Ahmet Akdemir Department of Mechanical Engineering, Faculty of Engineering-Architecture, Selcuk University, Konya, Turkey 42075 e-mail: aakdemir@selcuk.edu.tr akir Yazman S Mechanical Department, Ilgın Technical Science College of Selcuk University, Konya, Turkey 42075 e-mail: syazman@selcuk.edu.tr Hacı Saglam Mechanical Department, Technical Science College of Selcuk University, Kampus 42031, Konya, Turkey e-mail: hsaglam@selcuk.edu.tr Mesut Uyaner1 Department of Metallurgy and Material Engineering, Faculty of Engineering-Architecture, Selcuk University, Konya, Turkey 42075 e-mail: muyaner@selcuk.edu.tr 1 The Effects of Cutting Speed and Depth of Cut on Machinability Characteristics of Austempered Ductile Iron Ductile iron can acquire enhanced thermal and mechanical properties from austempering heat treatment. The present study aims to identify the function of different cutting parameters affecting machinability and to quantify its effects. Turning was performed to test machinability according to the ISO3685-1993 (E) standard. After austenitizing at 900  C for 90 min, austempered ductile iron (ADI) specimens were quenched in a salt bath at 380  C for 90 min. The cutting force signals along three directions were measured in real time, whereas flank wear and surface roughness were measured offline. For the cutting parameters, the cutting speed and depth of cut were varied, but the feed rate was kept constant. In the flank wear tests, machining length was corresponded to tool life. In addition, in order to find out the effect of cutting parameters on surface roughness (Ra), tangential force (Ft), and flank wear (VB) during turning, response surface methodology (RSM) was utilized by using experimental data. The effect of the depth of cut on the surface roughness was negligible but considerable in the cutting forces. The increased cutting speed produced a positive effect on surface roughness. It is found that the cutting speed was the dominant factor on the surface roughness, tangential force, and flank wear. [DOI: 10.1115/1.4005805] Keywords: casting, austempered ductile irons, machinability, tool wear, cutting forces, surface roughness Introduction The process of casting metal is the oldest and still one of the most widely used metal-forming processes. Good rigidity, compressive strength, and fluidity are typical properties of castable materials. Therefore, cast iron can be readily cast into complex shapes, whereas steel can only be cast into simple shapes, as it has low fluidity. Ductility and strength can be improved by various treatments that affect a material’s microstructure. Many researchers [1–3] have looked into improving the mechanical properties of ductile iron (DI) by austempering. The basic structural constituents of the different types of cast iron are ferritic, pearlitic, or a mixture of these. Cast iron with a ferritic matrix and little or no pearlite is easy to machine. The history of DI has been influenced by a number of technical developments that have resulted in new business opportunities for the foundry industry, the latest of which is the introduction of austempered ductile iron (ADI) [4]. In addition to having good mechanical properties, especially tensile strength and toughness, DI has been improved by austempering. Several studies [5–8] have been performed to determine the influence of alloying elements and heat treatments on the mechanical properties of ADI. ADI materials have found many applications in recent years due to their high strength and hardness, coupled with their substantial ductility and toughness. These materials have been used for a wide variety of applications in automotive, rail, and heavy engineering industries because of their excellent mechanical properties, such as high strength with good ductility, good wear resistance, and good fatigue properties [9]. 1 Corresponding author. Contributed by the Manufacturing Engineering Division of ASME for publication in the JOURNAL OF MANUFACTURING SCIENCE AND ENGINEERING. Manuscript received April 6, 2011; final manuscript received November 2, 2011; published online March 30, 2012. Assoc. Editor: Patrick Kwon. The remarkable properties of ADI are attributed to its unique microstructure consisting of high-carbon austenite and ferrite. When ductile or nodular cast iron is subjected to austempering heat treatment, it produces a microstructure consisting of ferrite and highcarbon austenite. This microstructure is different from that of steels. ADI has important advantages, as it requires short heat treatment periods due to its chemical composition and microstructure. Moreover, parts with intricate shapes can be cast with adequate strength and wear resistance properties similar to those of steels (with 10% lower density) [10]. In addition, ADI is better than forged aluminum with respect to its weight–strength ratio. Lerner and Kingsbury observed that ADI wear resistance is nearly four times greater than pearlite DI and more than 12 times that of leaded-tin bronze in dry-sliding wear mode [11]. The machinability of most types of cast iron involved in metal cutting production is generally good and is highly related to the structure of cast iron. Comparative machinability studies on DI samples that were austempered under different conditions have been carried out by several authors [12–15]. In these studies, the effect of heat treatment on machinability was searched by keeping the cutting parameters constant. As to our study, by keeping the cutting parameters in varying values, the effect of cutting parameters on machining of ADI and optimum cutting parameters was studied. From this aspect, it can obviously be said that this study is original. The aim of this experimental study was to investigate the effects of machining parameters such as different cutting speeds and depths of cuts on the machinability of ADI. The cutting forces, surface roughness, and tool wear were adopted as the evaluating criteria. 