The Pennsylvania State University. The Graduate School. College of Engineering A STUDY OF THE TURNING OF AUSTEMPERED DUCTILE IRON (ADI)

Transcription

1 The Pennsylvania State University The Graduate School College of Engineering A STUDY OF THE TURNING OF AUSTEMPERED DUCTILE IRON (ADI) GRADES WITH COATED CARBIDE TOOLS A Thesis in Industrial Engineering by Pei-Long,Ting 2016 Pei-Long,Ting Submitted in Partial Fulfillment of the Requirements for the Degree of Master of Science December 2016

2 ii The thesis of Pei-Long, Ting was reviewed and approved* by the following: Robert Voigt Professor of Industrial Engineering Thesis Advisor Edward De Meter Professor of Industrial and mechanical Engineering Janis Terpenny Professor of Industrial Engineering Head of the Department of Department or Graduate Program *Signatures are on file in the Graduate School

3 iii ABSTRACT Austempered Ductile Iron (ADI) is a relatively new material with the highest hardness and strength of any material in the cast iron family. Through the unique heat treatment - austempering, the ausferrite microstructure of ferrite and carbon-stabilized austenite along with graphite nodules is formed. Multiple strength levels can be produced by varying the austempering temperature or time. In general, ADI has a high strength-weight ratio, good toughness, and very high wear resistance compared to other ductile iron grades. In addition, the density of ADI is lower than steel but with approximately the same strength. These unique properties make ADI as an ideal material for manufacturing products requiring light weight but with high strength and toughness. On the other hand, ADI is difficult to machine because of its high hardness. This has impeded the application and the growth of market applications of ADI. The primary objective of this study was to evaluate the machinability of different grades of ADI (GR900, GR1050, GR1200) during high speed turning with coolant. Comprehensive turning experiments were conducted under a range of different machining conditions. The influence of cutting speed on tool life, surface roughness, and chip formation were analyzed during turning with coated carbide tools. The turning experiments were conducted on large diameter commercially produced, pre-machined cylinder castings at a constant feed rate of ipr and depth of cut of 0.06 inches. The cutting speed was varied for the different grades of ADI, from fpm and tool wear was measured at various time intervals. A Taylor tool life model was developed by measuring the tool life for a range of cutting speeds. This model was then used to generate general turning guidelines for the various grades of ADI based on tool life. Lastly, in order to benchmark the turning of ADI with other materials, turning studies with conventional Ductile Iron grade were also investigated under similar cutting conditions.

4 iv The chip formation for all grades of ADI and DI were discovered in the form of discontinuous c-shaped chips. As expected increasing cutting speeds accelerated the rate of tool wear. The surface roughness trend when machining GR900 and GR1050 are similar decreased cutting speed improved surface finish but a very low cutting speeds the surface finish of grade 1200 ADI also decreased.

5 v TABLE OF CONTENTS LIST OF FIGURES... vii LIST OF TABLES... x ACKNOWLEDGEMENTS... xiii Chapter 1 INTRODUCTION... 1 Background... 1 Problem statement... 1 Objectives... 2 Chapter 2 BACKGOUND... 3 Machining... 3 Cutting force... 7 Chip formation Turning Machinability Development of ADI Production of ADI Chapter 3 PREVIOUS RESEARCH Machining of ADI Cutting speed Cutting tool Depth of cut Chips formation and feed rate Chapter 4 RESEARCH PLAN Workpiece material characteristics Experimental platform Machinability metrics Chapter 5 RESULTS Chip formation Tool life Surface roughness Chapter 6 DISCUSSION Chapter 7 CONCLUSION AND FUTURE WORK References... 78

6 vi Appendix A - Tool wear measurements Appendix B - Surface roughness measurements Appendix C - Insert identification charts Appendix D Tool Wear Measurement Method

7 vii LIST OF FIGURES Figure 1: Illustration of a two-dimensional cutting process (orthogonal cutting) (a) with a well-defined shear plane (b) without a well-defined shear plane. [1]... 4 Figure 2: Oblique cutting. [2]... 5 Figure 3: The relationship of velocity Vs, V and Vc for orthogonal cutting. [4]... 7 Figure 4: Free body diagram of orthogonal cutting. [4]... 8 Figure 5: Merchant s circular force diagram. [4]... 9 Figure 6: Basic types of chips produced in metal cutting: (a) continuous chip with narrow, straight primary shear zone; (b) secondary shear zone at the tool-chip interface; (c) continuous chip with built-up edge; (d) segmented or nonhomogeneous chip; and (e) discontinuous chip. [1] Figure 7: Basic operations performed on a lathe. (a)facing (b)straight turning (c)taper turning (d)grooving and cutting off (e)threading (f)tracer turning (g)drilling (h)reaming (i) boring [3] Figure 8: Typical cutting conditions for common external turning operations. [6] Figure 9: Proper selection of inserts geometry for turning operations. [6] Figure 10: Crater wear of a turning tool. [1] Figure 11: Taylor tool life model (ln-ln coordinates) [3] Figure 12: Typical tool wear curves for different cutting velocities (V5 > V4 > V3 > V2 > V1) [3] Figure 13: Flank wear features for single-point-tool wear in turning operations [8] Figure 14: Surface finish representation: Arithmetical roughness [3] Figure 15: Austempered ductile iron heat treatment cycle Figure 16: The austempering reaction ADI processing window (a) High austempering temperature (b) Low austempering temperature Figure 17: The microstructure of each ADI grade Figure 18: SECO CNMG M5 TK2001 insert used in turning studies Figure 19: the SECO DCLNR2525M12-M tool holder fixed on the HAAS lathe turret Figure 20: The experimental ADI test cylinders prior to pre turning... 47

8 viii Figure 21: An ADI workpiece after initial surface removal prior to turning tests Figure 22: The Westover Portable Refractometer model RHB-32 used for cutting fluid characterization Figure 23: The procedure for measuring coolant concentration shown schematically Figure 24: Test cylinder placement for turning studies Figure 25: Nikon SMZ800 stereoscope and fixtures used for tool wear measurement Figure 26: Mahr Pocket Surf III profilometer and the calibration piece EPL-1691 Riser Plate used for surface roughness measurements Figure 29: Tool wear progressions for GR1200 at different cutting speeds, (Depth of cut = 0.06 in, feed rate = ipr) Figure 30: Tool wear progressions for GR1050 at different cutting speeds, (Depth of cut = 0.06 in, feed rate = ipr) Figure 31: Tool wear progressions for GR900 at different cutting speeds, (Depth of cut = 0.06 in, feed rate = ipr) Figure 32: Tool wear progressions for DI at different cutting speeds, (Depth of cut = 0.06 in, feed rate = ipr) Figure 33: Tool wear progressions for GR900, GR1050 and GR1200 at V = 375 fpm, (Depth of cut = 0.06 in, feed rate = ipr) Figure 34: Tool wear progressions for GR900, GR1050 and GR1200 at V = 500 fpm, (Depth of cut = 0.06 in, feed rate = ipr) Figure 35: Effect of cutting speed on tool life for each grade of ADI and DI (in terms of cutting time) (Depth of cut = 0.06 in, feed rate = ipr) Figure 36: Effect of cutting speed on tool life for each grade of ADI and DI (in terms of cutting length) (Depth of cut = 0.06 in, feed rate = ipr) Figure 37: Ln-Ln tool life plot for developing the Taylor tool life equation Figure 38: Effect of cutting length on surface roughness (Ra) for ADI GR900 at different cutting speeds (Depth of cut = 0.06 in, feed rate = ipr) Figure 39: Effect of cutting length on surface roughness (Ra) for ADI GR1050 at different cutting speeds (Depth of cut = 0.06 in, feed rate = ipr) Figure 40: Effect of cutting length on surface roughness (Ra) for ADI GR1200 at different cutting speeds (Depth of cut = 0.06 in, feed rate = ipr)

9 ix Figure 41: Effect of cutting length on surface roughness (Ra) for DI at different cutting speeds (Depth of cut = 0.06 in, feed rate = ipr) Figure 42: Effect of cutting speed on surface roughness (Ra) for all grades of ADI and DI (Depth of cut = 0.06 in, feed rate = ipr)

10 x LIST OF TABLES Table 1: Comparison of Brinell hardness ranges for the various ADI grades Table 2: Comparison of ISO, ASTM, SAE, and GB standard grades for ADI tensile strength (MPa) yield strength (MPa) elongation (%) Table 3: Chemical composition of ductile irons used in this study Table 4: Ferrite and austenite volume fraction of each ADI grade after heat treated as measured by X ray diffraction Table 5: Brinell hardness of each t material tested Table 6: Cutting parameters used in this study for each grade of ADI and for Table 7: Effect of cutting speed on chip form for ADI grades and DI (Depth of cut = 0.06 in, feed rate = ipr) Table 8: Tool wear progression polynomial fit equations and R2 values for different the grades of ADI and DI Table 9: Effect of cutting speed on tool life for each grade of ADI and DI in terms of cutting time and cutting length (Depth of cut = 0.06 in, feed rate = ipr) Table 10: Average surface roughness (Ra) for grades of ADI and DI (Depth of cut = 0.06 in, feed rate = ipr) Table 11: Flank wear for ADI GR900 at 375 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 12: Flank wear for ADI GR900 at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 13: Flank wear for ADI GR900 at 750 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 14: Flank wear for ADI GR900 at 1000 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 15: Flank wear for ADI GR1050 at 375 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 16: Flank wear for ADI GR1050 at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr)... 84

11 xi Table 17: Flank wear for ADI GR1050 at 750 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 18: Flank wear for ADI GR1050 at 1000 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 19: Flank wear for ADI GR1200 at 250 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 20: Flank wear for ADI GR1200 at 300 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 21: Flank wear for ADI GR1200 at 375 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 22: Flank wear for ADI GR1200 at 400 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 23: Flank wear for ADI GR1200 at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 24: Flank wear for DI at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 25: Flank wear for DI at 750 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 26: Flank wear for DI at 1000 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 27: Effect of cutting speed on tool life for ADI grades in terms of cutting time for all trials (Depth of cut = 0.06 in, feed rate = ipr) Table 28: Average surface roughness for ADI GR900 (Depth of cut = 0.06 in, feed rate = ipr) Table 29: Average surface roughness for ADI GR1050 (Depth of cut = 0.06 in, feed rate = ipr) Table 30: Average surface roughness for ADI GR1200 (Depth of cut = 0.06 in, feed rate = ipr) Table 31: Average surface roughness for DI (Depth of cut = 0.06 in, feed rate = ipr) Table 32: Surface roughness measurements for ADI GR900 at 375 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 33: Surface roughness measurements for ADI GR900 at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr)... 94

12 xii Table 34: Surface roughness measurements for ADI GR900 at 750 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 35: Surface roughness measurements for ADI GR900 at 1000 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 36: Surface roughness measurements for ADI GR1050 at 375 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 37: Surface roughness measurements for ADI GR1050 at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 38: Surface roughness measurements for ADI GR1050 at 750 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 39: Surface roughness measurements for ADI GR1050 at 1000 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 40: Surface roughness measurements for ADI GR1200 at 250 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 41: Surface roughness measurements for ADI GR1200 at 300 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 42: Surface roughness measurements for ADI GR1200 at 375 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 43: Surface roughness measurements for ADI GR1200 at 400 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 44: Surface roughness measurements for ADI GR1200 at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 45: Surface roughness measurements for DI at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 46: Surface roughness measurements for ADI GR1200 at 750 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 47: Surface roughness measurements for ADI GR1200 at 1000 fpm (Depth of cut = 0.06 in, feed rate = ipr) Table 48: Identification chart for inserts used to machine ADI GR Table 49: Identification chart for inserts used to machine ADI GR Table 50: Identification chart for inserts used to machine ADI GR Table 51: Identification chart for inserts used to machine DI

13 xiii ACKNOWLEDGEMENTS I would like to first acknowledge my appreciation for my adviser, Professor Voigt. His helpful comments on this paper and his patience over the past two years became the most crucial factor to complete this thesis. Also, I would like to thank and recognize the contributions of my senior Dika Handayani, support specialists Dan Supko and Travis Richner for their help in platform setup, experiment design, knowledge of manipulating the machine, and guidance in the analyzing methods used in this study In addition, I would like to express my thanks to Professor Edward De Meter for his assistance to review this study. This project was supported by Applied Process, Quaker Chemical Corp., and SECO Tools. My thanks also goes out to Kathy Hayryen, the advisor from Applied Process, for her suggestion, profession and guidance. Last but not least, I want to thank my parents, Sea-Chang, Ting and Wu Hsiang-Chun, Ting, who have supported me all the time and made this happen.

