Wear mechanisms of PVD-coated cutting tools during continuous turning of Ti-6Al-4V alloy

Accepted Manuscript Title: Wear Mechanisms of PVD-Coated Cutting Tools During Continuous Turning of Ti-6Al–4 V Alloy Author: Shuho Koseki Kenichi Inoue Katsuhiko Sekiya Shigekazu Morito Takuya Ohba Hiroshi Usuki PII: DOI: Reference:

S0141-6359(16)30221-5 http://dx.doi.org/doi:10.1016/j.precisioneng.2016.09.018 PRE 6465

To appear in:

Precision Engineering

Received date: Accepted date:

14-9-2016 29-9-2016

Please cite this article as: Koseki Shuho, Inoue Kenichi, Sekiya Katsuhiko, Morito Shigekazu, Ohba Takuya, Usuki Hiroshi.Wear Mechanisms of PVD-Coated Cutting Tools During Continuous Turning of Ti-6Al–4V Alloy.Precision Engineering http://dx.doi.org/10.1016/j.precisioneng.2016.09.018 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.

Wear Mechanisms of PVD-Coated Cutting Tools During Continuous Turning of Ti-6Al-4V Alloy Shuho Koseki*, Kenichi Inoue**, Katsuhiko Sekiya†, Shigekazu Morito†, Takuya Ohba††, and Hiroshi Usuki ††† |1 Shuho Koseki* Hitachi Metals, Ltd. Metallurgical Research Laboratory *Corresponding author 22 banchi, Hokuryou-cho, Matsue-shi, Shimane-ken, 690-0816 Japan TEL: +81-852-60-5050 (office), +81-90-1123-0758 (cell) FAX: +81-852-60-5055 E-mail: [email protected] Kenichi Inoue** Hitachi Metals, Ltd. High-Grade Metals Company, Engineering Department Shinagawa Season Terrace, 2-70, Konan 1-chome, Minato-ku, Tokyo 108-8224, Japan Katsuhiko Sekiya† Hiroshima University, Graduate School of Engineering, Department of Mechanical Systems Engineering 1-4-1 banchi, Kagamiyama, Higashi-Hiroshima-shi, Hiroshima-ken 739-8527, Japan Shigekazu Morito, Takuya Ohba†† Shimane University, Department of Materials Science 1060 banchi, Nishikawatsu-chou, Matsue-shi, Shimane-ken 690-0823, Japan Hiroshi Usuki††† †Shimane University, Department of Materials Creation and Circulation Technology 1060 banchi, Nishikawatsu-chou, Matsue-shi, Shimane-ken 690-0823, Japan

Highlights 

Damage of PVD-coated cutting tool during titanium alloy turning was

investigated. 

TiN coating on cutting tool wore faster than actual cemented carbide tool.



Damage mode of the coating was fracture without plastic deformation.



Coating damage dependent on interfacial strength with adhered materials.



Diagram for explaining coating damage during titanium alloy turning is proposed.

Abstract: The aim of this study was to investigate the damage of cutting tools coated by physical vapor deposition (PVD) during the continuous turning of a titanium alloy. The investigation utilized scanning electron microscopy (SEM), electron probe micro-analysis (EPMA), and transmission electron microscopy (TEM). It was found that a TiN coating on the tool wore faster than an uncoated cemented carbide tool. The damage mode of the coating on the rake face was fracture without plastic deformation. Additionally, there was a pattern to the crystal orientation relationship at some of the interfaces between the adhered workpiece material and the TiN coating. The crystal orientation relationship presumably produced a strong bond between the adhered material and the coating. The coating damage was thus caused by the force exerted by the adhered materials on the grain boundary on the damaged coating surface. A comparison of the tool damages during the machining of Ti-6Al-4V alloy and alloy 718 suggested that the damage of the coating depended on the interfacial strength between the adhered material and the coating, as well as the strength of the adhered material at a high temperature. Hence, to prevent the damage of the tool during the machining of a titanium alloy, it is preferable to use a ductile material (e.g., cemented carbide) rather than a brittle material (e.g., ceramic). Key words: Ti-6Al-4V alloy, continuous turning, PVD-coated cutting tool, damage, adhesion, adhesive wear, alloy 718