2 Machinability of Austempered Ductile Iron Austempering is an interrupted quenching process. Work piece must be quenched rapidly to the correct holding temperature so Journal of Manufacturing Science and Engineering C 2012 by ASME Copyright V APRIL 2012, Vol. 134 / 021013-1 Downloaded 21 Jun 2012 to 194.27.12.67. Redistribution subject to ASME license or copyright; see http://www.asme.org/terms/Terms_Use.cfm Fig. 2 Fig. 1 Isothermal transformation (IT) diagram for DI [23] that the cooling curve is not permitted to intersect the nose in the transformation diagram (Fig. 1). The presence of graphite particles in gray cast iron, malleable cast iron, and DI renders materials with good machinability according to nearly all criteria, especially when compared to steels. Low rates of tool wear, high rates of metal removal, relatively low cutting forces, and power consumption are some of the positive characteristics of cast iron [12]. The surface of machined cast iron, however, is rather matte in character. When cast iron is machined, graphite particles determine the cutting forces and surface roughness, whereas the matrix determines the tool life (T) [16,17]. The available information about the machinability of ADI and, more specifically, about the factors affecting it, is, to the best of our knowledge, very scarce and only covers specific technical cases in which the data reported have not been obtained following a particular standard [15]. The best machinability characteristics of ADI have been achieved by a matrix structure (ferrite þ austenite with high carbon content) with austempering. Although this for ADI has obtained good mechanical properties compared to DI, its enhanced mechanical properties present some machining problems. It is necessary to obtain optimum values of the cutting parameters. Ferritic cast iron with increasing silicon content is stronger and less ductile, tends to give less Built-Up Edge (BUE). While machinability is significantly affected by graphite, ferrite, pearlite, and carbide content in DI, some alloying elements improve the machinability, while others reduce it [18,19]. Although there is no definite information in the published literature that DI has better machinability than steels, data obtained from manufacturers do support this. Parts manufactured from DI show improved tool life by 20–900% [20] when compared to heat-treated forged steels. Very fine surface finishing can be obtained on ductile iron. The surface roughness of machined parts depends on the grain size of the machining medium and the applied finishing process. Surface roughness of 0.1 lm (Ra) or better on DI parts can be achieved by grinding and honing [21]. When machining ductile irons, it is possible to find some practical values for the cutting parameters from a machining handbook [21,22]. It is also possible to adopt suitable cutting parameters for ADI from those for ductile iron derivatives of ADI. It can be reasoned that ADI poses certain problems during machining due to its enhanced strength and wear resistance when 021013-2 / Vol. 134, APRIL 2012 Relative machinability of several ferrous materials [7] compared to the other types of cast iron. It has previously been reported that unreacted (or untransformed) retained austenite can reduce the machinability of ADI due to its stress-induced transformation to martensite during machining [15]. However, softer grades of ADI machinability are the same as or better than those of steels, which are similar to ADI in strength [7]. A comparison between the relative machinability of various iron-based engineering materials is shown in Fig. 2. Seah and Sharma conducted machining tests on ADI alloyed with nickel and austempered at 300, 340, 360, and 380  C for 120 min [14]. They determined the machinability based on the material removal rate and the unit power consumed at various cutting speeds and feed rates. It was found that the machinability index of this material increases significantly with an increase in austempering temperature due to a severe drop in hardness. Nickel content, however, had only a very marginal effect on hardness. Moncada et al. investigated the influence of some variables affecting the machinability of ADI in turning, using a standard testing procedure [15]. They measured the machinability for a wide range of machining variables, such as cutting speed, feed rate, depth of cut, and lubrication and some cutting tool characteristics, such as the carbide grade, rake angle, and coating. The effect