14 1 Chapter 1 INTRODUCTION Background Austempered ductile iron (ADI) is a relatively new material in cast iron family. In recent years, ADI has been much more widely used in industry. ADI use is driven by its special mechanical properties. Compared to ductile iron, ADI has improved strength, toughness and wear resistance that are developed during on austempering heat treatment. In addition, ADI has other advantages comparing to other competing materials. ADI components can often be produced at costs that are less than forged steel, with similar strength levels and lower density. Also, the strength, fatigue resistance, and wear resistance of ADI are higher than for aluminum alloys. With these advantages, ADI has quickly developing into an important automotive material as well as an important material in the railroad and defense industries for applications such as gears and ordinance. ADI casting can be processed by variety of machining process such as turning, milling, and drilling; however, its high hardness makes it difficult to machine. Even when rough machining is performed prior to heat treatment, final machining after heat treatment can be difficult. Problem statement In many cases, current and future applications of ADI are limited by difficulties in machining the material. There is incomplete information in the proper cutting speeds and feed rates when machining the various grades of ADI to achieve the desirable tool life and surface

15 2 finish to reduce machining cost. Therefore, manufacturers are often hesitant to completely machine castings after austempering treatment. In order to decrease the overall cost and increase the machining efficiency of the different grades of ADI, machinability studies must be conducted to develop effective machining guidelines. Objectives The primary object of this study is to characterize the turning behavior of common grades of ADI (GR900, GR1050, GR1200) under production turning conditions. Turning studies with contemporary coated carbide tools, coolant, and fixed feed rates and depths of cut were performed to develop the relationship between tool life and cutting speed. In addition, the effect of the different cutting speeds on surface finish and chip formation will also be evaluated.

16 3 Chapter 2 BACKGOUND Machining Machining is a cutting process where in the surface layer of a work piece will be removed by the cutting tool in chip form. [1] Figure 1 [1] shows this schematically. The unnecessary material is removed and slides along the rake face of the cutting tool with a new face being formed by the flank face. [1] Based on the geometric relationship of the tool to the workpiece, the cutting process can classify into two different type of process - orthogonal cutting and oblique cutting. [1] The orthogonal cutting is a fundamental two-force cutting geometry with the cutting edge oriented perpendicular to the cutting direction. The analysis of orthogonal cutting is relatively simple comparing to oblique cutting, because the cutting force can be modeled in only two dimensional. Because it is easier to build and evaluate cutting models, the orthogonal cutting is widely using in theoretical and experimental work. However, the majority of machining process are threedimensional oblique cutting. [1] [2] Unlike orthogonal cutting, in oblique cutting there is an inclination angle to the cutting edge. In other words, the cutting edge is not perpendicular to the cutting direction as showing in the Figure 2. [2]

17 4 Figure 1: Illustration of a two-dimensional cutting process (orthogonal cutting) (a) with a welldefined shear plane (b) without a well-defined shear plane. [1]

18 5 Figure 2: Oblique cutting. [2] Machining mechanics The mechanism of metal cutting is dominated by localized shear deformation in the workpiece at the cutting edge of the tool. Two different types of deformation could be classified during machining. [3] The primary deformation is the interaction between the cutting tool and the workpiece causing shear deformation and ultimate fracture which forms the chip. The secondary deformation is when chips passed over the rake face of the cutting tool resulting in additional deformation because of shearing and sliding of the chip against the tool. [3] These two deformations are mutually dependence. Because the material that rubs the rake face generates heat and plastic deformation during the primary process, the secondary deformation will be affected by the formation of a shear plane. On the other hand, the shear direction is affected by deformation and friction at the rake face which effects the heat and strain of the chips in the primary process. [3] During machining, the chips is formed by the interaction of the tool and the workpiece. The cutting process is shown schematically in Figure 1. [1] In this illustration, the tool is assumed to travel relative to the workpiece with a velocity V, the cutting speed. The tool

19 6 geometry is defined by the tool rake angle α and tool clearance angle γ. The shear process occurs in the shear plane with a shear angle ϕ and a shear velocityf s. Also, the uncut chip thickness, t 0 and the chip thickness which is assumed without deformation, t c both are shown in the figure. The chips have a velocity V c. Figure 2 shows that the V s is the vector velocity of V and V c. [4] The cutting ratio, or chip-thickness ratio r is important and useful for evaluating the cutting condition as a parameter, and is always greater than unity. [1] The cutting ratio is defined by relationships between the rake angle and the shear angle. Therefore the cutting ratio can also be presented in the form of velocity and chip thickness as well, equation 1 and 2: Tan ϕ = R Cosα / (1 - R Sinα) (equation 1) R = t 0 /t c = Sin ϕ / Cos (ϕ - α) = V c / V (equation 2) [4] Note that since the R is always greater than one, the chip velocity is greater than the cutting velocity. The velocities are also very important in estimating the power, temperature and vibration analysis associated with chatter during chip formation. Velocity relationship can be also be described with respect to the rake angle α and shear angle ϕ in Figure 3 [4]. The equation of the relationship of the three velocities can be expressed as: V / Cos (ϕ - α) = V s / Cos α = V c / Sin ϕ (equation 3) The shear strain, γ on the shear plane in cutting can also be expressed as:

20 7 γ = Cot ϕ + Tan (ϕ - α) (equation 4) The resultant strain rate γ r can be expressed as: γ r = V s / Y = [Cos α / Cos (ϕ - α)]* (V/ Y) (equation 5) Figure 3: The relationship of velocity V s, V and V c for orthogonal cutting. [4] Cutting force The cutting force is a critical factor because it starts the process of chip formation and controls both the flow of the chips and the thermal gradients involved in machining. [32] The interaction of chip formation and other machining variables has a significant influence on cutting force. Generally, the cutting forces are less for discontinuous chips than continuous chips. Also when cutting force decreases, the relative chip thickness also decreases because of an increase in

21 8 cutting ratio or shear angle. [2] For orthogonal cutting, the entire force system lies in a single plane as depicted in Figure 4.[4] Overall the cutting force could be separated into two basic forces - the friction force F and the normal force N acting on the tool-chip interface contact area. The resultant force R can also be resolved into two forces - shear force F s and normal force F n acting on the shear plane area As.[4] Figure 4: Free body diagram of orthogonal cutting. [4] However, neither of these two sets of forces can be measured directly. Merchant s circular force diagram (Figure 5) can be used to analyze the resultant force R. As shown in the Figure 5, the cutting force, R could be separated by cutting force F c and the tangential forcef t, which are measurable. The two forces F c and F t can then be used to calculate the resultant cutting force by

22 9 Equation 6: F= F c 2 + F t 2 (equation 6) Figure 5: Merchant s circular force diagram. [4] The relations between the measurable forces (F c, F t ) and the other components of the resultant force during orthogonal machining helps to calculate the forces in the primary and secondary deformation zones. [28] The friction forces F and the normal forces N is shown in Equation 7 and 8: F= (F c Sin α) + (F t Cos α) (equation 7) N = (F c Cos α) (F t Sin α) (equation 8)

23 10 The forces parallel and perpendicular to the shear plane can be described as: F s = (F c cos ϕ) - (F t sin ϕ) (equation 9) F n = (F c sin ϕ) + (F t cos ϕ) (equation 10) Where ϕ is the shear angle and α is the rake angle. The shear stress τ s can be defined by equation 11 and 12: τ s = F s /A s (equation 11) Where A s is the shear plane area, could be expressed as equation 12: A s = t 0 w / sin ϕ (equation 12) Note that t 0 was the uncut chip thickness and w was the width of the workpiece. The shear stress can also be expressed as: τ s = (F c Cos ϕ Sin ϕ -F t Sin ϕ 2 ) / t 0 w (equation 13)

24 11 Chip formation The surfaces of chip formed by machining are different in the constrained and unconstrained surfaces. The chip surface contacting with rake surface has a shiny appearance and is smooth. The unconstrained surface is rough. The basic nature of the chips forming during machining can be classified into four different types - continuous, build up edge (BUE), discontinuous, and serrated (segment). [1] Continuous chips usually form when machining a ductile material at high cutting speeds with high rake angle tools. [1] The deformation of the workpiece occurs in a narrow shear zone, which is called the primary deformation zone (Figure 6a) [1]. Continuous chips can potentially cause a second deformation zone due to the high friction at the rake surface. The continuous chip usually result in good surface finish of the workpiece. [1] However, continuous chip are typically not desirable. Continuous chip may tangle around the tool holder, workpiece and the disposal system. Tangling can result in stoppage time to clean the chip tangle. Tangling can be surmised by changing the cutting parameter such as cutting speed, feed, depth of cut or using cutting fluid. [1] The BUE chip is named because it has a layer of the material from the workpiece which is progressively deposited on the tip of the cutting tool as shown in Figure 6c. [1] When the BUE chip becomes larger, it will be unstable and break off the tool. Part of the BUE is carried away from the tool side of the chip and randomly deposited on the workpiece surface. This cycle of BUE formation and destruction will repeat continuously during machining. [1] The BUE effectively changes the geometry of the cutting edge and makes it dull. Therefore, it has a negative effect on the surface finish. However, if the BUE is stable, the tool wear will actually decrease by protecting the rake face of the tool from wear. The tendency of BUE formation is generally decreased by increasing cutting speed, rake angle and decreasing the depth of cut.

25 12 Using the sharp cutting tool and proper cutting fluid also decreases the tendency of BUE formation. [1] A discontinuous chip, shown in Figure 6e, is defined as chip segments connected firmly or loosely to each other. [1] Discontinuous chip formation during the machining is because of a series of discrete chip segments that fracture as stress concentrates in the workpiece. [2] Discontinuous chip formation usually takes place when machining a material with inclusions, or points of stress concentration such as cast iron due to the brittle graphite phase that doesn t have the capacity to undergo the high shear strain involved in cutting. [1] [2] Certain cutting conditions increase the tendency of forming discontinuous chips. The conditions of very low or very high cutting speed, large depth of cut, low rake angles and lack of proper cutting fluid will all increase the tendency of discontinuous chip formation. [1] Sometimes a discontinuous chip has a semi-continuous appearance such as in Figure 6d which is called a serrated (segmented, nonhomogeneous) chip. The serrated chip has large zones of low shear strain and small zones of high shear strain. (Shear localization). The serrated chip usually occurs when machining low thermal conductivity and low strength materials decreasing significantly with temperature (thermal softening) such as titanium, resulting in a sawtooth like chip during machining. [1]

26 13 Figure 6: Basic types of chips produced in metal cutting: (a) continuous chip with narrow, straight primary shear zone; (b) secondary shear zone at the tool-chip interface; (c) continuous chip with built-up edge; (d) segmented or nonhomogeneous chip; and (e) discontinuous chip. [1] Due to the structure of discontinuous and serrated chips, the cutting force will oscillate when producing those types of chips. [1] Consequently, the stiffness or rigidity of the cutting equipment such as the tool holder, the workpiece holding device, and the cutting tool are important. If the stiffness of the equipment is not sufficient, the cutting tool will vibrate and chatter, leading to poor surface finish, poor dimensional accuracy, and premature tool wear. [1]

27 14 Turning Turning is a widely used material removal process generally rotating the workpiece on a lathe and feeding the cutting tool mounted on the lathe into the workpiece to remove the unnecessary part of workpiece to the required shape. [2] All turning operations can be divided into two categories - the workpiece is situated in the center between chucks or fixed at the one end without support at the other end. [3] The workpiece is held and rotated as the cutting tool moves with a certain depth of cut travel in one direction under the certain velocity. Other than the depth of cut and cutting speed (velocity), feed rate is also an important cutting parameter of turning. The feed rate is the distance of tool travels horizontally or vertically per unit revolution of the workpiece. [2] Figure 7 showing the common operations performing that could be achieved by turning on lathe with different kinds of cutting equipment. [3] The common machining process could be done by lathe are: 1. Turning: to produce straight, conical, curved or grooved of workpiece such as shafts, spindle, and pins. 2. Facing: to produce a flat surface which is parallel to the rotating direction at the end of the workpiece for the parts needed to be assembled with other components. 3. Boring: to enlarge a hole or cylindrical cavity made by previous process or to produce circular internal grooves. 4. Drilling: to produce a hole or improve the dimensional accuracy and surface finish of boring hole. 5. Parting: (as cut off); the process is to cut a piece from the end of the workpiece.