|2

1. Introduction Titanium alloy is a typical difficult-to-cut material [1,2]. During the machining of the alloy, it is necessary to maintain a cutting speed lower than that used for carbon steels. Many works attribute the poor machinability to the high temperature generated at the cutting edge due to the low heat conductivity and low density of the alloy [3], and the high stress on the tool face due to the short chip-tool contact length [4]. The poor machinability may also be due to the fluctuation of the cutting force caused by the saw-tooth shape of the chip resulting from adiabatic plastic deformation [3,5,6], the chatter caused by the fluctuation, and the low Young’s modulus of the alloy [5]. Many engineers are of the opinion that an effective method for lengthening the life of a tool used to machine a titanium alloy is the maintenance of a sufficiently low cutting temperature. The conditions under which the cutting temperature is sufficiently low have been proposed [7], as well as the use of high-heat-conductivity tools to enhance heat dissipation [2,8,9]. The tool wear is dominated by the chipping of the cutting edge due to the adhesion between the tool and the workpiece material [1,8,10]. It has also been suggested that the adhesion of the chip to the tool face is caused by the high chemical affinity between the tool material and the chemically activated titanium alloy at high temperatures [1]. Observations during machining and welding experiments have also led to the proposition that the generation of titanium carbide at the tool-chip interface is due to the interface adhesion between the alloy and the tool material such as diamond or cemented carbide [8,11–13]. The chipping is caused by the detachment of the adhered material due to the frictional force generated by the chip flow. Cemented carbide tools with a ceramic coating have been generally found to have long service lives. However, their lives are almost the same compared to uncoated carbide tools when used for the turning of titanium alloys [2]. In many

|3

machine shops, uncoated carbide tools are mainly used for the machining of titanium alloys. The relatively poor performance of a coated tool for the machining of a titanium alloy is probably due to the adhesion of the alloy to the coated layer of the tool. There, however, have been a few reports about the detailed microsco pic observation of phenomena that occur at the interface between the chip and the coated layer. The aim of the present study was to determine the primary factor that controlled the wear of a coated tool during the machining of a titanium alloy. Ti-6Al-4V alloy was specifically used for the experiments of this study, and the wear of the coated layer was microscopically observed with the purpose of establishing a more suitable tool material and cutting method compared to present practice. Both coated and uncoated tools were examined and the observations were compared with those of our previous work [14], which considered the machining of a nickel-based super alloy.

2.

Experimental Setup

2.1 Cutting conditions The workpiece material was an annealed Hv320 Ti-6Al-4V alloy, which is a typical  type alloy. Fig. 1 shows the microstructure of the workpiece material. The material had an equiaxed α structure and crystal grain size of approximately 20–100 μm. The actual workpiece was a 100-mm-diameter cylinder. The cutting tool was fabricated from cemented carbide (WC with a grain size of about 0.8 μm and 6 mass% Co-Cr). The surface of the cutting edge was coated with approximately 5-µm-thick TiN. The coating was deposited by PVD cathodic arc ion plating. TiN coating was selected for this study because it is one of the typically employed PVD coating material. Fig. 2 shows a cross-sectional image of the cemented carbide and coating. The cutting tool geometry was ISO-compliant, with a CNMG 120408 turning insert, a