of tool geometry [24,25], coating, and coolant effects on machinability were considered as well. As a result, higher austempering temperature produces lower machinability. S eker and Hasirci measured the cutting forces and surface roughness to determine the effect of microstructure on the mechanical properties of austempered ductile irons (ADIs) [12]. Specimens were prepared under different austempering times with the addition of various amounts of Cu and Ni. Six different specimen groups were austenitized at 900  C for 90 min and then austempered in molten salt at 370  C for 60, 90, 180, and 200 min. The results were evaluated after machining tests, which were carried out in accordance with ISO 3685. Austempering heat treatment resulted in considerable improvement of the surface quality when compared to as-cast specimens, while the changes in the cutting forces remained at about 20% for different specimens. In terms of both criteria, the best result was obtained from specimens austempered for 60 min with addition of 0.7% Ni and 0.7% Cu. The temperature and period of austempering are the main determinants that affect the mechanical properties of ADI. High austempering temperature generates high ductility, low yield strength with good fatigue, and impact strength. When the austempering time is kept short, there is insufficient carbon diffusion to the austenite to stabilize it, and martensite may form during cooling. The resultant microstructure would have a higher hardness but lower ductility and fracture toughness [26]. 3 Experimental Procedure 3.1 Preparation of Test Specimens. The cast DI specimen bars with ferrite and spherical graphite were produced by the horizontal continuous casting method. For experimental studies, the specimens were prepared nearly 38 mm in diameter and 254 mm Transactions of the ASME Downloaded 21 Jun 2012 to 194.27.12.67. Redistribution subject to ASME license or copyright; see http://www.asme.org/terms/Terms_Use.cfm Table 1 Element % Chemical composition of the ductile iron specimens C Si P S Mn Mg 3.80 2.87 0.030 0.010 0.18 0.054 Fig. 3 Schematic diagram of the heat treatment process in length. The chemical composition of the DI specimens is given in Table 1. Following each heat treatment, before machinability test, each specimen was turned with a 1.8 mm depth of cut to avoid possible decarburization and oxidization. 3.2 Heat Treatment. Heat treatment is an essential requirement to improve the mechanical and thermal properties of ADI. A schematic diagram of the heat treatment process is shown in Fig. 3. Temperature measurement during heat treatment was conducted with a K-type thermocouple welded to one side of the specimen. Austenitizing of the DI specimens was carried out in a muffle furnace. All specimens were treated in a heat bed of cast iron chips to minimize decarburization. After austenitizing at 900  C for 90 min, the specimens were quenched in a 50% NaNO3þ50% KNO3 salt bath at 380  C for 90 min. The mechanical properties of the specimens before and after austempering are given in Table 2. When the values in the table are closely observed, it is seen that proof strength, tensile strength, and hardness increased considerably. Hardness is a good indication of machinability, which deteriorates with increasing Brinell-hardness. Abrasive hardness due to sand inclusions and free carbides is very negative for machinability. Hence, the higher the hardness and strength that a type of cast iron has, the lower the machinability and shorter the tool life that can be expected from inserts and tools produced from that material. 3.3 Machinability Tests. Machinability tests were conducted with uncoated cemented carbide tools on a CNC lathe (Johnford TC-35) in accordance with ISO 3685. Cutting tools were selected according to ISO 1832 as recommended by ISO 3685. In machining tests, carbide inserts with positive rake angles were used. The positive-rake-angle tools achieved better machinability than negative-rake-angle tools. This result is significant, as the literature recommends the use of negative rake angles for cast iron, in general [24]. The inserts were mechanically clamped onto a rigid tool holder (SSBCR 2525M-12 according to ISO 5608). The cutting tool grade (ISCAR IC 20 that is equivalent to K10, K10/20, and K20 grade in ISO) was selected by taking into consideration the mechanical properties of the ADI. Before testing, a metal Table 3 Experimental details for the machining tests Machine tool Tool holder Tool inserts Cutting parameters Cutting speed (v) Feed rate (f) Depth of cut (a) Rake angle (c) Tool cutting edge angle (jC) Nose radius Cutting condition Johnford TC-35 CNC Lathe SSBCR 2525M-12 SCMT 120408-19 (K10, K10/20, K20) 50; 100; 150; 200; 250 m/min 0.12 mm/rev 1.0; 1.5; 2.0; 3.0 mm 6 deg 75 deg 0.4 mm Dry layer, 1 mm in depth, was removed from the surface of all heattreated bars with cylindrical turning method to avoid possible decarburization and oxidization. Thus, the form errors and adverse effects of inhomogeneous hardness on the outer layer were removed, and a homogeneous chip section for cutting force measurements was provided. The tests were carried out under dry conditions. The details of the experimental conditions