28 15 Figure 7: Basic operations performed on a lathe. (a)facing (b)straight turning (c)taper turning (d)grooving and cutting off (e)threading (f)tracer turning (g)drilling (h)reaming (i) boring [3] Cutting tool geometry Most turning is done as single-point turning. [1] [4] Single point of turning could be further classified into rough and finish turning based on the different feed rate and depth of cut. [1] [6] Finish turning is after initial roughing cuts improve dimensional accuracy and improves surface finish. [1] [4] The tool geometry is important to effectively control chip formation. [6] Rake angle controls the chip flow direction and the tool tip strength. In general, positive rake angles reduce cutting force and temperature. However, positive rake angles have the risk of

29 16 premature tool chipping and failure. For brittle, high strength materials, the rake angle should be negative. Increasing the wedge angle (the angle between the top face and the clearance angle) will create a stronger cutting edge but require greater power to machine. [6] The clearance angle is also important. If the clearance angle is too large, chipping is more likely. However, if it is too small, the flank wear would increase and the finish surface will be adversely affected. Another important factor of tool geometry is nose radius. Nose radius can significantly affect the surface finish and tool tip strength. Smaller nose radius will reduce surface finish at high feed rates. [1] When roughing the workpiece, the nose radi should be larger to increase the strength of the tool point, Figure 9. [6] In addition, the machining with bigger nose radi is more desirable at higher feed rates. [6] In general, it is recommended that the feed rate for roughing should be 50% of the size of the nose radius in order to have expected surface finish. [6] Figure 8: Typical cutting conditions for common external turning operations. [6] Figure 9: Proper selection of inserts geometry for turning operations. [6]

30 17 Cutting parameters and material removal rates The key cutting parameter in turning are cutting speed, feed rate and depth of cut. [4] The surface cutting speed is built on the RPM of the rotating workpiece and diameter of the workpiece V max = π D 0 N s (equation 14) Where D 0 is the original cutting diameter (mm/in) during turning process, N s is the rotation speed of the workpiece. (rpm) Or V avg = π D avg N s (equation 15) Where D avg is the average cutting diameter (mm/in) during turning process, the D avg could be calculated by equation 16: D avg = (D 0 +D f ) / 2 (equation 16) Where D f (mm/in) is the final diameter. The feed rate, the distance of tool travels per unit revolution (F r ) is given by equation 17:

31 18 d = DOC = (d 1 +d 2 ) / 2 (mm/in) (equation 17) Note that d 1 is the original diameter before cutting and d 2 is the new diameter after machining. The length of cut is not only the distance traveled parallel to the perpendicular axis of machining the workpiece but also the sum of allowance or overrun distance before or after machining. [4] The cutting time (T), can be calculate by equation 18: T = (L+A) / F r N s (equation 18) Where Ns is the rotation speed of workpiece (RPM), A is the distance before or after machining. [4] The material removal rate (MRR) is the volume of material removed per unit time (mm 3 or inch 3 / min), as follows: MMR = volume removed / time = π [(d 1 /2) 2 - (d 2 /2) 2 ] F r N s (equation 19) If the average MRR is required, the average diameter is used instead of the single diameter of a particular cut. The average MRR can be described as: MRR = π D avg d F r N s or d F r V (equation 20)

32 19 Cutting fluid The cutting fluid is also a very important during turning. The basic two functions of cutting fluid are cooling and lubrication which decreasing the friction force and temperature of the interface between tool and workpiece. [3] At moderate cutting the lubrication and cooling are both important, and for low speed cutting, the function of cutting fluid is mainly for lubrication. [3] Compared to dry cutting, use of a cutting fluid can increase the productivity and reduce the cost by making possible of using higher the depth of cut, cutting speed, and feed rate. The effective application of cutting fluid can increase the tool life, dimensional accuracy, as well as decrease the surface roughness and power consumption. The choice and effectiveness of cutting fluid can be determined by the condition of the machine tool, cutting parameters (speed, feed rate, depth of cut), composition, finish, geometry of the cutting tool, mode of fluid application, geometry of material is machined, surface coatings, and the composition microstructure, and residual stress distribution of the workpiece. [3] Machinability The machinability of a material is not only based on the metallurgical properties such as hardness, microstructure, and chemical composition but also on compatibility with the machining process including the shape of cut, the cutting tool and the cutting fluid. Different machining environments can have a significant effect on machinability rankings. [2] Furthermore, no single dependent variable in machining is universally accepted as a measure of machinability. As a result, machinability ratings for a material will vary depending on the different methods used to evaluate machinability such as surface finish, tool life, cutting force, and chip formation.

33 20 Chip formation The shape of chip formation, can be an important indicator of machinability. The formation of short chips are considered as a better result than long, unbroken or small, powder like chips. [28] A ductile material that forms long, unbroken, or continuous chip has more tendency to burrs, especially as tools wear. [5] The burr formation are often used to evaluate the machinability of soft, ductile materials such as aluminum. In addition, comparing chip type (the continuous chip or discontinuous chip) indicates ease of chip recycling and tendency to clog the machine. [26] Cutting force The cutting force is another method used to evaluate machinability. Various cutting parameters have a significant influence on cutting force. Typically the feed rate, depth of cut has more of an effect on cutting forces than cutting speed. [5] The cutting forces increase significantly with flank wear. In general, low cutting force result in the longer tool life and improved dimensional accuracy. [5] The tool geometry also affect the cutting force. The cutting force is reduced when the rake angle increases. [5] In general, the use of force as an indicator of relative machinability is based on the assumption that more difficult to machine metals require more cutting force. [2] However, the tool force is not typically used as the sole measurement of machining quality. [2] [7] Tool life Tool life studies are the most direct method to evaluate the machinability of materials. When tool life increases, the effective machinability also increases. [28] However, tool wear can occur by various mechanism: abrasion, adhesion, diffusion, chemical wear, and oxidation wear. [2] Abrasion wear takes place when hard material and hard particles on the surface of the chip slide on the tool faces and remove tool material. The hard particles could be fragments of the BUE or

34 21 abrasive inclusions within the workpiece material. [2] Adhesion wear occurs when the frictional environment at the interface between the tool and chip, caused by pressure and high temperature, causes welding at the interface of chips and tool face. When the welds fracture, a small part of tool material is carried away with the chips. [2] Diffusion wear occurs when the atoms of a metallic crystal lattice diffuse from an area of high atomic concentration in the workpiece to an area of low concentration in the tool, due to the high temperature at the interface between the tool and the chip on the workpiece. Chemical wear is a chemical reaction that wears the tool by corrosion. The oxidation wear occurs at very high temperature at the point at which the structure of the cutting tool is weakened through oxidation reactions. The high temperature softens the microstructure of cutting tool and weakens the edge of the cutting tool. [2] The tool wear can be classified into many different types, such as flank wear, crater wear, nose wear, notching, plastic deformation of the tool tip, chipping and gross fracture. [1] Among the various type of wear, flank wear and crater wear are commonly used to characterize tool wear in most studies, Figure 10 [1] shows characteristic of crater wear and the flank wear on a turning tool. [1] [3] Crater wear is observed on the rake face of the cutting tool and flank wear is occurred on relief face of the cutting tool. Crater wear is a result of both of abrasion and diffusion [6] and is believed to be a thermochemical reaction. [2] The evaluation of crater wear can be done by measuring the location of the maximum depth of the crater, KT, Figure 10. KT is also coincides with the location of the maximum temperature at the tool-chip interface. [1] Crater wear is normally generated by diffusion mechanisms. With the increased tool temperature, the crater wear increases as well. [1] Crater wear is normally used for evaluating tool wear in extreme conditions. [28] However, flank wear is more typically used to evaluate tool wear. [40] Flank wear generation is caused by the rubbing of the tool along the machined surface causing adhesive and/or abrasive wear. [1] Various tool life models have been used to predict the tool life

35 22 based on tool wear. One of the earliest successful tool life model was developed by F.W Taylor in In this model, the relationship between cutting velocity and tool life is used and to optimize the tool performance during machining. The Taylor s tool life equation is generally expressed as: [3] VT n = C (equation 21) Where T is the tool life, V is cutting speed, and C and n are material constants. When time to develop a given amount of flank wear is fixed, recommended cutting speeds can be estimated. [1] The constants n and C are obtained by testing cutting tools at different cutting conditions and using the tool life criterion to establish a point at which the defined end of tool life is achieved. [3] The exponent n primarily depends on the tool material and other factors such as workpiece material and cutting conditions. [1] The n value range for common tool materials are: for high-speed steels, for carbides, for coated carbides, and for ceramic cutting tool. [1] The constant C is numerically equal to the value of the cutting speed when T is equal to one minute. It is widely used as a measure of machinability for particular cutting tools and cutting conditions. [3] Figure 10: Crater wear of a turning tool. [1]

36 23 Figure 12 [3] shows typical tool wear curves at different cutting velocities. The shape of the curve can be broken into three sections - initial wear, where the sharp edge wears quickly, the steadystate wear in the middle, where wear progresses at a constant rate, and accelerated wear, where the tool wear rapidly increases just before total failure. Flank wear (W1, Figure 12) is typically the criteria for limiting tool wear. The width of flank wear land is an appropriate wear criteria for single point cutting. The wear land can be classified into three zones. The first zone, C, is located at the tool corner. The second and third, zones B and A, are on the cutting faces, which stop at the zone N. In the zone B, the wear land is relatively uniform. Two different wear land width can be identified, the maximum width of flank wear VB B max and the average wear land width VB. Either criteria can be considered identified as critical indicators of tool wear. [8] When the wear land is uniform, the VB criteria is used for measuring the tool life. If the uniform wear land width reaches the value of 0.3mm, it is considered as tool failure. When the wear land is not uniform, VB B max is considered. When the VB B max reaches 0.6mm, the tool is considered as reaching the end of its life. [8] Figure 11: Taylor tool life model (ln-ln coordinates) [3]

37 24 Figure 12: Typical tool wear curves for different cutting velocities (V 5 > V 4 > V 3 > V 2 > V 1 ) [3] Figure 13: Flank wear features for single-point-tool wear in turning operations [8]

38 25 Surface finish Surface finish plays an important role in influencing the dimensional accuracy, properties, and performance of machined parts. [1] The surface finish is typically characterized by measuring surface roughness. The surface roughness has many different designations, including the arithmetical roughness R a, root mean square (rms) roughness R q, maximum peak to valley roughness height, R y or R max, ten-points height, R z. [3] Among the various designations, the arithmetical roughness Ra is the most common method of designating surface roughness in the US. [3] The R a is the average of absolute deviation of the workpiece surface from the centerline as shown in the Figure 14 [3] and can be expressed as: R a = 1 L L 0 y dx (equation 24) Where L is length of the measurement and y is the ordinate of profile from the centerline. Figure 14: Surface finish representation: Arithmetical roughness [3]

39 26 Surface roughness is a function of tool feed and the geometry and is also influenced by BUE, chatter and inaccuracies in machine tool movement. [9] [10] The roughness value is also influenced by the feed rate and tool nose radius. The equation of ideal surface roughness is shown in equation 25 [3]: R a = f 2 / r (equation 25) The theoretical surface roughness can be estimated from feed rate, tool radius, and cutting edge angle for cases when the feed rate does not exceed the theoretical (nose radius) limit. If feed rates exceed the limit, the resulting scallop is not a strict function of the nose radius but also depends on the end cutting edge angle. The actual surface using a given tool could be worse than estimated because of BUE generation during the machining process. [3] However, in other situations, the actual surface roughness is also less than the theoretical surface roughness because of tool wear on the cutting edge that produces a finer finished surface. The wear land of the tool provides the wiping action that tend to smooth out the surface irregularities. [3] Good surface finish is an indicator of machinability. With its significant influence on tool-tip, the BUE profile can also have a great effect on the surface finished. [1] Therefore, ceramic and diamond tools generate better surface finishes than other tools due to the in lower tendency to form a BUE. [1] The tool tip radius is typically large in relation to the depth of cut. Otherwise the tool will rub over the surface and generate heat and residual surface stress sometimes causing tearing and cracking. As a result, depth of cut should be larger than the radius of cutting edge. In addition,

40 vibration and chatter during machining can also adversely affect surface finish because vibration and chatter of the tool will periodically change the dimension of the cut. [1] 27 Development of ADI ADI develops an ausferrite matrix structure after heat treatment, a mixed microstructure of acicular ferrite and carbon enriched stabled austenite. The austempering process was first developed by 2 researchers working of United States Steel Laboratories, Davenport and Bain, in the early 1930 s for heat treating steels. [22] Much later, the unique ausferrite banite microstructure of cast iron was discovered. [24] However, commercialization of austempered cast irons has lagged. [23] In 1948, ductile iron (DI) was first developed by British Cast Iron Research Association. Ductile iron is a unique material with spherical shaped graphite structure. DI has high strength, has higher elastic modulus and substantial ductility. Soon after DI was developed, International Harvester and General Motors started to use the austempering process on DI in research studies in the 1960 s.[26] However, ADI was not produced commercially until the early 1970 s. In 1972, the Tecumseh Products announced the first commercial ADI product a compressor crankshaft. [27] Since then, more and more companies have started to use ADI. The Finnish company KymiKymmene Metal used ADI to replace forged steel for a set of gears in the early 1970 s. In late 1970 s, General Motors replaced a wide range of steel gears, pinions, and velocity joints in light vehicles with ADI. [27] Since the 1980 s ADI applications have been continuous increasing in North America. The Cummins Engine started to make ADI timing gears produced to AGMA class 8 standards, in their B and C series diesel engines in [27] The main reason for the market expansion is cost

41 28 reduction. The price per unit of mass for ductile iron was typically half that of steel. When the extra cost of austempering process was added, the ADI products are still much more economical than steel. [26] Because of the increasing application of ADI in 1980 s, ADI material specifications were developed. In 1990, the first US standards - ASTM 897/897M for ADI were released. [26] In this standard various grades of ADI are specified for many applications. Updated ADI standards were released in 2000 s. Currently, there are four common standards - ASTM (A897/A879M) first issued in 2002, SAE (J2477) issued in 2003, ISO (17804) issued in 2005, and China standard GB/T24733 issued in The Table 1 shows the comparison of the hardness of the different standard grades. Table 2 shows the yield strength (MPa) - elongation (%) for the various grades. Table 1: Comparison of Brinell hardness ranges for the various ADI grades Hardness Range (HB) Grades ISO SAE J2477 ASTM A897/A897M *Based on tensile strength (MPA)

42 29 Table 2: Comparison of ISO, ASTM, SAE, and GB standard grades for ADI tensile strength (MPa) yield strength (MPa) elongation (%) ISO Issued 2005 SAE J2477 Issued 2003 Revised 2004 ASTM A897/A879M 2006 GB/T24733 Issued In 2003, the annual worldwide production of ADI was estimated as 125,000 tons and it was predicted to exceed 300,000 tons by However, the annual world production of ADI has grown at a faster pace. In the 2009, the estimated annual world production of ADI had already reached 300,000 tons in the end of [25] Today, ADI is widely used for agricultural equipment, construction equipment and gear or powertrain components. Also heavy truck or trailers, light vehicles and buses, mining or forestry equipment, railway, equipment, farm and oilfield machinery, conveyor and tooling equipment, defense, energy generation and sporting goods all have ADI components. Without a doubt, the production and application of ADI will continue to grow in future. [28] Production of ADI The production of ADI is a two-stage isothermal heat treatment process austenitization followed by austempering. During the austenitizing stage, the matrix microstructure of ductile iron will be transformed into austenite followed by rapid cooling to the austempering temperature. In the austempering process stage, the reaction is transformation of austenite (γ) to form ferrite (α) and carbon- enriched stabilised austenite (γ s ). Excessive holding at the

43 30 austempering temperature will result in the carbon- enriched stabilised austenite decomposing to ferrite and carbide (conventional bainite). The entire production process could be separated into five stage, shown in Figure 15 [14]. In the first austenitization stage (section A-B and B-C in Figure 15) [14], the casting is heated to an austenitizing temperature of between 1550 to 1750 for sufficient time (one to three hours) in order to completely austenitize the microstructure. Figure 15: Austempered ductile iron heat treatment cycle The next step is the quenching stage, the section of C-D in Figure 15 [14]. The temperature of casting must be dropped to 450 to 750 in the austempering temperature range. In this stage, the quenching rate is very important. The quenching rate must be high as possible to reach the austempering temperature without transforming any of the austenite to pearlite.