|4

chip breaker, and a rake angle of 13°. The roundness of the cutting edge was approximately 30 μm. The inserts were fixed to a tool holder with an approach angle of −5°, and back and side clearance angles of 6°. The actual rake angle of the tool after mounting on the holder was therefore approximately 7°. The cutting conditions are listed in Table 1. The lateral surface of the workpiece was machined with a cutting speed V of 40 and 60 m/min, respectively. These speeds are commonly employed in industry for machining. The cutting test was carried out by a CNC lathe machine (LB4000EXY, Okuma Corporation) under wet conditions. The wear morphologies for the two speeds did not differ. A diluted water-soluble cutting oil (Daphne Master Cool WT, Idemitsu Kosan Co., Ltd.) was employed. The feed f and depth of cut d were 0.2 mm/rev and 0.5 mm, respectively. The cutting length was 10 m, which was sufficiently long to observe the damage of the PVD-coated cutting tool. The damage of the uncoated tool during the cutting test using a cutting speed V of 60 m/min was also observed and compared with that of the PVD-coated tool. 2.2 Observation and analysis of damage modes After the turning test, the surface and cross-sectional damages of the cutting tools were observed by scanning electron microscopy (SEM) and backscattered electron imaging (BEI). The contrasts in the BEI image were determined by the composition (average atomic number of the materials). The dark parts in the images indicated the presence of contrasting substances such as light elements. Electron probe microanalysis/wavelength dispersive X-ray spectrometry (EPMA/WDS) was also used to investigate the distribution and components of the adhered material. Figure 3 shows a diagram of a sample cross-sectional observation of the cutting tool edge. The observed sample was prepared by cutting it at the center of the cutting edge, which is near the center of the chip width in the chip-tool contact area. It was considered that the contact force between the tool surface and the back surface of the chip at this point produced a stronger adhesion than elsewhere. The normal and shear stresses at this point were higher than those at points perpendicular to the chip flow direction. The point

|5

was thus more suitable for analysis of the wear. The cut surface was polished to a mirror finish using the slurry of 1-μm-grain diamond. To observe the damaged TiN coating and investigate the mechanism of the adhesion at the interface between the workpiece material and the coating, a transmission electron microscope (TEM) and scanning transmission electron microscope (STEM) were employed. The TEM sample was prepared by a focused ion beam (FIB) microsampling technique. A bright-field TEM image was used to observe the crystalline morphology of the coating and adhered materials. The contrasts in the dark-field STEM image were influenced by the composition of the material. The structure of the interface between the damaged TiN coating and the adhered material was also evaluated by high-resolution TEM observation. Before the observation, the TEM sample was carefully polished by ion milling using a low acceleration voltage. This suppressed the effect of the texture changes and prevented selective material removal. In addition, we investigated the compositions of the adhered material and conducted a structural analysis using energy dispersive X-ray spectroscopy (EDS), selected-area diffraction (SAD), and nanobeam diffraction (NBD). The analytical conditions are presented in Table 2.

3.

Results and discussion

3.1 Observation of the damaged TiN coating Figure 4 shows the results of the SEM observation, and the compositional BEI images of the rake face of the cutting edge after Ti-6Al-4V alloy turning (cutting speed V = 40 m/min, cutting length L = 10 m). Although the cutting length of 10 m was short, the region where the crater wear appeared to begin was observed slightly away from the tip of the cutting edge. The initial crater wear morphology reflected the progress of the local wear on the rake face. It indicated the occurrence of local wear over the entire chip contact area, rather than a mere loss of the coating. After the coverage of a long cutting distance, the wear region grew to form a large crater. The maximum depth of the wear was observed at a distance of approximately 2/3 (180 μm) of the chip contact length,