are summarized in Table 3. In these tests, the cutting speed, v, depths of cut, a, and machining length, Lc, were varied in the ranges 50–250 m/min, 1–3 mm, and 50–1200 mm, respectively, while the feed rate, f (0.12 mm/rev), was kept constant. The tangential, feed, and radial cutting force components acting on the tool holder were measured with a three-component dynamometer (Kistler 9257B). Surface roughness values were measured at three noncollinear points on the machined surface using a profilometer (Taylor-Hobson Surtronik 10). During machining, the cutting process was halted at different machining lengths (50–1200 mm), and the rates of tool flank wear (VB) occurring on the tool side-face were measured periodically by optical microscopy (Scherr Tumico) following the metal removing processes. As the tool life criteria, a value of 0.3 mm of average flank wear land (VB) that was established by ISO 3685 was used (Fig. 4). Quick stop device (QSD), the schematic working principle of which is shown in Fig. 5, was used in order to examine cutting mechanism at the time of machining and the morphology of chip that came up as a result. This device can be used approximately at a 160 m/min peripheral cutting speed [27]. While turning the specimens at different cutting speeds for 2 s, some chip root samples were produced by using the QSD. Then, the metallography of these specimens was performed. 3.4 Metallographic Inspection. Standard procedures were followed in preparing the specimens. The number of graphite spheres per unit area in 1 mm2 was determined from the sample surface etched with nital by optical microscopy (LEICA DM 4000M). The average phase volume fractions were also determined from the sample surface by using the line-crossing method with measurements taken from five different zones in the field of view of the optical microscope (Table 2). Metallographic examination showed that the austemperability was sufficient in all cases to obtain typical ADI structures on the surface of the test bars. It is apparent from the micrographs that the amount of retained austenite varies depending on the austempering time and temperature. Generally, the amount of retained austenite increases with decreasing austempering time. Changes in the phase volume fractions of as-cast Table 2 Phase volume fractions, nodule count, and mechanical properties of as-cast and ADI specimens Specimens DI ADI Graphite (%) Ferrite (%) Carbon-enrichment austenite (%) Untransformed austenite (%) Nodule count (nodule/mm2) Proof strength (MPa) Tensile strength (MPa) Hardness (BHN) 14.38 10.76 85.62 42.87 — 39.04 — 7.3 235 214 358 769 454 818 151 313 Journal of Manufacturing Science and Engineering APRIL 2012, Vol. 134 / 021013-3 Downloaded 21 Jun 2012 to 194.27.12.67. Redistribution subject to ASME license or copyright; see http://www.asme.org/terms/Terms_Use.cfm Fig. 4 Typical profile of flank wear [13] Fig. 7 The variation of cutting forces with cutting speed (a 5 1 mm) Fig. 5 The schematic working principle of QSD and austempered specimens are shown in Table 2, and microstructural changes are shown in Fig. 6. 4 Results and Discussion In metal cutting, two important quality characteristics are taken into consideration: namely, the desired geometry and surface roughness (Ra). Despite the complex structure, partial control of the surface roughness can be attained with the selection of optimum values of three cutting parameters. Depending on this approach, the effects of cutting speed and depth of cut on the machinability were investigated. Focus was then turned to the determination of the cutting parameters that produce better results from the point of view of the machinability criteria. The machinability evaluation was performed with reference to the cutting forces, surface roughness, and tool wear. 4.1 Evaluation of Cutting Forces With Cutting Speed and Depth of Cut. The variation of the cutting force components measured as tangential force (Ft), feed force (Ff), and radial force Fig. 6 Fig. 8 The variation of cutting forces with depth of cut (v 5 200 m/min) (Fr) depending on different cutting speeds and depths of cuts during the machining of the ADI samples are given in Figs. 7 and 8. In this study, the cutting force components along three directions were recorded online for evaluation. When the graph showing the variation of the cutting forces with the cutting speeds is closely examined (Fig. 7), it is seen that the maximum cutting forces were (a) Micrographs of an as-cast specimen and (b) austempered ductile iron 021013-4 / Vol. 134, APRIL 2012 Transactions of the ASME Downloaded 21 Jun 2012 to 194.27.12.67. Redistribution subject to ASME license or copyright; see http://www.asme.org/terms/Terms_Use.cfm Fig. 9 Flank wear rate at different machining lengths (a 5 1 mm, v 5 250 m/min) Table 4 Variation of the cutting forces depend on the depths of cuts Variation of depth of cut, d (mm) 1.0–1.5 1.5–2.0 2.0–3.0 Raise in Ft (%) Raise in Ff (%) 73 32 34 89 44 24 measured at low cutting speeds (50 m/min). The decreasing in Ft is 47% as the cutting speed is increased from 50 to 200 m/min. The cutting forces show a decreasing trend up to a cutting speed of 200 m/min. The existence of steady BUE on the interface