44 31 After reaching to the temperature between 450 and 750. The components held for a period of time (D-E in figure X). This is known as the austempering time. The holding time in this stage is very crucial due to the heat treatment processing window, showing in the Figure 16, is defined as the gap of time period between the first reaction is completed ( γ α + γ s+c ) and the before the second reaction starts, ( γ α + Fe 3 C). During the austempering stage, because of the silicon content of ductile iron, the conventional bainite reaction that occurs in is suppressed, causing the carbon rejected by bainite to ferrite enrich the carbon content of remaining austenite until the austenite matrix is stabilized. [20] To achieve desirable properties, the austempering time should stop before the second stage starts. However, if the austempering time is too short, the higher carbon content austenite is not fully stabilized and will transform into martensite upon cooling. The bainite and martensite structures, which should be avoided, would both affect the mechanical properties of ADI. The last stage (E to F, in Figure 15) is cooling to room temperature, resulting in a matrix of ausferrite, consisting of ferrite needles in carbon- enriched, stabilized austenite. Figure 16: The austempering reaction ADI processing window (a) High austempering temperature (b) Low austempering temperature

45 32 The three main factors that affect success are austempering time and temperature, austenitizing time and temperature and a cooling rate sufficient for the casting/alloy combination. [27] The Austenitizing time and temperature are factors that cannot be ignored when producing ADI. The austenitizing temperature is related to dissolved carbon, and the reduction in the rate of austempering. [28] The purpose of the austenitizing is to saturate the austenite matrix with carbon. Austenitizing temperature affects the final carbon content of the austenite. Increasing the austenitizing temperature causes an increase in the carbon content of austenite, leading to higher hardness and strength after austempering, [27] but slowing the transformation during austempering and sometimes reducing the mechanical properties after austempering. High carbon content austenite needs more time to transform to ausferrite [27]. During the austempering process, decreasing the austempering temperature will delay the austempering transformation start and completion times. At lower temperature the austenite structure may not complete stabilize. [28] On the other hand, lower temperature also create lower austenite contents with fine ferrite needles with some bainitic ausferrite formation. The higher austempering temperature leads to a coarser ausferrite microstructure, which consist of coarse and distinct bainitic ferrite needles, with more ductility but less strength and hardness than ADI s transformed at lower austempering temperatures [29]. The proper austenitizing temperature is dependent on the chemical composition of DI. Elements such as manganese, silicon and molybdenum play an important role in selecting the proper austenitizing temperature. [18] The upper critical austenitizing temperature (UCT) is the lowest temperature for forming austenite + graphite (γ + G) as described on the Fe C phase diagram. The manganese will decrease the UTC while the silicon and molybdenum will raise it. [18]. However, if the austenitizing is performed below the UTC, proeutectoid ferrite will be present in

46 33 the final microstructure, resulting in a lower strength and hardness material after austempering. [20] The time of austenitization is also very important. Insufficient holding time would result in an austenite matrix that is not fully saturated with carbon. The time is affected by composition; heavily alloyed irons will take more time to austenitize. [18] Han claims that an austenitization temperature at 900 for 1.5 hours will increase the size of austempering process window and will shorten the time for decomposition of austenite. [30] The austempering temperature and time also have a significant effect on ADI performance. These two factors will influence the mechanical properties, because they affect the composition of the micro constituents. Figure 16 [29] shows the influence of temperature on microstructure. At low austempering temperatures, the carbon diffusion rate is low, resulting in a limited amount of austenite carbon enrichment in stage 1 and accompanied by the co-precipitation of carbides in the ferrite. On the other hand, high austempering temperatures will increase the diffusion rate of carbon. Therefore, these microstructure would have more carbon - enriched stabilised austenite. The matrix of ADI can have up to 50% of carbon - enriched stabilised austenite. [29] The austempering temperature is normally controlled between 450 and 750. If the temperature dropping too low during the quenching stage, it will generate martensite. Martensite will adversely affect the mechanical properties of ADI. Therefore the austempering temperature must be controlled above the martensite start temperature. On the other hand, if austempering temperature is too high or the quenching process is slow, pearlite will form instead of ausferrite. [16] Due to the different proportion of austenite and ferrite in microstructure at different austempering temperatures, the mechanical properties vary as austempering temperature changes. At high

47 austempering temperature larger proportions of stabilized austenite are produced and the hardness and strength are lower but the ductility is improved. [11] [12] [13] [29] [31] 34 The austempering time is another critical factor to affect the production of ADI. Insufficient austempering holding time causes insufficient carbon stabilisation of austenite and transforms the microstructure to martensite during subsequent cooling to room temperature. The formation of martensite will increase the strength and hardness, but will significantly decrease the ductility resulting in brittle performance. [14] [28] [29] [32] Also, long austempering time should be avoided. Excessive austempering times will lead to decomposition of stabilised austenite to form ferrite and carbide [29] [33] and decrease the ductility and toughness [29], but not significantly affect the hardness and strength. [11] [15] [33] The desired ADI matrix should be thought of as a metastable or intermediate microstructure of ferrite and stabilized austenite with a temperature time austempering processing window defining successful heat treatment. [29] The austempering stage 1 (within the processing window) and undesirable Stage 2 reactions (over - austempering) are shown in Figure 16. Stage 1: γ (Austenite) α (Ferrite) + γ s (carbon enriched stabilised austenite) Stage 2: γ s (carbon enriched stabilised austenite) α (Ferrite) + ε (Carbide) Alloy elements influence the size of the heat treatment processing window for ADI and thus can have a significant influence on the microstructure and mechanical properties. [31] Alloying element in ADI help to prevent pearlite and ferrite formation upon cooling from the austenitizing temperature and make the ADI production easier to control in the processing widow by shifting it

48 35 to the right and delaying carbide precipitation during the bainitic transformation. [21] However, an improper alloy composition would cause significant alloy segregation, which will affect the microstructure and mechanical properties adversely. Ni, Si, Mn, Cu, Mo are the common alloying elements that have significant effect on ADI. [30] Nickel (Ni) tends to prolong the decomposition of the carbon with austenite during the austempering process. The use Ni is limited because of its high cost. The roles of Manganese (Mn) and Molybdenum (Mo) are similar. They are used for preventing pearlite during the austempering process and stabilised the austenite, making the carbon difficult to decompose. However, Mn and Mo can also cause the serious segregation and formation of martensite, which will limit mechanical properties. Since Mo segregates more than Mn, the amount of Mo should be controlled carefully. The amount of Mo is normally limited to between wt%. Copper (Cu) can also be used for slowing the second stage of austempering process, creating a wider process window with only a small influence the mechanical properties. Silicon (Si) is an important element for production of DI as well as for controlling the austempering reaction. The main function of Si during austempering process is to stabilize the austenite and decrease the carbide formation at the second stage. The amount of Si normally limited between 2.0% and 2.5% for effective DI production. [30]

49 36 Chapter 3 PREVIOUS RESEARCH Machining of ADI The machinability of a metal is very important in industry. During the past thirty years, many ADI machining studies have published. Most of these studies have focused on the lower strength grades of ADI. However, the growth of ADI markets are still impeded due to insufficient information on the machining of ADI. There are three different strategies for machining ADI. [14] [35] [36] The three strategies are: 1. Machining before heat treatment 2. Rough machining prior to heat treatment followed by final machining after heat treatment. 3. Machining after heat treatment. The first method is primary to circumvent the difficulty of machining ADI. The process can be widely used depending on the proper design and predictable part growth during the austempering process of the component. Normally, this method is using for wide dimensional tolerance parts. The second method is using when tight tolerance and surface finishes can t held during heat treatment. One of the advantage for using this method is improvement of final part fatigue strength. However, the process is costly and can cause logistical problems. [14] Machining after heat treatment avoids the logistical problems and cost of machining twice while maintain tight tolerances and good surface finish. [14] The two main difficulties when machining ADI are the high hardness of ADI and the straininduced transformation from the austenite structure to the harder and more brittle martensite on

50 37 the surface of the ADI component. [14] [36] Strain-induced transformation of ADI is because the austenite structure will transform to martensite structure under plastic strain directly beneath and ahead of the cutting tool. The strain induced transformation of ADI severely resists further fine machining, leading to overall poor tool life and tool failure. [41] From the view of material removal rates, ADI is not as easily machined as pearlitic or ferritic ductile irons, but does machine comparably to 30Rc hardened steel [14]. The recommended material removal rates of ADI are reported to be 75% of that of pearlitic ductile irons [36]. Also, the recommended setting machining parameters for machining ADI are 50% less speed and 50% deeper depth of cut when compared to other materials with similar hardnesses. [14] [36] The effect of different cutting parameters on the machinability is one of the main targets of this ADI turning research. Turning tool life depends on speed, feed rate, cutting tool material and depth of cut. The choice of cutting tool is also a critical factor affecting machinability. [37] [38] Cutting speed The cutting speed has a significant influence on turning tool life. The evaluation of turning tool life is typically based the tool wear. [39] Flank wear is typically used to evaluate tool wear instead of crater wear because the flank wear during turning directly affects the accuracy of the product. [40] As cutting speed increases, the tool life will decrease significantly. [31] [32] [33] [34] [35] The cutting speed will also influence chip formation. Austenite on ADI surface transforms to martensite due to the associated high surface strains. The brittle, hard martensite that forms accelerates tool wear and forms discontinuous chips. [44]. The surface roughness will also be influenced by cutting speed because high cutting speeds result in higher temperatures on the surface of tool and at tool-chip interface. [41] [42] [45] Ucun and Gök used CBN cutting tools at various cutting speeds (200, 300,400 m/min) to machine ADI austempered at different

51 38 austempering temperatures (250, 350 ). The result shows that the higher cutting speed and associated high strain rates, lead to the formation of martensite and discontinuous, smaller chips and shorten tool life. [44] In research by Aslantas and Ucun [45], two type of cutting tools (nose radius 0.4 mm) - ceramic (Al2O3 based) and cermet (TiCN+TiN coated), were compared under various cutting speeds, constant depth of cut and feed rate when machining two different grades of ADI. The results show that from the view of tool life, ceramic tools performed better than cermet tools. Ceramic cutting tools were not suitable for machining ADI at low cutting speeds based on the criteria of tool life and surface roughness. However both ceramic and cermet tools were suitable at high cutting speeds. [45] Cutting tool Choice of the proper cutting tool is an important factor for cost effective machining. The high strength and ductility of ADI adversely affects tool wear during machining. The cutting tools for machining ADI must have high wear resistance. Tool geometry is also important. The tool life has also been observed to increase 70 % and by 100% when dry turning ADI and when wet turning with prefered tool geometries. [46] Tool nose radius also impacts on surface roughness. The research shows that increasing the nose radius impacts surface roughness. [47] Carbide cutting tools is typically recommended for machining the various ADI grades, with K-grade carbide tools used for turning with cutting fluid and P-grade cutting tools used for dry cutting. [14] Carbide tools with various coatings have good tool life when machining all grades of ADI under a wide range of cutting conditions. [28] Other material such as ceramic and CBN have also been used for cutting ADI. Al2O3 ceramics can be used for continuous cutting processes such as turning. However, Si3N4 ceramics and PCBN cutting tools are not suggested for discontinuous cut machining of ADI. [14]