|6

from the tip of the cutting edge. The compositional BEI image revealed the disappearance of the TiN coating and exposure of the cemented carbide in the area of the maximum wear (bright region). The adhered material (gray region in the compositional BEI image) was observed on the cutting edge near the tip. It has been reported that the mechanical load during the machining of the Ti alloy is higher near the cutting edge, while the thermal load is highest at some distance from the cutting edge [5], as shown in Fig. 4(c). In addition, some works on the stress distribution generated on the rake face posit that the normal stress increases as the cutting edge is approached, while the shear stress remains constant at a high value from the cutting edge to the middle of the chip contact length, after which it gradually decreases towards the vicinity of the chip detachment point [15,16]. The position of the maximum wear depth probably corresponds to the location where the shear stress is higher than the normal stress and the cutting temperature is highest. Figure 5 shows the enlarged SEM images and compositional BEI images of the coating lesions in regions A and B in Fig. 4(a). Region A is located closer to the cutting edge, while region B is near the position where the chip detaches from the rake face. In the SEM images (upper photos in Fig. 5), the yellow lines indicate the boundaries of the adhesion areas of the chip material, the TiN coating, or the cemented carbide matrix. These areas are also identified by the compositional images (lower photos in Fig. 5). The material that adheres to the coating surface appears to be stretched by the chip flow. The fracture morphology of the TiN coating can be observed, while the delamination at the interface between the coating and the cemented carbide matrix is hardly discernible. This indicates that the adhesion strength of the TiN coating to the matrix is sufficiently strong for the cutting tool. The morphology of the damaged part of the coating indicated brittle fracture. The TEM image is shown in Fig. 10. It was thus concluded that the damage was brittle fracture without plastic deformation. Especially in region B, some parts of the adhered material were observed to bite into the damaged coating. It is considered that the high wear rate of the coating was caused by a chain of damage processes induced by the

|7

adhesion and loss of the adhered material. The damage rate was also presumably increased by adhesive wear during the machining. The destruction of the coating caused by the biting of the adhesive material has been previously observed in the damage of the flank face during the machining of alloy 718 [14,17]. In the case of machining involving easy adhesion of the workpiece material to the cutting tool, the destruction of the coating might be initiated by defects such as those caused by droplets of the coating material. Figure 6 shows SEM images of the damaged flank face after the turning of Ti-6Al-4V alloy. As with the chip detachment area around B in Fig. 4(a), the flank face also has a brittle fracture morphology. In addition, no smooth wear morphology was observed on both faces during Ti-6Al-4V alloy machining. It was confirmed that the coating was damaged around the lower end of the adhered material. This was because the material flow pulled the adhered material downwards, resulting in the generation of bending stress in the adhered material and coating. If the stress is smaller than the adhesion strength at the interface but greater than the strength of the coating, the adhered material would pull off the coating. This is the determinant of the morphology of the damaged coating in Fig. 6. The damage of both the rake and flank faces may be due to the falling off of the adhered material. The higher the cutting pressure and temperature—which dominate the adhesion strength at the interface—the higher the wear rate of the rake face. Ti-6Al-4V alloy exhibits high work hardening and thermal softening [18]. Moreover, the thermal softening of the Ti-6Al-4V alloy, which contained a small amount of hard particles within its matrix, hardly contributed to the abrasive wear of the flank face. Figure 7 shows the WDS mapping of the rake and flank faces after the turning of Ti-6Al-4V alloy over a length of 10 m. Ti, Al, and V were detected in these areas. The distribution of the adhered material was represented by Al, which was present in only the adhered material. The distribution of the cemented carbide substrate was identically represented by W. The observation of O emission also indicated the presence of oxygen, which may have also affected the tool damage, e.g., through the enhancement of corrosive wear processes such as oxidation wear. It could also have reduced the wear