of tool/chip increases the active rake angle of the cutting tool. The increased rake angle increases shear plane angle, which decreases shear plane area and, therefore, cutting forces. Moreover, the existence of BUE on the chip surface of tool cuts down tool/chip contact area. For this reason, decreased friction forces cause decreasing of cutting and feed forces too [17]. The decrease in the cutting forces with increasing cutting speed can be explained by the partially reduced contact region on the cutting tool rake surface and also by the partially reduced yield strength of material with increasing temperature in the cutting zone during metal removal [28]. For this reason, the cutting forces show a rapid decrease beyond approximately 150 m/min. The observed increase in the cutting forces with increasing cutting speed from 200 m/min to 250 m/min can be attributed to the probable increase in tool wear at high cutting speeds (Fig. 9), and consequently, this leads to an increase in the cutting forces due to the wearing away of the cutting tool [16]. In order to evaluate the effect of the depth of cut on the cutting forces, the cutting speed was held constant at 200 m/min for a series of depths. The variation of cutting force with depth of cut at Fig. 10 The effect of cutting speed on surface roughness (a 5 1 mm) Journal of Manufacturing Science and Engineering v ¼ 200 m/min is shown in Fig. 8. It can be seen that the cutting forces increased with increasing depth of cut. Hence, the removed chip volume, the removed chip rate, and also the specific cutting energy, all increase with increasing depth of cut. The maximum increase was measured for tangential/main cutting force (Ft), and feed force (Ff), but the increase in the radial force was negligible. When the depth of cut was increased from 1 mm to 1.5 mm, Ft and Ff increased by 73% and 89%, respectively. The other variations in the cutting forces depending on the depth of cut are given in Table 4. 4.2 Evaluation of Surface Roughness With Cutting Speed and Depth of Cut. One of the other factors determining surface quality is surface roughness. The parameters widely used for Fig. 11 The effect of depth of cut on surface roughness (v 5 100 m/min) Fig. 12 Flank wear rate dependence at different machining lengths on cutting speed (a 5 1 mm) APRIL 2012, Vol. 134 / 021013-5 Downloaded 21 Jun 2012 to 194.27.12.67. Redistribution subject to ASME license or copyright; see http://www.asme.org/terms/Terms_Use.cfm Fig. 13 Flank wear rate dependence at different machining lengths on depth of cut (v 5 100 m/min) surface finish are average surface roughness (Ra) and maximum roughness depth (Rt). During the machining of the ADI samples, the variations in the average surface roughness at different cutting speeds and depths of cuts are given in Figs. 10 and 11. In Fig. 10, the surface roughness values with increasing cutting speed decrease until a certain cutting speed and then start to increase. At low cutting speed (50 m/min in the experiment), the maximum surface roughness was obtained, while the minimum surface roughness was obtained at high cutting speed (200 m/min). A decrease in surface roughness until 200 m/min occurs with a decrease in the BUE along the Fig. 14 Micrographs showing the morphologies of chip roots at various cutting speeds [28] cutting edge. The temperature in the cutting area increases with increasing cutting speed, and for a cutting process maintained at a raised temperature, the strength of the BUE is reduced. Therefore, BUE becomes more unstable and needs to be removed from the cutting tool more often. Because BUE formation and its separation period from the tool become shorter, its size also decreases Fig. 15 Response surface of (a) surface roughness (Ra), (b) tangential force (Ft), and (c) flank wear (VB) 021013-6 / Vol. 134, APRIL 2012 Transactions of the ASME Downloaded 21 Jun 2012 to 194.27.12.67. Redistribution subject to ASME license or copyright; see http://www.asme.org/terms/Terms_Use.cfm and, therefore, reduces the roughness at the surface. The temperature on the tool face also plays a major role with respect to the size and stability of the BUE. The surface roughness of the austempered specimens was significantly better than those of as-cast specimens [12]. When the influence of the depth of cut on the surface roughness is examined (Fig. 11), the average surface roughness is shown to increase with increasing depth of cut. The increase in cutting forces and tool wear resulting from the increase in depth of cut will negatively affect surface roughness. During the cutting process, cutting forces cause several changes in the shapes of both tool and work piece and change the location (position) of tool/ work piece and affect cutting quality. In fact, work piece-toolmachine tool chain is a flexible system. So, during the cutting process, due to the variability of cutting force, there may emerge vibrations. In case of these vibrations being intensive, chatter that causes a bad surface quality may be encountered. The observed increase in the surface roughness with increasing cutting speed from 200 m/min to 250 m/min can be attributed to the probable increase in tool wear at high cutting speeds, and