52 39 Depth of cut Depth of cut is a very important variable for setting cutting parameters. Decreasing the depth of cut tends to reduce tool life for ADI grades at constant material removal rates [35]. Avishan, Yazdani, and Vahid studied relationship between depth of cut and machinability. Decreasing depth of cut in general reduced tool life. They recommend depths of cut between 0.5 and 1 mm for machining ADI. [35] Akdemir,Yazman, Saglam and Uyaner also reported the influence of depth of cut when machining ADI. When the depth of cut increased, the surface roughness also increased. However, they found that reduced depths of cut did not significantly affect tool wear. [42] Chips formation and feed rate The chips formed during the machining ADI are discontinuous. [14] [44] [47] Compared to the continuous coil - like chips that typically result when machining other materials such as steel, the discontinuous chips formed when machining ADI are considered beneficial. Discontinuous chips are easier to handle and recycle and, in highly automated machining centers, the smaller chips tend to not to clog the equipment. [14] Chip shape is influenced by feed rate and cutting speed, but is typically not affected by depth of cut. When feed rate increases, the thickness of chips also increases. [47] This can change the chip shape from a coiled chip to C - shaped chip as feed rate increases. When the cutting speed is increased, the chip shape turned from crack-like and C - shaped to coiled chips. [48] The feed rate would also affect the tool life and surface finish of ADI. In research by Polishetty, tool wear increases when the feed rate increases, especially when cutting with PCBN tools. It was observed that feed rate also demonstrated a dominant effect on the surface roughness. [41]

53 40 Polishetty machined ASTM grade 3 ADI using PCBN and ceramic cutting tools for both rough turning and fine turning. [41] Katuku, Koursaris, and Sigalas researched the machinability of ASTM grade 2 ADI, using PCBN cutting tools with the constant depth of cut and feed rate at different cutting speeds. These investigation showed that the optimized cutting speed was between 150 and 500 m/min for dry turning. C. Wang, Guo, W. Wang and Dong studied tool life during turning with CC650 ceramic tools at various cutting speeds, feed rates and depths of cut. [48] Carvalho, Montenegro, Gomes investigated the machinability of ASTM grade 2 and 3 using carbide tools at a constant cutting speed with various depths of cut, feed rates and tool nose radi. [47] Akdemir, Yazman, Saglam and Uyaner investigated the effect of machinability by using carbide inserts at a constant feed rate. [42] In all of these studies, optimal turning conditions were reported at various cutting speeds and depths of cut. Parhad, Likhite, Bhatt, Peshwe evaluated the tool life of TiAlN-coated tungsten carbide inserts, at a constant feed rate at two levels of depth of cut (1, 2 mm) and three levels of speed when turning ADI-250. In the term of surface quality, the combination of a depth of cut of 2 mm and feed rate of 0.1 mm/rev resulted in the best surface quality when the cutting speeds is in the range m/min. [49] However, in total, past studies have not resulted in a comprehensive understanding of the machinability of ADI grades. Although some studies have focused on were conducted in particular grade of ADI, together these studies were comprehensive enough to establish robust turning guideline for ADI grades. In addition, past ADI studies have not used contemporary cutting tools. Only a few studies have evaluated common coated tungsten carbide inserts. Also, even studies using coated tungsten carbide inserts were conducted without cutting fluids (dry turning). Clearly information is lacking on the turning ADI grades using contemporary coated carbide inserts and contemporary cutting fluids.

54 41 Chapter 4 RESEARCH PLAN The primary objective of the research is to evaluate and compare the machinability of ADI GR900, GR1050, GR1200 to ductile iron by conventional turning. The turning methodology and the machinability evaluation metric would be discussed in this chapter. The machinability metrics used for evaluation are tool life, surface roughness, and chip formation. In addition, the machining specifics - pre-set of the machine platform, cutting tool information, machining parameters (cutting speed, feed rate, depth of cut), cutting fluid (coolant), and sample preparation will be described. Methods of calibrating coolant concentration and the procedures for collecting, analyzing the results of tool wear and surface roughness studies will be described. Workpiece material characteristics Commercially produced ductile iron test cylinders, austempered by a commercial heat treater (project sponsor), were used in this study. Also similar as cast samples (ductile iron) were used for comparison testing. The chemical compositions of ductile iron used in this experiment were shown in table 3. Test cylinders used in this study were produced in two different heats with only slight chemical composition differences between the heats. During austempering, all the ductile irons were austenizing at 1625 for 128 minutes but quenched and held at different austempering temperature and times in order to produce the different grades of ADI from the same base iron. The GR900 was austempering for 134 minutes at 720F, GR 1050 was austempering for 171 minutes at 670, and GR1200 was austempering for 225 minutes at 600. The microstructure was also evaluated. The nodule count of ductile irons used in this

55 42 studies was nodules mm 2 with a nodularity of 93%. The austenite content remaining in the ADI microstructure after heat treatment was measured by X-ray diffraction (XRD). The XRD was collected with a Scintag XRD-2000 θ/θ diffractometer with a copper target x-ray tube from 70 to 105 2θ, with a step size of 0.05, and 30 second count time. The microstructure of different grades of ADI is shown on Figure 17. The results of XRD analysis is shown in Table 4. It demonstrates that the GR900 matrix is 40% austenite and 60% ferrite, the GR1050 matrix is 37% austenite and 63% ferrite, and the GR1200 matrix is 30% austenite and 70% ferrite all typical values. Table 3: Chemical composition of ductile irons used in this study Element Heat 1(%) Heat 2(%) C S P Si Mn Cr Ni Mo <0.01 <0.01 V < Al Cu Mg Ti Sb <0.005 <0.005 Ce Sn <0.005 <0.005 Table 4: Ferrite and austenite volume fraction of each ADI grade after heat treated as measured by X ray diffraction

56 Sample *Ferrite *Austenite 95% Confidence volume fraction volume fraction 95% Confidence GR GR GR * Matrix volume fraction Figure 17: The microstructure of each ADI grade Hardness testing of different types of material was performed using the Brinell hardness test with a 3000 kg load scale and a 10 mm steel ball indenter. The hardness measurement results are shown in table 5. The result shows that the hardness increased from GR900 to GR1200 with the being the ductile iron being the softest material. By comparing the result of the hardness measurement to the ASTM standard datas for ADI, the hardness of all grades of ADI was within specification.

57 44 Table 5: Brinell hardness of each t material tested GR900 GR1050 GR1200 DI test test test test test Average ASTM A897/A879M Experimental platform All experiments were conducted on the HAAS ST-20 CNC lathe with the maximum spindle speed of 4000 rpm. Single point straight turning was conducted under the different cutting parameters and configurations. The cutting tools used in the experiments - SECO CNMG M5 TK2001 (ISO type K), tungsten carbide general cutting inserts coated with CVD Ti (C, N) + Al2O3 coatings, with 0.3 mm land width, 5 land angle, and 0.8 mm nose radius, Figure 18. A new cutting tool edge was used for each single turning experiment. The tool holder holding the cutting inserts (SECO DCLNR2525M12-M), Figure 19 resulted in a rake angle of -6 orthogonal, 95 of cutting angle, inclination angle of -6, side cutting angle of 5, and end cutting angle of 5. [52]. The ADI workpiece was machined in the center of the workpiece. Before each experiment started, the ADI workpieces were prior machined by facing and light turning in order to remove the heat treated outer layer. The cutting parameters used to pre-turn the workpiece were at a constant feed rate of ipr, and a depth of cut of 0.06 inches (greater than the total nose radius - 0.8mm). However different cutting speeds, based on the different grades of material were used for pre-turning fpm for ADI GR900, GR1050, and ductile iron ; and 250 fpm for ADI GR1200. In addition, actual surface finish is compared to the theoretical surface

58 45 roughness. In this study, the constant feed rate (0.012 ipr) exceed the theoretical (0.8 mm nose radius) limit, which can be calculate by equation 26. Therefore, the scallop generation will be the scenario of advanced end cutting edge intersects the point of tangency of the side cutting edge. Based on the constant feed rate and nose radius, the theoretical surface roughness R a (115 uin) can be calculated by equation 27 and 28: NR limit = 2 NR sin(ecea) equation 26 Feed rate = NR 2 (NR H) 2 + NR sin(ecea) + NR cot(ecea) ( H NR 1 + cos(ecea)) equation 27 R a = H/4 equation 28 Where NR is nose radius, H is scallop height, and ECEA is end cutting edge angle. Figure 18: SECO CNMG M5 TK2001 insert used in turning studies

59 46 Figure 19: the SECO DCLNR2525M12-M tool holder fixed on the HAAS lathe turret The test workpiece was a thick wall, hollow cylinder. The dimension of workpiece before premachining was 7 inches outer diameter, 4.5 inches inner diameter, and 8 inches of the length as shown in Figure 20. The original workpiece was turned down to an outside diameter of 6.85 inches, Figure 21, with a separate tool insert before conducting the turning tool life experiments.. The cutting length was 7.5 inches for each pass. When the cutting tool had traveled to the end of the pass (7.5 inches), the cutting tool returned to the zero and the next pass would be started. The workpieces started at the diameter of 6.85 inches and stopped when the outside diameter was 5.41 inches. (12 test passes) If the cutting tool wear did not reach the limitation of tool life, another new workpiece would be fixtured to conduct until the tool life limits were reached. The data for tool wear, surface roughness, and chip formation samples was collected every two turning passes.

60 47 Figure 20: The experimental ADI test cylinders prior to pre turning Figure 21: An ADI workpiece after initial surface removal prior to turning tests The coolant used for all experiment was Quakercool 7020-CG. The concentration of the coolant was controlled between 7 and 8 percent for whole experiments. Two measurement methods were used to verify the concentration of the coolant. Both methods for measuring the concentration of coolant are conducted using a Westover Portable Refractometer model RHB-32, Figure 22. The principle of obtaining the concentration of the solution by refractometer was to measure the ratio of water and non-water in the solution. The measurement process is shown in Figure 23. The test sample was dropped on the prism of the refractometer and the result would be displayed on the

61 48 percentage slide which was built inside the refractometer. The coolant concentration was first measured when mixing the coolant solution at 7 percent before placing it the machine system. Figure 22: The Westover Portable Refractometer model RHB-32 used for cutting fluid characterization Figure 23: The procedure for measuring coolant concentration shown schematically Also the concentration was again measured during pre-machining to assure the concentration of the coolant solution is between 7 to 8 percent during machining trials. The machining setup of the experiment is shown in figure 19. One side of the ADI workpiece would be fixed in the chuck of the headstock and centered in the tailstock for the other side of the workpiece as shown in figure 24. The tool holder and the coolant nozzle were fixed on the turret and traveled in the direction of the tool feed.

62 49 Figure 24: Test cylinder placement for turning studies Machinability metrics The machinability of GR900, GR1050, GR1200, and ductile iron was evaluated by measuring tool life, surface roughness, and chip formation during straight turning experiments. Three speeds were investigated of each grade of the ADI and ductile iron with three replications of every condition. In addition, one extra cutting speed experiment for GR900 and GR1050, and two for GR1200 were conducted after preliminary data analysis in order to build more reliable turning models. The end of tool life using was defined as when the maximum width of flank wear land (VB B max) was reached 0.6mm or when the uniform width of flank wear land VB reached 0.3mm. [8] The tool wear measurements were taken after each tool pass was completed (7.5 inches) or when the tool reaching a limitation condition as indicated by sound change or spark generation during machining. The cutting tool would be taken off the tool holder after the end of the pass and

63 50 placed on a measurement fixture, Figure 25, to measure the tool wear. The equipment for measuring the tool wear was a Nikon SMZ800 stereoscope, at a magnification of 63X. The magnification of stereoscope was verified using a standard glass to calibrate. In addition, the chip formation was observed by the Nikon SMZ800 stereoscope at 1X. After the completely cleaning the workpiece, a sample of the chips generated in the first turning pass was also collected and analyzed. The surface roughness also were measured and recorded by the Mahr Pocket Surf III profilometer, Figure 26, after each pass. Two replicates were measured at three points and one replication was measured at five points. The profilometer was set to travel a length of 0.5mm with 0.8mm of cutoff wavelength to collect the surface roughness data. The profilometer would be calibrated by the EPL-1691 Riser Plate shown in Figure 26 before measuring the workpiece. Initial cutting speeds for the various materials were based on prior literature data. The cutting speed was set at a constant surface speed regardless of workpiece diameter. The GR900 and GR1050 ADI experiments were conducted at speeds ranging from 375 to 1000 fpm, the GR1200 ADI were conducted at speeds ranging from 250 to 500 fpm, and speeds ranging from 500 to 1000 fpm for the ductile iron. All other cutting factors were held constant. All samples were machined at a constant feed rate of ipr and depth of cut of 0.06 inches. A summary of the cutting parameters used for the experiments for each grade of ADI and ductile iron are shown in the Table 6.

64 51 Table 6: Cutting parameters used in this study for each grade of ADI and for ductile iron. GR900 GR1050 GR1200 Ductile iron cutting speed (FPM) feed rate (ipr) depth of cut (inch) cutting length (inch) Number of Replicates

65 Figure 25: Nikon SMZ800 stereoscope and fixtures used for tool wear measurement. 52

66 Figure 26: Mahr Pocket Surf III profilometer and the calibration piece EPL-1691 Riser Plate used for surface roughness measurements. 53

67 54 Chapter 5 RESULTS The results of the experiment, machinability evaluations of ADI and ductile iron, are presented in three parts. The first part evaluates chip formation from all grades of ADI and the ductile iron during turning. The second part discusses the surface roughness analysis for each experiment for all material types. Lastly, the tool life under the different cutting conditions for all materials is presented. Chip formation Representative chips produced from turning experiments are shown in the Table 7. The results show the effect of different turning speeds from 250 fpm to 1000 fpm on the chip formation for all materials studied. In general, the chips generated from all grades of ADI work and the ductile iron at all speeds are short segmented. Cutting speed had little or no influence on chip morphology. The length and shape of chips of all grades of ADI and ductile iron at all cutting speeds was similar with classic c-shaped or connected c-shaped chips.

