with 5 mm die radius, i.e. CR5 curve, is selected as the original un-optimised curve. The blank material is AHSS DP780 with a width of 25 mm. Table 8.1 shows the material properties of the blank and the tools. Table 8.1 Material properties of DP780 blank and die Blank (DP780) Die Material definition Elastic-plastic Elastic Young s Modulus, E 205 GPa 210 GPa Poisson s ratio, v Yield strength, 480 MPa - Ultimate Tensile Strength, 780 MPa - The die radius curve is divided into 30 divisions (See Figure 8.7), and the effective radius R at these 31 division points can be adjusted after each simulation. After each simulation, the average accumulated wear work is calculated, and then the effective radius R of these points is adjusted according to the prescribed 143

162 algorithm. Due to the time-consuming simulations (approximately 5 hours for each single simulation), only 20 simulation loops are performed during the optimisation. The optimised die radius profile will be selected from the results of these simulations using lowest mean value. Figure 8.7 Divisions of die radius profile Results and discussion Table 8.2 summaries the results of the simulation for the effective radius R at the 31 division points for CR5 curve and the optimised curve. Figure 8.8 illustrates the positions of 31 division points of optimised curve. 144

163 Table 8.2 Effective radius R for CR5 and optimised curves Point i R CR5 R Optimised Point i R CR5 R Optimised

164 Figure 8.8 Positions of division points of optimised curve Mean value of wear work for CR 5 Mean value of wear work for optimised curve Figure 8.9 Wear work over die radius for CR5 and optimised curves Figure 8.9 shows the plots of the wear work over the die radius computed for CR5 curve and the optimised curve. It is noticed that the wear work distribution is not uniform for the circular radius profile (CR5), and it has a peak wear work of GPamm near 0 die angle. But when using the optimised radius profile, the wear work distribution becomes more uniform and oscillating, as noted in the wear 146

165 work distribution of Figure 8.9. It shows a peak of GPamm at 5 die angle, then goes down to near zero and peaks again with a magnitude of 6.08 GPamm at 37 die angle. It is concluded that the maximum wear work was reduced from GPamm to GPamm. However, besides the reduction of the maximum wear work, another peak value of the wear work occurs at the area between 30 to 50 with the value of 6.08 GPamm. Though another peak value of the wear work appears by applying the optimised die radius profile, the reduction of the maximum wear work compensates it to achieve relative low average wear work. Figure 8.9 also shows that the average wear work obtained from the optimised radius profile is much lower than the average wear work given by the un-optimised circular radius profile. Table 8.3 shows a comparison of maximum wear works and average wear work obtained by the unoptimised circular profile and optimised profile of the developed computer routine. Table 8.3 Comparison of wear work of un-optimised circular profile with optimised one Die radius profile Maximum wear work (GPamm) Average wear work (GPamm) Un-optimised circular profile Optimised profile Due to the considerable computation time, it was not possible to work out wear 147

166 work distribution for high elliptical and flat elliptical die radius profile, but it is concluded that the proposed methodology of the die radius optimisation will deliver similar reduction on wear work distribution for these two geometries as well. 8.5 Summary In this chapter, to minimise and achieve uniform contact pressure (wear distribution), a methodology based on a specialised software routine was introduced for optimisation of die radius profiles using Python programming language, which was fully integrated with Abaqus software. The algorithm provides the following functions: (1) To provide a user-friendly Graphical User Interface for pre-processing of data input for users who have less experiences and skills; (2) To optimise a die radius profile according to the control parameters that users input. The case study discussed in the chapter shows that the routine was suitable for optimisation of a die radius profile, thought it may require time-consuming iterative simulations in Abaqus software. 148

167 CHAPTER 9 CONCLUSIONS AND FURTHER RESEARCH 9.1 Overview This research has presented an investigation on the development of tool wear prediction model to study the influences of draw die geometry on the wear distribution over the draw die radius for AHSS material. The research also present a methodology for optimising the draw die geometry to reduce wear using numerical methods by developing a specialised software routine using Python programming language and implemented in Abaqus finite element analysis software. 9.2 Major Research Outcomes Tool wear predictions on automotive sheet metal forming die and recommended protections of the tool surface under the initial production conditions were obtained from AutoForm simulation software. Effects of lubrication coefficients, binder pressure loads and die coating on tool wear distributions were investigated as well. It is concluded that the areas that are most sensitive to the tool wear occur at the locations corresponding to the large gradient of drawing depth. To study the tool wear distributions for more common stamping parts, a numerical tool wear model was developed and applied using the commercial software 149