consequently, this leads to an increase in the surface roughness due to the wearing away of the cutting tool. 4.3 Cutting Tool Wear. Tool wear is the product of a combination of load factors, such as mechanical, thermal, chemical, and abrasive, on the cutting edge. Tool wear is a natural result of the interaction between the tool and work piece couple and machining conditions. In metal cutting, flank wear is the most common type of tool wear and results mainly from the abrasive wear mechanism. In determining tool life, flank wear is usually the most important parameter that is to be taken into account. It affects the dimensional accuracy and surface quality. Excessive flank wear will lead to poor surface quality, inaccuracy, and increasing friction as the cutting edge changes shape. Temperature is also an accelerating factor in this process. Following the primary stage, the tool progressively wears at almost a constant rate until it reaches the final stage during which wear develops at an increasing rate, leading to a tool failure [29]. As a result of load factors, friction, and temperature, some basic wear mechanisms dominate metal cutting. The main wear types encountered when machining cast iron are abrasion, adhesion, and diffusion wear. The abrasion is produced mainly by carbides, sand inclusions, and harder chilled skins. Adhesion wear with BUE formation takes place at low cutting speeds and, consequently, at low machining temperatures. The ferrite portions of DI are welded onto the inserts easily but can be eliminated with increasing speed. On the other hand, diffusion wear is temperature-related and occurs at high cutting speeds, especially in high-strength DI grades. The dependence of the flank wear rate of the cutting tool at different machining lengths on the cutting speed and depth of cut are shown in Figs. 12 and 13, respectively. In these graphs, machining length corresponds to tool life. Cutting speed has a major effect on tool wear; thus, excessive tool Fig. 16 Response optimization plot for (a) surface roughness (Ra), (b) tangential force (Ft), and (c) flank wear (VB) parameter components Journal of Manufacturing Science and Engineering APRIL 2012, Vol. 134 / 021013-7 Downloaded 21 Jun 2012 to 194.27.12.67. Redistribution subject to ASME license or copyright; see http://www.asme.org/terms/Terms_Use.cfm Table 5 Response optimization for surface roughness (Ra), tangential force (Ft), and flank wear (VB) parameter components Obtimum cutting parameters Parameters Ra (lm) Ft (N) VB (mm) Goal a (mm) v (m/min) Lower Target Upper Pre. response Desirability Minimum Minimum Minimum 1 1 1 185.35 167.17 50 1.079 243.51 0.086 1.079 243.51 0.086 3.7 747.62 0.654 0.8836 266.74 0.104 1 0.95392 0.96881 wear occurred at the cutting speed of 200 m/min, while less tool wear occurred at low cutting speed, namely 50 m/min and 100 m/min. While the increase in the flank wear of the cutting tool with cutting speed shows approximately uniform character up to a machining length of 200 mm, beyond a machining length of 300 mm, the maximum wear limit for flank wear is exceeded due to the temperature generated in the cutting area. By decreasing the cutting speed from 200 m/min to 50 m/min, it was observed that the tool life rose to a maximum level. In this study, the temperature at the tool-chip interface was not measured; instead, the effects of temperature on cutting edge were analyzed. The interaction between tool wear and depth of cut shows a decreasing trend. The tool wear increases with increasing depth of cut (Fig. 13), but this effect is not as significant as compared with the effect of cutting speed on tool wear. 4.4 Chip Root Morphologies and Built-Up Edge Formation. The factors affecting chip morphology are related to the chemical composition of the material, microstructure, and metal-removal parameters [30]. Micrographs depicting the morphologies of chip roots formed at various cutting speeds are shown in Fig. 14. BUE is existent at maximum size at 20 m/min, which is the lowest cutting speed used in experiments [Fig. 14(a)]. With the increasing cutting speed, BUE dimension gets smaller. BUE is about to get disappear [Fig. 14(d)]. At low cutting speeds, irregularity on the machined surfaces of the samples is observed. Besides, chip that emerges at low cutting speeds is irregular, and shear plane angle is wide. It is a noticeable fact that, with the increasing cutting speed, shear plane angle narrows, and although chip should come up more regularly, even at high cutting speeds, chip that emerges is irregular. The reason for that is, as can clearly be seen in Fig. 14, probably the overlapping of spheroidal graphite by breaking off and their displaying of a solid lubricant feature. 