68 55 GR900 GR1050 GR1200 DI V = 375 fpm V = 375 fpm V = 250 fpm V = 500 fpm V = 500 fpm V = 500 fpm V = 300 fpm V = 750 fpm V = 750 fpm V = 750 fpm V = 375 fpm V = 1000 fpm V = 1000 fpm V = 1000 fpm V = 400 fpm V = 500 fpm Table 7: Effect of cutting speed on chip form for ADI grades and DI (Depth of cut = 0.06 in, feed rate = ipr)

69 56 Tool life Tool life is typically the most important indicator of machinability. In this research, the tool wear criteria used to estimate tool life is when the flank wear land width reached 0.3 mm uniform wear (VBBmax). The tool wear was measured after each 7.5 in of workpiece traverse. However, sometimes the tool life was ended mid pass by catastrophic tool failure. The typical catastrophic failure observed was when the cutting edge suddenly fractured during machining and generated visible sparks. Although in most cases progressive tool wear was observed, in the case of machining GR1050 and GR1200, catastrophic tool failure happened occasionally. Typical examples of flank wear measurements are shown in Figure 27 and Figure 28, which show examples of catastrophic tool failure during machining GR1050 at cutting speed of 1000 fpm and progressive wear when machining GR900 at a cutting speed of 1000 fpm respectively. Figure 27: Example of tool wear development for ADI GR1050 at 1000 fpm (Depth of cut = 0.06 in, feed rate = ipr), at cutting lengths (1) 7.5 in, (2) 15 in, (3) 22.5 in, and (4) 30 in, after tool tip fracture (63x)

70 57 Figure 28: Example of tool wear development for ADI GR1050 at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr), at cutting lengths (1) 7.5 in, (2) 15 in, (3) 22.5 in, (4) 30 in, (5) 37.5 in, (6) 45 in, (7) 52.5 in, (8) 60 in, (9) 67.5 in, and (10) 75 in, (63X) As expected, the result of experiments show that the cutting speed has a very significant influence on tool wear rates for every grade of ADI and for DI The higher the cutting speed, the faster the tool wear rate. From the tool wear data at each cutting condition, tool wear progression for each grade of ADI and DI can be evaluated by fitted cubic polynomial as shown in

71 58 Figures In these figures, VBBmax is shown as the function of cutting time. The fitted trend line polynomial equation and the R2 values for each grade of ADI and DI are presented in Table 8. Since the R2 values for each fitted curve are high, the fitted wear trend lines of different cutting speed ( fpm for GR900 and GR1050, for GR1200, and for DI ) accurately describe tool wear progression. Table 8: Tool wear progression polynomial fit equations and R2 values for different the grades of ADI and DI Cutting speed [fpm] Cubic polynomial fit equations ln V = 8E-05(ln T) (ln T) (ln T) ln V = 5E-05(ln T) (ln T) (ln T) ln V = (ln T) ln T) (ln T) ln V = (ln T) ln T) (ln T) ln V = 9E-05(ln T) ln T) (ln T) ln V = (ln T) ln T) (ln T) ln V = (ln T) ln T) (ln T) ln V = (ln T) (ln T) ln V = 3E-06(ln T) (ln T) (ln T) ln V = 4E-05(ln T) (ln T) (ln T) ln V = (ln T) (ln T) (ln T) ln V = -8E-05(ln T) (ln T) (ln T) ln V = (ln T) (ln T) (ln T) DI DI DI ln V = 2E-06(ln T) (ln T) (ln T) ln V = (ln T) (ln T) (ln T) ln V = (ln T) (ln T) (ln T) R2

72 Flank wear (mm) Flank wear (mm) 59 GR R² = R² = R² = R² = R² = Time (min) VBBmax Figure 27: Tool wear progressions for GR1200 at different cutting speeds, (Depth of cut = 0.06 in, feed rate = ipr) 0.6 GR R² = R² = R² = R² = Time (min) VBBmax Figure 28: Tool wear progressions for GR1050 at different cutting speeds, (Depth of cut = 0.06 in, feed rate = ipr)

73 Flank wear (mm) Flank wear (mm) R² = R² = GR900 R² = R² = VBBmax Time (min) Figure 29: Tool wear progressions for GR900 at different cutting speeds, (Depth of cut = 0.06 in, feed rate = ipr) DI R² = R² = R² = VBBmax Time (min) Figure 30: Tool wear progressions for DI at different cutting speeds, (Depth of cut = 0.06 in, feed rate = ipr)

74 61 Figure 29 displays tool wear progression for GR1200 at cutting speeds ranging from 250 to 500 fpm. In Table 8, the high R2 values when machining GR1200 at each cutting speed indicate that the wear progression curve and cubic polynomial equations are accurate. Figure 29 shows that increasing cutting speeds progressively shorten the tool life. In terms of tool wear progressions, except for 400 fpm, all the wear curves show initial rapid tool wear followed by a gradual linear wear profile until tool wear reaches the final limit of tool life. At a cutting speed at 400 fpm, the tool wear did not steadily reach the limitation of the tool life before catastrophic tool failure occurred in one of the testing trials. Therefore, the lowest cutting speed (250 fpm) is considered as the most cost-effective speed of machining GR1200 ADI. The tool wear progression curve for GR1050 ADI at cutting speeds ranging from 375 to 1000 fpm is shown in Figure 30. Similar results were observed for ADI GR1200. The high R2 values for machining GR1050 in Table 8 indicate that the fitted tool wear progression curves are accurate. Figure 30 shows that tool life is shorter when turning at high cutting speed (750, 1000 fpm) than at low cutting speeds (375, 500 fpm). The wear progression curves at each cutting speed all show rapid increases in tool wear initially followed by a linear wear rate until the limits of tool life are reached. Catastrophic tool failure often took place when the cutting speed was high (750 fpm) during the turning of GR1050 ADI. However, the measured tool wear for ADI GR1050 was significantly less than for GR1200. The tool life at a cutting speed at 375 and 500 fpm was approximate the same in terms of cutting time, but the tool life at a cutting speed at 500 fpm was higher than at 375 fpm if measured by material removal rates. Figure 31 shows VBBmax as the function of effective cutting time when turning GR900 ADI at cutting speeds ranging from 375 to 1000 fpm. The results are very similar to the results for GR1050 in general. The correlation coefficient for each experiment were also high for GR900 ADI. Also to GR1050, the wear progression curves increased rapidly initially and then exhibited

75 Flank wear (mm) 62 characteristic linear wear rates until the end of tool life. Tool life was low when cutting speeds were high (750, 1000 fpm). However, only small differences in tool life when turning GR900 ADI at cutting speeds at 500 fpm and 375 fpm were observed. Analysis shows that 500 fpm is the best cutting speed for ADI GR1050. Figure 32 shows VBBmax as a function of effective cutting time for DI Similar to the result of machining all grade of ADI, the fitted tool wear progression curves and polynomial equations when machining at cutting speeds ranging from 500 to 1000 fpm had characteristic shapes. This tool wear progression can be presented in other form by comparing wear rates for the different grades of ADI and DI at the same cutting speed. Figures 33 and 34 show similar tool wear curves at cutting speeds of 375 and 500 fpm fpm GR900 GR1050 GR1200 VBBmax Time (min) Figure 31: Tool wear progressions for GR900, GR1050 and GR1200 at V = 375 fpm, (Depth of cut = 0.06 in, feed rate = ipr)

76 Flank wear (mm) fpm Time (min) GR900 GR1050 GR1200 VBBmax Figure 32: Tool wear progressions for GR900, GR1050 and GR1200 at V = 500 fpm, (Depth of cut = 0.06 in, feed rate = ipr) Tool life predictions can be established based on wear progressions to compute the time or length of machining until the limiting tool life criteria (0.3 mm uniform flank wear) has been reached, Table 8. Based on Table 8, the tool life in terms of cutting time and cutting length are shown in Figure 35 and Figure 36 respectively. Tool life for GR1200 ADI ranged from 64.6 to 6.6 minutes with tool life decreasing dramatically as cutting speed increased from 250 to 375 fpm, followed by linear decrease in tool life of higher cutting speeds ( fpm). The tool life of GR1050 ADI gradually decreased from 22.4 to 5.33 minutes when the cutting speed was increased. For GR900, the tool life increased slightly as the cutting speed increased from 375 to 500 fpm and then gradually decreased to 5.1 minutes as cutting speed increased.

77 Cutting length Cutting time (min) Cutting speed (fpm) DI Figure 33: Effect of cutting speed on tool life for each grade of ADI and DI (in terms of cutting time) (Depth of cut = 0.06 in, feed rate = ipr) DI Cutting speed (fpm) Figure 34: Effect of cutting speed on tool life for each grade of ADI and DI (in terms of cutting length) (Depth of cut = 0.06 in, feed rate = ipr) The tool life prediction graph clearly shows that when machining all grades of ADI the same speed, GR1200 has the shortest tool life. Comparing the GR900 and GR1050, the tool life of GR1050 was approximately the same as GR900 but slightly better than GR900 at low cutting

78 65 speeds. But when the cutting speed was increased, the GR900 exhibited better tool life than GR1050. Other method for estimating tool life is to calculate the total cutting length. (at constant depth of cut) The tool life of GR1200 ranged from 22.5 to in, as cutting speed decreased. However, for GR900 and GR1050, cutting length increased in the beginning and followed by linear life. Comparable turning data is shown for as cast DI in Table 8 and Figure 32. In terms of cutting length, the tool life ranged from to 45 in when the cutting speed was increased. As expected, the tool life when machining DI is much greater than for any grade of ADI at any cutting speed. The tool life prediction equations for each grade of ADI were developed using the Taylor tool life equation. The relationship between cutting speed and tool life, in terms of time, was used to build the equations. The constant n and C in the Taylor equation for each grade of ADI were determined from the tool life data. The relationship between tool life and cutting speed was plotted on a ln ln graph, as shown in Figure 37. The n value was estimated from the slope and the C value was determined from Y-intercept of the linear fitted line of Figure 38. The developed tool life equations for GR 900, GR1050, and GR1200 are presented below. TGR900 = V 1/0.55, R² = (equation 26) TGR1050 = V 1/0.434, R² = (equation 27) TGR1200 = V 1/0.303, R² = (equation 28) Where V is measured in rpm and T is measured in minutes.

79 Cutting speed (rpm) (ln V) ln V = (ln T) GR900 ln V = (ln T) GR1050 ln V = (ln T) GR1200 GR GR1050 GR Tool life (min) (ln T) Figure 35: Ln-Ln tool life plot for developing the Taylor tool life equation Surface roughness The results of surface roughness of each grade for ADI and DI based on cutting length at different cutting speeds is shown in Figures The effect of cutting speed on surface roughness for each grade of ADI and DI are shown in Table 9 and Figure 42. The average surface roughness of GR1200 was slightly different than GR1050 and GR900. The average surface roughness of GR1200 ranged from 113 to 62 uin as cutting speed increased. The surface finish dropped from 113 to 110 uin, for the speed range of fpms, and then significantly decreased to 62 uin at low cutting speeds. ( fpm) And as cutting speed increased, the surface roughness of GR1050 and GR900 materials decreased at low speeds ( fpm) followed by only slightly linear decrease in surface roughness at high cutting speeds ( fpm). The surface roughness of GR 1050, GR900, and DI ranged from 106

80 67 to 61 uin, 105 to 50 uin, and 93 to 59 uin respectively as cutting speed increased. The surface finish of DI also decreased as cutting speed increased. In Figures 38-41, the surface roughness is displayed in terms of cutting length. The surface finish for all grades of ADI gradually decreased at high cutting speeds but gradually increased at low cutting speeds except for GR1200 at a cutting speed of 300 fpm. The surface roughness of the GR1200 increased at the beginning and then decreased until the tool broke.

81 68 Figure 36: Effect of cutting length on surface roughness (Ra) for ADI GR900 at different cutting speeds (Depth of cut = 0.06 in, feed rate = ipr). Figure 37: Effect of cutting length on surface roughness (Ra) for ADI GR1050 at different cutting speeds (Depth of cut = 0.06 in, feed rate = ipr).

82 Figure 38: Effect of cutting length on surface roughness (Ra) for ADI GR1200 at different cutting speeds (Depth of cut = 0.06 in, feed rate = ipr). 69

83 Surface roughness 70 Figure 39: Effect of cutting length on surface roughness (Ra) for DI at different cutting speeds (Depth of cut = 0.06 in, feed rate = ipr) Cutting speed (fpm) DI Figure 40: Effect of cutting speed on surface roughness (Ra) for all grades of ADI and DI (Depth of cut = 0.06 in, feed rate = ipr).