168 package Abaqus. Channel tests are carried out using an Erichsen sheet metal tester with high pressure prescale films to verify the numerical model results. Comparing the results obtained from the prescale film with the results from the simulation, it is concluded that the contact pressure distributions indicated by the prescale film are consistent with those from the simulation. Various geometries of radius arc profiles, including standard circular profiles, high elliptical profiles, and flat elliptical profiles, were numerically investigated using the tool wear model developed, and the contact pressure distribution and tool wear work along the radii were determined. The following conclusions were reached from the investigations: (1) The area (as plotted by colour contour) of the high contact pressure on the die radius can be divided into three distinct zones of high pressure and tool wear; (2) The dominant zone leading to maximum contact pressure and tool wear severity depends on the geometry of die radius profile under the same material and process conditions; (3) The geometry of draw die radius has a significant influence on the tool wear and standard circular and elliptical curves can lead to the achievement of reduced and uniform contact pressure distribution (wear distribution) along most of zones of the draw die radius arc. The results suggest that to minimise contact pressure and tool wear using this approach it would be necessary to optimise the shape of the die for a particular combination of material type, thickness and forming process. 150

169 Effects of control parameters, such as blank geometry, punch geometry, deep-drawing process parameters, blank material and tool material, on wear behaviour in deep-drawing for various shape of die radius were then investigated to provide guidelines for impacts of these parameters. A specialised software routine was then compiled for optimisation of die radius profiles to minimise and achieve uniform contact pressure (wear distribution) using Python programming language. The routine was fully integrated with Abaqus software and has the following functions: (1) To provide a user-friendly Graphical User Interface for pre-processing data input for users who have less experiences and skills; (2) To optimise a die radius profile according to the control parameters that users input. The following major research outcomes were achieved in this project: (1) Prediction and identification of critical tool worn area on GM Holden s sheet metal forming die using AutoForm simulation software; (2) Establishment of a numerical tool wear prediction model of deep-drawing process using Abaqus simulation software for a common part and experimental validation by a series of channel bending tests; (3) Determination of the relationship between different die profile shape and tool wear distribution for deep-drawing process; (4) Determination of the relationship between different control parameters (with the same type shape, e.g. elliptical, circular) and tool wear distribution for deep-drawing process; 151

170 (5) Development of a specialised algorithm for achieving minimised and uniform wear distribution by changing the die profile shape for deep-drawing process using Python programming language. 9.3 Recommendations for Future Work Future work can be focused on the tool wear prediction modelling which is closely linked with: Effect of lubricant; Effect of alternative die material; Effect of spring back. Due to the limitation of the computation time, the simulation in this project is 2D. In the future, with the high performance computer technology, the investigation, modelling and prediction conducted in this project can be conducted using 3D model, from which can be obtained the effects of the variation of geometries and control parameters along z axis on the tool wear distribution. The algorithm developed in Chapter 9 has a practical application for the optimisation of the die radius profile in the industries. In the future, the practical application of the algorithm in the industries can be studied in detail and a manufacture system can be designed using the algorithm. 152

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181 APPENDIX A LIST OF PUBLICATIONS Peer-reviewed Journal Papers: [1] X. Z. Wang and S. Masood, "Investigation of die radius arc profile on wear behaviour in sheet metal processing of advanced high strength steels," Materials & Design, vol. 32, pp , [2] X. Z. Wang and S. Masood, "A study on tool wear of sheet metal stamping die using numerical method," Materials Science Forum, vol , pp , [3] X. Z. Wang, S. H. Masood, and M. Dingle, "Numerical simulation and optimisation of sheet metal forming for auto-body panel using AutoForm software," Materials Science Forum, vol , pp , Peer-reviewed Conferences Papers: [4] X. Z. Wang, S. H. Masood, and M. Dingle, "An investigation on tool wear prediction in automotive sheet metal stamping die using numerical simulation," in the Proceedings of 2009 IAENG International Conference on Industrial Engineering, Hong Kong, 2009, pp [5] X. Z. Wang, S. H. Masood, and M. Dingle, "Tool wear prediction on sheet metal forming die of automotive part based on numerical simulation method," in the Proceedings of 5th Australasian Congress on Applied Mechanics, Brisbane, 2007, pp

182 APPENDIX B MOMENTARY PRESSURE CHART 164