4.5 Optimization of Response. Response surface methodology (RSM) is a statistical technique based on simple multiple regressions. The effect of two or more factors on quality criteria can be investigated, and optimum values are obtained by utilizing this technique [31]. In RSM design, there should be at least three levels for each factor. In this way, the factor values that are not actually tested using fewer experimental combinations and the combinations themselves can be estimated [32]. The results are expressed in 3D series or counter map. The effect of different cutting parameters, cutting speed, and depth of cut (v, a) on surface roughness (Ra), tangential force, and flank wear (VB) are shown in Fig. 15. Figure 15(a) slightly displays that the value of surface roughness (Ra) increases with increase in the depth of cut, a. It goes down at first and then recovered gradually while the cutting speed v increases. In Fig. 15(b), as the depth of cut, a, increases, the tangential force (Ft) increases. However, it is flactuated while the cutting speed v increases. From Fig. 15(c), it is observed that if cutting speed and depth of cut (v, a) all together are increased, the flank wear (VB) increases too. One of the most important aims of experiments related to manufacturing is to achieve the desired surface roughness, tangential force, and flank wear of the optimal cutting parameters [33]. To this end, the response surface optimization is one of the techniques for determination of the best cutting parameters in turning. Here, 021013-8 / Vol. 134, APRIL 2012 the goal is to minimize surface roughness, tangential force, and flank wear. RSM optimization result for surface roughness, tangential force, and flank wear parameter are shown in Fig. 16. Optimum cutting parameters obtained in Table 5 are found to be the depth of cut (a) of 1 mm and the cutting speed (v) of 185.35 m/min for surface roughness (Ra). They are realized for the tangential force (Ft) and flank wear (VB) as 1 mm, 167.17 m/min and 1 mm, 50.00 m/min, respectively. 5 Conclusion In the present study, specimens were first prepared by austenitizing DI at 900  C for 90 min and were then quenched in a 50% NaNO3þ50% KNO3 salt bath at 380  C for 90 min. Then, the effects of cutting parameters on the machinability of ADI were investigated, and the measured outputs, such as cutting forces, surface roughness, and tool wear, were analyzed. By selecting convenient tool material, geometry, and machining conditions, ADI specimens were machined using a cutting speed in the range of 50–250 m/min, depth of cut in the range of 1–3 mm, and a constant feed rate of 0.12 mm/rev for a machining length of 1200 mm. The following conclusions have been drawn from the present study: 1. The cutting forces were measured along three directions. When the variation of the cutting forces with cutting speed was investigated, it was seen that the maximum cutting force was measured at lower cutting speed (50 m/min), whereas minimum force was obtained at the 200 m/min cutting speed. The observed increase in the cutting forces with increasing cutting speed from 200 m/min to 250 m/min can be attributed to the probable increase in tool wear at high cutting speeds, and consequently, this leads to an increase in the cutting forces due to the wearing away of the cutting tool. 2. The cutting force components increased proportionally with increasing rate at depths of cuts. When the depth of cut was increased from 1 mm to 1.5 mm, Ft and Ff increase by 73% and 89%, respectively. 3. Surface roughness is an important factor that determines surface quality. It varies mainly with the machining method, type and tool condition, cutting parameters (especially tool tip radius and feed rate), work piece material, and overall stability. The maximum surface roughness was obtained at the lowest cutting speed used in the experiment. BUE formation has a negative effect on surface quality such that the machined surface deteriorates and the temperature is increased. 4. When the depth of cut is increased, its influence on surface roughness is negative, but this effect is negligible. The increase in cutting forces and tool wear resulting from the increase in depth of cut will negatively affect surface roughness. 5. Cutting speed has a major effect on tool wear; thus, excessive tool wear occurred at the cutting speed of 200 m/min, while less tool wear occurred at low cutting speed, namely 50 m/min and 100 m/min. The tool wear increases with increasing depth of cut, but this effect is not as significant as compared with the effect of cutting speed on tool wear. 6. In the experiments performed, it was observed that spheroidal graphite overlap by breaking off in the processing of austempered ductile iron, and they make the chip to emerge unstably by displaying a solid lubricant feature. Transactions of the ASME Downloaded 21 Jun 2012 to 194.27.12.67. Redistribution subject to ASME license or copyright; see http://www.asme.org/terms/Terms_Use.cfm 7. Optimum cutting parameter values are obtained by utilizing RSM. They are found to be the cut of depth (a) of 1 mm and the cutting speed (v) of 185.35 m/min for surface roughness (Ra). They are realized for the tangential force (Ft) and flank wear (VB) as 1 mm, 167.17 m/min and 1 mm, 50.00 m/min, respectively. Acknowledgment This study was carried out as M.Sc. Thesis by S akir Yazman in the Graduate School of Natural and Applied Science, University of Selcuk, Konya, 2006 (in Turkish). References [1] Janowak, J. F., and Gundlach, R. B., 1983, “Development of a Ductile Iron for Commercial Austempering,” Trans. Am. Foundry Soc., 54(83), pp. 377–388. [2] Shea, M. M., and Ryntz, E. F., 1986, “Austempering Nodular Iron for Optimum