84 Table 9: Effect of cutting speed on tool life for each grade of ADI and DI in terms of cutting time and cutting length (Depth of cut = 0.06 in, feed rate = ipr) 71 Grade Cutting speed [fpm] Average Tool life, cutting time [min] Standard Deviation [min] Average Tool life, cutting length [in] Standard Deviation [in] * * DI DI DI

85 Table 10: Average surface roughness (Ra) for grades of ADI and DI (Depth of cut = 0.06 in, feed rate = ipr) 72 Cutting Speed [fpm] Average Surface Roughness Standard Deviation * DI DI DI

86 73 Chapter 6 DISCUSSION This research thoroughly evaluated of machinability during turning for different grades of ADI (GR900, GR1050, GR1200) and compared them to DI In this section, the results of machining trials with ADI will be discussed and analyzed. Tool wear measurements are an important indicator for evaluating the machinability of ADI during turning. Flank wear measurement is commonly observed on both end flank and side flank of a turning tools. In this study the measurement value of end flank wear had a strong correlation to the side flank wear, with the side flank wear being much more consistent from tool to tool. Therefore, side flank wear was used as the critical flank wear criteria for evaluating tool wear in this study. The flank wear patterns for insert when machining GR1050 ADI at low cutting speeds and high cutting speeds until the tool failure are shown in Figure 27 and Figure 28, respectively. These wear patterns were typical for all grade of ADI at different cutting conditions. When machining ADI at lower cutting speeds, the cutting edge stayed in a stable wear condition - gradually increasing the wear land at a wide range of cutting times, as shown in Figures 38-1 to 8. This stable increasing abrasive wear continued until a sufficient number of mechanical impact loading cycle occurred, resulting in the coating being worn away. [18] As shown in Figure 38-9 and 10, the abrasive wear then increased significantly causing rapid tool failure. The Taylor tool life equation was used to realize and evaluate the relationship between tool life and cutting speed. [50] In this trial, the n values were quite different for the three different grades of ADI. These results suggest that the tool material is most sensitive to cutting speed. Tool wear for the GR900 material was the most sensitive to cutting speed (n = 0.55), followed by GR1050 (n = 0.434), and GR1200 (n = 0.303). The cutting tool SECO CNMG M5 TK2001

87 74 showed the larger n value compared to uncoated tungsten carbide (0.2) and TiC or TiN coated tungsten carbide (0.3) used in previous studies. [51] Therefore, it can be assumed that the SECO cutting tool is less sensitive to cutting speed than uncoated tungsten carbide and TiC or TiN coated tungsten carbide cutting tools. Katuku, Koursaris, and Sigalas [33] reported a Taylor equation with the n value of 0.51 and C value of 3*10^7 for machining GR900 with a PCBN cutting tool. (finishing cut, dry turning) The n value for the PCBN cutting tool (0.51) and SECO CNMG M5 TK2001 in this study (0.55) is very close, suggesting that level of sensitivity for machining GR900 ADI is similar. The value for the SECO tool is much smaller than the PCBN cutting tool when machining GR900 ADI. [43] The significant difference in C values could come from differences in depth of cut, feed rate, and material removal rates between two research studies as well as differences in tool material. The Taylor tool equation can be used to estimate expected tool life under various cutting conditions. Predicted tool life for 10, 30, and 60 minute tool life criteria for fixed cutting conditions (depth of cut = 0.06 in, feed rate = ipr) are based on the developed equations for different grades of ADI. The GR900 would result in 10 minutes of the tool life at a cutting speed of 677 fpm, 30 minutes tool life at a cutting speed of 370 fpm, and 60 minutes tool life at a cutting speed of 253 fpm. When machining GR1050, a cutting speed of 598 fpm would result in 10 minutes of tool life, the cutting speed of 371 fpm would result in 30 minutes of tool life, and the cutting speed of 275 fpm would result in 60 minutes of tool life. For machining GR1200, when the desired tool lives are 10, 30, and 60 minutes, the cutting speed should be 437, 313, and 254 fpm respectively. DI is the preferred as cast grade used as the starting material before ADI heat treatment, had a lower hardness than any grade of ADI. As expected, the expected tool life of DI is longer than ADI both in terms of cutting length and cutting time for all ADI grades.

88 75 The chip formation analysis is another important way to characterize machinability. From Table 7, the chips under all cutting conditions are short segmented with only slight geometry differences. Short segmented chips rather than discontinuous chips were generated because the insert used had a chip breaking geometry. There was no significant difference in chip formation between the beginning of the cut and the end of tool life. The tool wear profiles suggest that both nose radius and side cutting edge had significant influence on chip formation. The cutting speed has only a minor effect on the chip formation. At high feed rates, it has been reported that chip shape changed from coiled to C-shaped. [48] However, C. Wang, Guo, W. Wang, and Dong [48] reported that individual C - shaped chips formed during ADI turning at high feed rates in this research (0.012 ipr) the chips were also C - shaped. The C shaped chips formed in this research not only existed in single pieces but also in connected C shape segments (2 5 pieces). Also, the size and length of the chips were approximately the same between the different cutting speed as feed rate for any grade of ADI and DI There was no observable uneven surface finish on the turned surfaces due to chatter for any of the ADI grades on ADI machining conditions used. Though in the Figure 41, there is a slight decrease in surface roughness when cutting speed increases of all ADI grades and DI Also, there were also no significant difference in surface roughness between different materials turned at the same cutting speed. Initial rapid tool wear observed on nose radius of the tool after the first pass of cutting. Creating a measured surface roughness value for ADI grades at all cutting conditions that were less than the theoretical average surface roughness 115 uin (feed rate = ipr, nose radius = 0.8 mm). This also suggests BUE formation during turning was not significant. The characteristic rapid tool wear on nose can be expected to cause a surface wiping effect that slightly improves surface finish as the tool wears. This effect which is likely the reason for the decreasing trend of surface roughness observed during tool wear studies.

89 76 Chapter 7 CONCLUSION AND FUTURE WORK Due to the lack of technical information on the machinability of ADI via turning, applications of ADI for new products has been limited. The present research has addressed this need and was started to analyze the machinability of different grades of ADI (GR900. GR1050, GR1200) via turning. The effects of machining parameters on machining performance was measuring by tool life, surface roughness and characterization the chip formation during lubricated turning trials with coated carbide tools. All machining chip, for all grades were discontinuous. The generation of discontinuous chips could be attributed to both the graphite nodules and the hardness of the ADI matrix structure. In general, from the view of tool life, The GR900 exhibited longer tool life than GR1050 and GR1200. In terms of resultant surface roughness, GR1050 showed a slightly better surface finish that the other grades of ADI. The as cast DI had the longest tool life. However, the surface roughness was slightly higher than any other grades of ADI. Taylor tool life coefficients for the ADI grades were calculated. The constant C was shown to be slightly dependent on the grades of ADI. The constant n was very different for each grade of ADI. In addition, the influence of cutting speed on tool life, surface roughness, and chip formation during turning were investigated. The tool wear studies have characterized the relationship between cutting speed and tool life. At high cutting speed, the tool wear is more rapid. Flank wear was the dominant wear mechanisms; however, some fracturing of the cutting edge was observed during aggressive machining. Cracking was observed at high cutting speeds in all grades of ADI. This change in tool failure mechanism was associated with very high chip and tool temperatures.

90 77 Future extensions of this research should further investigate the machinability of ADI under different turning conditions such as the depth of cut and feed rate. Also, the possible effect of undesirable martensite transformation on the surface of ADI during turning should be thoroughly investigated by careful microstructure characterization of machined surfaces.

91 78 References [1] S. Kalpakjian, S. R. Schmid, Manufacturing engineering and technology, 6 th edition, [2] W. h. Chubberly, R. Bakerjian, Tool and manufacturing engineers handbook, desk edition, five-volume 4 th edition, 1989 [3] ASM, Metals handbook, volume 16, machining, 9 th edition, [4] J. T. Black, R. A. Kohser, Degarmo s materials and processes in manufacturing, 11 th edition, [5] D. A. Stephenson and J. S. Agapiou, Metal Cutting Theory and Practice, Boca Raton, FL: Taylor &Franics Group, LLC, [6] G. Smith, Cutting Tool Technology, London: Springer-Verlag, 2008 [7] Murphy, D.W.; Aylward, P.T. Machinability of Steels, 1998 [8] ISO 3685,InterationalStandard, [9] Krizbergs, J. &Kromanis, A Method for prediction of the surface roughness 3D parameter according to technological parameters, 2006 [10] A. Bewoor, V. Kulkarni, Metrology & Measurement, 2009 [11] O. Eric, M. Jovanovic, L. Sidanin, D. Rajnovic and S. Zec, "The austempering study of alloyed ductile iron," Materials & Design, Volume 27, pp , [12] M. C. Cakir, A. Bayram, Y. Isik and B. Salar, "The effects of austempering temperature and time onto the machinability of austempered ductile iron" Materials Science & Engineering A, vol. 407, no. 1-2, pp , [13] P. Shanmugam, P. Prasad rao, K. RajendraUdupa, N. Venkataraman "Effect of microstructure on the fatigue" Journal of Materials Science, Volume 29, Issue 18, pp , 1994 [14] K. Brandenberg, "Machining austempered ductile iron.," Society of Manufacturing Engineering., vol. 128, no. 5, pp , [15] M. Baydogan and H.Cimenoglu "The effect of austempering time on mechanical properties of a ductile iron" Scandinavian Journal of Metallurgy Volume 30, Issue 6, pages , December [16] S. K. Swain, R. K. Panda, J. P. Dhal, S. C. Mishra and S. Sen "Phase Investigation of Austempered Ductile Iron" Orissa Journal of Physics Vol. 19, No.1 February 2012 pp [17] A. Abedi, S.P.H. Marashi, K. Sohrabi, M. Marvastian, S.M.H. Mirbagheri"The Effect of Heat Treatment Parameters on Microstructure and Toughness of Austempered Ductile Iron (ADI)" Advanced Materials Research, Vols , pp , 2011 [18] K. L. Hayrynen"The Production of Austempered Ductile Iron (ADI)" World Conference on ADI, 2002

92 79 [19] Y.J. Kim, H.Shin, H. Park, J. D. Lim "Investigation into mechanical properties of austempered ductile cast iron (ADI) in accordance with austempering temperature"materials Letters Volume 62, Issue 3, 15 February 2008, Pages [20] B. V. S. Kovacs, "Austempered Ductile Iron: Fact and Fiction," Modern Casting, pp , 1990 [21] A. A. Zhukov, A. Basak& A. B. Yanchenko (1997)" New viewpoints and technologies in field of austempering of Fe C alloys" Materials Science and Technology, 13:5, [22] J. Zimba, D.J. Simbi, E. Navara "Austempered ductile iron: an alternative material for earth moving components" Cement & Concrete Composites Volume 25, Issue 6, Pages [23] J. R. Keough, K. L. Hayrynen "Automotive Applications of Austempered Ductile Iron (ADI): A Critical Review," Society of Automotive Engineers, [24] J. R. Keough, K. L. Hayrynen, G. L. Pioszak "Designing with Austempered Ductile Iron (ADI)" AFS Proceedings 2010 [25] A.A. Nofal, L. Jekova "Novel processing techniques and applications of austempered ductile iron" Journal of the University of Chemical Technology and Metallurgy, 44(3), pp , 2009 [26] K. L. Hayrynen and J. R. Keough, "Austempered Ductile Iron-The State of the Industry in 2003," in Keith D. Millis Symposium, Louisville, 2003 [27] "Ductile iron Data for Design Engineers - section IV. Austempered ductile iron" Ductile iron society, [Online]. Available: [Accessed 3 April 2016]. [28] A.Sinlah Jr."A Study of Machinability in Milling of Austempered Ductile Iron (ADI) ADI 900, ADI 1050, And ADI 1200 with Carbide Tools." Master thesis, Pennsylvania state University, 2014 [29] R. C. Voigt, "Austempered Ductile Iron - Processing and Properties," Cast Metals, vol. 2, no. 2, pp , [30] C. F. Han "Effects of Nickel and Cobalt Elements and Variable Austempering on Microstructure and Mechanical Properties of Ductile Iron" Master thesis in Chinese, Tatung University, 2006 [31] R. C. Voigt, C.R. Loper, Jr. "Austempered Ductile Iron - Process control and Quality Assurance" J. heat treating, vol. 3, no. 4, pp , 1984 [32] A. Polishetty, "Machinability and microstructural studies on phase transformations in Austempered Ductile Iron," Doctorial Thesis, Auckland University of Technology, [33] R. C. Voigt "Microstructural Analysis of austempered ductile iron using the scanning electron microscope" AFS Transactions, pp.83-89, 1982 [34] K.L.Hayrynen,D.J.Moore and K.B.Rundman "More about the Tensile and Fatigue Properties of Relatively Pure ADI" AFS Transactions, pp , 1993

93 80 [35] B. Avishan, S. Yazdani and D. JalaliVahid, "The influence of depth of cut on the machinability of an alloyed austempered ductile iron.," Materials Science & Engineering, no. 523, pp , [36] J. Pilc, M. Šajgalík, J. Holubják, M. Piešová, L. Zaušková, O. Babík, V. Kuždák, J. Rákoci "Austempered Ductile Iron Machining" Technological Engineering. Volume 12, Issue 1, Pages 9 12, 2015 [37] W.H. Yang, Y.S. Tarng"Design optimization of cutting parameters for turning operations based on the Taguchi method" Journal of Materials Processing Technology, Volume 84, Issues 1 3, Pages , 1998 [38] G. H. Gowd, M. V. Goud, K. D. Theja, M. G. Reddy "Optimal Selection Of Machining Parameters In CNC Turning Process Of EN-31 Using Intelligent Hybrid Decision Making Tools" Procedia Engineering, Volume 97,Pages , 2014 [39] M.E.R. Bonifacio, A.E. Diniz "Correlating tool wear, tool life, surface roughness and tool vibration in finish turning with coated carbide tools"wear, Volume 173, Issues 1 2, Pages , 1994 [40] S.K Choudhury, K.K Kishore "Tool wear measurement in turning using force ratio"international Journal of Machine Tools and Manufacture, Volume 40, Issue 6, Pages , 2000 [41] A. Polishetty, M. Goldberg, G. Littlefair "Wear Characteristics of Ultra-Hard Cutting Tools when Machining Austempered Ductile Iron" International Journal of Mechanical & Mechatronics Engineering 10(01):1-6, 2010 [42] A. Akdemir, S. Yazman, H. Saglam and M. Uyaner, "The effects of cutting speed and depth of cut on machinability characteristics of austempered ductile iron," Journal of Manufacturing Science and Engineering-Transactions of the Asme, vol. 134, 2012 [43] K. Katuku, A. Koursaris and I. Sigalas, "Wear, cutting forces and chip characteristics when dry turning ASTM grade 2 austempered ductile iron with PcBN cutting tools under finishing conditions," Journal of Materials Processing Tech, vol. 209, no. 5, pp , 2009 [44] K. Aslantas and I. Ucun, "Evaluation of the performance of CBN tools when turning austempered ductile iron material," Journal of Manufacturing Science and Engineering, vol. 130, no. 5, pp , 2008 [45] K. Aslantas and I. Ucun, "The performance of ceramic and cermet cutting tools for the machining of austempered ductile iron," The International Journal of Advanced Manufacturing Technology, vol. 41, no. 7, pp , 2009 [46] M.Arft, F.Klocke "High Performance Turning of Austempered Ductile Iron (ADI) with adapted Cutting Inserts" 14th CIRP Conference on Modeling of Machining Operations (CIRP CMMO), 2013 [47] M. V. De Carvalho, D. M. Montenegro and J. D. Gomes, "An analysis of the machinability of ASTM grades 2 and 3 austempered ductile iron," Journal of Materials Processing Technology, vol. 213, no. 560, 2013

94 81 [48] C. Wang, X. Guo, W. Wang, Q. Dong "Study on the Influence of Cutting Parameters on Cutting Forces and Chip Shape of Austempered Ductile Iron (ADI)" The third International Multi-Conference on Engineering and Technological Innovation: IMETI 2010 [49] P. Parhad,A. Likhite, J. Bhatt, D. Peshwe"The Effect of Cutting Speed and Depth of Cut on Surface Roughness During Machining of Austempered Ductile Iron" Transactions of the Indian Institute of Metals, Volume 68, Issue 1, pp , 2015 [50] I.V. Rao G.L. Lal, International journal of machine tool design and research volume 17, issue 4, 1977, page [51] D. A. Stephenson and J. S. Agapiou, Metal Cutting Theory and Practice, Boca Raton, FL: Taylor & Franics Group, LLC, 2006 [52] SECO tool online, Available:

95 82 Appendix A - Tool wear measurements The following tables present the flank wear width measurement for all grades of ADI and DI In the case of the ADI, three trials were conducted at each cutting speed (except for extra experiments of GR1050 at the speed of 375 fpm) and two trials were conducted for the DI The limitation of the tool are when the tools have an average flank wear width of 0.3 mm, a single width of 0.5 mm, or the cutting edge are signified tool failure. Table 11: Flank wear for ADI GR900 at 375 fpm (Depth of cut = 0.06 in, feed rate = ipr) Pass Initial Diameter Final Diameter Speed (fpm) Cutting length Cut time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm) Flank Wear Trial 3 (mm) Table 12: Flank wear for ADI GR900 at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) Pass Initial Diameter Final Diameter Speed (rpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm) Flank Wear Trial 3 (mm)

96 Table 13: Flank wear for ADI GR900 at 750 fpm (Depth of cut = 0.06 in, feed rate = ipr) 83 Pass Initial Diameter Final Diameter Speed (rpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm) Flank Wear Trial 3 (mm) Table 14: Flank wear for ADI GR900 at 1000 fpm (Depth of cut = 0.06 in, feed rate = ipr) Pass Initial Diameter Final Diameter Speed (rpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm) Flank Wear Trial 3 (mm) Table 15: Flank wear for ADI GR1050 at 375 fpm (Depth of cut = 0.06 in, feed rate = ipr) Pass Initial Diameter Final Diameter Speed (fpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm)

97 Table 16: Flank wear for ADI GR1050 at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) 84 Pass Initial Diameter Final Diameter Speed (rpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm) Flank Wear Trial 3 (mm) (F) Table 17: Flank wear for ADI GR1050 at 750 fpm (Depth of cut = 0.06 in, feed rate = ipr) Pass Initial Diameter Final Diameter Speed (rpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm) Flank Wear Trial 3 (mm) Table 18: Flank wear for ADI GR1050 at 1000 fpm (Depth of cut = 0.06 in, feed rate = ipr) Pass Initial Diameter Final Diameter Speed (rpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm) Flank Wear Trial 3 (mm)

98 Table 19: Flank wear for ADI GR1200 at 250 fpm (Depth of cut = 0.06 in, feed rate = ipr) 85 Pass Initial Diameter Final Diameter Speed (fpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm) Flank Wear Trial 3 (mm) (B)

99 Table 20: Flank wear for ADI GR1200 at 300 fpm (Depth of cut = 0.06 in, feed rate = ipr) 86 Pass Initial Diameter Final Diameter Speed (fpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm) Flank Wear Trial 3 (mm) Table 21: Flank wear for ADI GR1200 at 375 fpm (Depth of cut = 0.06 in, feed rate = ipr) Pass Initial Diameter Final Diameter Speed (fpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm) Flank Wear Trial 3 (mm)

100 Table 22: Flank wear for ADI GR1200 at 400 fpm (Depth of cut = 0.06 in, feed rate = ipr) 87 Pass Initial Diameter Final Diameter Speed (fpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm) Flank Wear Trial 3 (mm) (F) Table 23: Flank wear for ADI GR1200 at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) Pass Initial Diameter Final Diameter Speed (rpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm) Flank Wear Trial 3 (mm) (F) (F)

101 Table 24: Flank wear for DI at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) 88 Pass Initial Diameter Final Diameter Speed (fpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm)

102 Table 25: Flank wear for DI at 750 fpm (Depth of cut = 0.06 in, feed rate = ipr) Pass Initial Diameter Final Diameter Speed (rpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm)

103 90 Table 26: Flank wear for DI at 1000 fpm (Depth of cut = 0.06 in, feed rate = ipr) Pass Initial Diameter Final Diameter Speed (rpm) Cutting length Cutting time (min) Flank Wear Trial 1 (mm) Flank Wear Trial 2 (mm) Table 27: Effect of cutting speed on tool life for ADI grades in terms of cutting time for all trials (Depth of cut = 0.06 in, feed rate = ipr) Grade Cutting Speed [fpm] Trial 1 [min] Trial 2 [min] Trial 3 [min] Average Tool life [min] Standard Deviation [min] * * * DI DI DI * * *

104 91 Appendix B - Surface roughness measurements The following tables show the overall average surface roughness for each trial and for each grade of ADI and ductile iron , as well as the actual three or five readings for each surface profile. Table 28: Average surface roughness for ADI GR900 (Depth of cut = 0.06 in, feed rate = ipr) Cutting Speed [fpm] Average Surface Roughness Trial 1 [μin] Average Surface Roughness Trial 2 [μin] Average Surface Roughness Trial 3 [μin] Average Surface Roughness [μin] Standard Deviation [μin] Table 29: Average surface roughness for ADI GR1050 (Depth of cut = 0.06 in, feed rate = ipr) Cutting Speed [fpm] Average Surface Roughness Trial 1 [μin] Average Surface Roughness Trial 2 [μin] Average Surface Roughness Trial 3 [μin] Average Surface Roughness [μin] Standard Deviation [μin] * * *

105 Table 30: Average surface roughness for ADI GR1200 (Depth of cut = 0.06 in, feed rate = ipr) 92 Cutting Speed [fpm] Average Surface Roughness Trial 1 [μin] Average Surface Roughness Trial 2 [μin] Average Surface Roughness Trial 3 [μin] Average Surface Roughness [μin] Standard Deviation [μin] Table 31: Average surface roughness for DI (Depth of cut = 0.06 in, feed rate = ipr) Cutting Speed [fpm] Average Surface Roughness Trial 1 [μin] Average Surface Roughness Trial 2 [μin] Average Surface Roughness [μin] Standard Deviation [μin]

106 Table 32: Surface roughness measurements for ADI GR900 at 375 fpm (Depth of cut = 0.06 in, feed rate = ipr) 93 Cutting length Reading 1 (μin) Reading 2 (μin) Reading 3 (μin) Reading 4 (μin) Reading 5 (μin) Average roughness (μin) standard deviation (μin) Trial 1 Trial 2 Trial * * * * * * * * * * * * * * * * * * * * * * * * * * * *

107 Table 33: Surface roughness measurements for ADI GR900 at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) 94 Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation Trial 1 Trial 2 Trial * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * *

108 Table 34: Surface roughness measurements for ADI GR900 at 750 fpm (Depth of cut = 0.06 in, feed rate = ipr) Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation Trial 1 Trial 2 Trial * * * * * * * * * * * * * * * * * * * * * * * * * *

109 Table 35: Surface roughness measurements for ADI GR900 at 1000 fpm (Depth of cut = 0.06 in, feed rate = ipr) 96 Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation Trial 1 Trial 2 Trial * * * * * * * 15 * * * * * * * 22.5 * * * * * * * 30 * * * * * * * * * * * * * * * * * * * Table 36: Surface roughness measurements for ADI GR1050 at 375 fpm (Depth of cut = 0.06 in, feed rate = ipr) Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation Trial * * * * * * * *

110 Table 37: Surface roughness measurements for ADI GR1050 at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) 97 Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation Trial 1 Trial 2 Trial * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * *

111 Table 38: Surface roughness measurements for ADI GR1050 at 750 fpm (Depth of cut = 0.06 in, feed rate = ipr) 98 Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation Trial 1 Trial 2 Trial * * * * * * * * * * * * * * * * * * * Table 39: Surface roughness measurements for ADI GR1050 at 1000 fpm (Depth of cut = 0.06 in, feed rate = ipr) Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation Trial 1 Trial 2 Trial * * * * * * * * * * * *

112 Table 40: Surface roughness measurements for ADI GR1200 at 250 fpm (Depth of cut = 0.06 in, feed rate = ipr) 99 Trial 1 Trial 2 Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * *

113 100 Trial * * * * * * * *

114 Table 41: Surface roughness measurements for ADI GR1200 at 300 fpm (Depth of cut = 0.06 in, feed rate = ipr) 101 Trial 1 Trial 2 Trial 3 Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * *

115 Table 42: Surface roughness measurements for ADI GR1200 at 375 fpm (Depth of cut = 0.06 in, feed rate = ipr) 102 Trial 1 Trial 2 Trial 3 Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * *

116 Table 43: Surface roughness measurements for ADI GR1200 at 400 fpm (Depth of cut = 0.06 in, feed rate = ipr) 103 Trial 1 Trial 2 Trial 3 Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation * * * * * * * * * * * * * * * * * * * * Table 44: Surface roughness measurements for ADI GR1200 at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) Trial 1 Trial 2 Trial 3 Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation * * * * * * * * * * * * * *

117 Table 45: Surface roughness measurements for DI at 500 fpm (Depth of cut = 0.06 in, feed rate = ipr) 104 Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation Trial * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * *

118 105 Trial * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * *

119 * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * *

120 Table 46: Surface roughness measurements for ADI GR1200 at 750 fpm (Depth of cut = 0.06 in, feed rate = ipr) 107 Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation Trial 1 Trial * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * * *

121 Table 47: Surface roughness measurements for ADI GR1200 at 1000 fpm (Depth of cut = 0.06 in, feed rate = ipr) 108 Cut length Reading 1 Reading 2 Reading 3 Reading 4 Reading 5 Average roughness standard deviation Trial 1 Trial * * * * * * * * * * * * * * * * * * * * * * * *

122 109 Appendix C - Insert identification charts The following tables show the insert identification methodology used to identify each insert and cutting edge for each cutting condition in the all turning experiments. Table 48: Identification chart for inserts used to machine ADI GR900 Cutting speed [fpm] Insert number Cutting edge 375 Trail Trail Trail Trail Trail Trail Trail Trail Trail Trail Trail Trail Table 49: Identification chart for inserts used to machine ADI GR1050 Cutting speed [fpm] Insert number Cutting edge 375 Trail Trail Trail Trail Trail Trail Trail Trail Trail Trail

123 110 Table 50: Identification chart for inserts used to machine ADI GR1200 Cutting speed [fpm] Insert number Cutting edge 250 Trail Trail Trail Trail Trail Trail Trail Trail Trail Trail Trail Trail Trail Trail Trail Table 51: Identification chart for inserts used to machine DI Cutting speed [fpm] Insert number Cutting edge 500 Trail Trail Trail Trail Trail Trail

124 111 Appendix D Tool Wear Measurement Method Step 1: Remove the tool from the machine Step 2: Identify the flank wear

125 112 Step 3: Place the insert on the fixture plate Step 4: Measure the face B tool wear and record the result