Toughness,” Trans. Am. Foundry Soc., 86(125), pp. 683–688. [3] Ozel, A., 1994, “Investigation of the Effect of Austempering Heat Treatment on the Impact Strength and Impact Transition Temperature in GGG 40-80 Class Ductile Irons,” PhD thesis, ITU-Institute of Science and Technology, Istanbul, Turkey. [4] Trudel, A., and Gagne, M., 1997, “Effect of Composition and Heat Treatment Parameters on the Characteristics of Austempered Ductile Irons,” Can. Metall. Q., 36(5), pp. 289–298. [5] Putatunda, S. K., and Gadicherla, K., 2000, “Effect of Austempering Time on Mechanical Properties of a Low Manganese Austempered Ductile Iron,” J. Mater. Eng. Perform., 9(2), pp. 193–203. [6] Rafaey, A., and Fatahalla, N., 2003, “Effect of Microstructure on Properties of ADI and Low Alloyed Ductile Iron,” J. Mater. Sci., 38, pp. 351–362. [7] Delia, M., Alaalam, M., and Grech, M., 1998, “Effect of Austenitizing Conditions on the Impact Properties of an Alloyed Austempered Ductile Iron of Initially Ferritic Matrix Structure,” J. Mater. Eng. Perform., 7(2), pp. 265–272. [8] Gagne, M., 1985, “The Influence of Manganese and Silicon on the Microstructure and Tensile Properties of Austempered Ductile Iron,” Trans. Am. Foundry Soc., 85(133), pp. 801–812. [9] Bosnjak, B., Radulovic, B., Pop-Toner, K., and Asanovic, V., 2001, “Influence of Microalloying and Heat Treatment on the Kinetics of Bainitic Reaction in Austempered Ductile Iron,” J. Mater. Eng. Perform., 10(2), pp. 203–211. [10] Degarmo, E. P., Black, J. T., and Kahser, R. A., 1997, Materials and Process in Manufacturing, 8th ed., Prentice-Hall, NJ. [11] Lerner, Y. S., and Kingsbury, G. R., 1998, “Wear Resistance Properties of Austempered Ductile Iron,” J. Mater. Eng. Perform., 7(1), pp. 48–52. [12] S eker, U., and Hasirci, H., 2006, “Evaluation of Machinability of Austempered Ductile Irons in Terms of Cutting Force and Surface Quality,” J. Mater. Process. Technol., 173(3), pp. 260–268. Journal of Manufacturing Science and Engineering [13] Cakir, M. C., and Isik, Y., 2008, “Investigating the Machinability of Austempered Ductile Irons Having Different Austempering Temperatures and Times,” Mater. Des., 29, pp. 937–942. [14] Seah, K. H. W., and Sharma, S. C., 1995, “Machinability of Alloyed Austempered Ductile Iron,” Int. J. Mach. Tools Manuf., 35(10), pp. 1475–1479. [15] Moncada, O. J., Spicacci, R. H., and Sikora, J. A., 1998, “Machinability of Austempered Ductile Iron,” Trans. Am. Foundry Soc., 98(10), pp. 39–45. [16] Ductile Iron Data for Design Engineer, 2001, Ductile Iron Society (DIS), Montreal, Canada. [17] Trent, E. M., 1984, Metal Cutting, Tanner Ltd., London. [18] Modern Metal Cutting—A Practical Handbook, 1997, Sandvik-Coromant Co. Inc., Sweden. [19] Die and Mould Design Handbook, 2000, Sandvik-Coromant Co. Inc., Sweden. _ and Hasirci, H., 2003, “The Effect of Alloying Elements on [20] S eker, U., C iftçi, I., Surface Roughness and Cutting Forces During Machining of Ductile Iron,” Mater. Des., 24, pp. 47–51. [21] The Iron Casting Handbook, 1981, Iron Casting Society, Inc., Ohio, USA. [22] Machining Ductile Irons, 2001, International Nickel Co. Inc., New York. [23] ASM Handbook, 1991, Heat Treating, 10th ed., Vol. 4, ASM International, Ohio, USA. [24] Saglam, H., Unsacar, F., and Yaldiz, S., 2006, “Investigation of the Effect of Rake Angle and Approaching Angle on Main Cutting Force and Tool Tip Temperature,” Int. J. Mach. Tools Manuf., 46, pp. 132–141. [25] Saglam, H., Yaldiz, S., and Unsacar, F., 2007, “The Effect of Tool Geometry and Cutting Speed on Main Cutting Force and Tool Tip Temperature,” Mater. Des., 28, pp. 101–111. [26] Sorelmetal, 1990, Ductile Iron Data for Design Engineer, Ductile Iron Society, pp. 84–155. [27] Özçatalbas, Y., and Ercan, F., 2003, “The Effects of Heat Treatment on the Machinability of Mild Steels,” J. Mater. Process. Technol., 136, pp. 227–238. [28] Yazman, S., 2006, “The Effects of Cutting Parameters on Machining for Austempered Ferritic Ductile Iron,” (in Turkish) MSc thesis, Selcuk UniversityInstitute of Science and Technology, Konya, Turkey. [29] Oraby, S. E., and Hayhurst D. R., 1991, “Development of Models for Tool Wear Force Relationships in Metal Cutting,” Int. J. Mech. Sci., 33(2), pp. 125–138. [30] Dubensky, W. J., and Rundman, K. B., 1985, “An Electron Microscope Study of Carbide Formation in Austempered Ductile Iron,” Trans. Am. Foundry Soc., 64(85), pp. 389–394. [31] Neseli, S., Yaldız, S., and Türkes, E., 2011, “Optimization of Tool Geometry Parameters for Turning Operations Based on the Response Surface Methodology,” Measurement, 44, pp. 580–587. [32] Kwak, J. S., 2005, “Application of Taguchi and Response Surface Methodologies for Geometric Error in Surface Grinding Process,” Int. J. Mach. Tools Manuf., 45, pp. 327–334. [33] Bouacha, K., Yallese, M. A., Mabrouki, T., and Rigal, J. F., 2010, “Statistical Analysis of Surface Roughness and Cutting Forces Using Response Surface Methodology in Hard Turning of AISI 52100 Bearing Steel With CBN Tool,” Int. J. Refract. Met. Hard Mater., 28, pp. 349–361. APRIL 2012, Vol. 134 / 021013-9 Downloaded 21 Jun 2012 to 194.27.12.67. Redistribution subject to ASME license or copyright; see http://www.asme.org/terms/Terms_Use.cfm View publication stats