|8

through the generation of oxide belag. The composition of the adhered material was almost the same as that of Ti-6Al-4V alloy, and the material adhered evenly over the crater wear area. The slight amount of oxygen detected, as shown in Fig. 7, indicates that oxidation was not the main cause of the coating damage. Furthermore, W was not detected on the flank face, as shown in Fig. 7(b), indicating that the coating on the cemented carbide did not peel off; rather, there was fracturing of the coating layer. The distribution of the adhered material on the flank face was very similar to that on the rake face. Figure 8 shows the cross-sectional SEM images of the cutting edge after the turning of Ti-6Al-4V alloy. As can be observed, the coating on the flank face and on the surface near the cutting edge was retained, and wear can be hardly observed on the flank face. Although there was a little chipping at the tip of the cutting edge, its presence before the preparation of the cross-section sample for SEM observation could not be confirmed. The chipping possibly occurred during the sample preparation. The coating on the rake face disappeared during the initial wear stage, exposing a cemented carbide area with planar damage. The damage indicated a very slow wear rate of the cemented carbide substrate. An uncoated cemented carbide tool is thus considered to be more suitable than a coated one for machining Ti alloys. Adhered material was confirmed only on the coating surface. It is presumed that the adhered material dropped off with the broken coating. Figure 9 shows the WDS mapping of the cutting edge cross section of the TiN-coated cutting tool after the turning of Ti-6Al-4V alloy. The analyzed area corresponds to the wear propagation areas shown in Fig. 8(c)-C. The adhered material was observed to be densely attached to the coating without any gap. Interfacial enrichment of specific elements (including oxygen) and the reaction layer was not confirmed. However, oxygen was slightly detected on the surface layer, probably adsorbed after the turning, and had caused partial oxidation of the surface. The oxidation did not directly affect the coating damage during continuous turning. 3.2 Evaluation of adhesion by TEM

|9

Figure 10(a) shows the cross-sectional TEM image of the TiN coating on the rake face after the turning of Ti-6Al-4V alloy. The observed area in the figure corresponds to that shown in Fig. 8(c)-C (the crater wear progress area). Fig. 10(b) shows the cross-sectional TEM image of the damaged coating during the turning of alloy 718 (V = 30 m/min, f = 0.2 mm/rev, d = 0.5 mm, cutting length L = 10 m). The two images show that the wear mechanism of the coating during the turning of a titanium alloy is not plastic deformation as in the machining of nickel alloy, but brittle fracture. In our previous work [19], the average cutting temperatures during the cutting of Ti-6Al-4V alloy and alloy 718 were 674 and 684 °C, respectively, as determined by the tool-workpiece thermocouple method. This means that the tool surface was exposed to almost the same temperature range during the machining of the two materials. The difference between the corresponding wear mechanisms was thus not due to differing cutting temperatures. Additionally, adhered material was observed above the damaged portion of the coating in the case of alloy 718, as shown in Fig. 10(b), while no adhered material was observed above the fractured portion of the coating for Ti-6Al-4V alloy, as can be observed from Fig. 10(a). Hence, if the machining of the Ti-6Al-4V alloy is continued, the adhesion and falling off of the adhered material is expected to be repeated. It is thus believed that the adhered material constantly occurred on the damaged coating during the turning of both alloys. To clarify the damage mechanisms of the coating, it is necessary to establish the behavior of the material that directly adheres to the damaged portion under high-temperature cutting conditions. Figure 11 shows a high-magnification cross-sectional TEM image of the coating portion with the adhered material. The particle size of the adhered material was observed to be 100 nm or less, with the material distributed uniformly in the thickness direction. Similar to the formation of the ultra-fine grains observed in the surface layer of drilled carbon steel [20], the mechanism of the fine grain formation is thought to be an extension of grain subdivision [21]. Although the miniaturization of the crystal grains significantly increases the yield

| 10

strength, the resultant increase in the tensile strength is small and work hardening is less likely to occur. It has been reported that the elongation of the material is limited by the plastic instability [22]. The characteristics of the deformation stress during the cutting were investigated by Obikawa et al. [18] through the conduction of a shock compression test. They observed increases in the work hardening and yield stress of Ti-6Al-4V alloy, substantial thermal softening, and a small fracture strain. A workpiece material with a fine microstructure that adheres to the tool cutting edges is thus considered to have a small elongation and low fracture strength, which implies that its properties affect the wear mechanism of the cutting edge. Incidentally, the tensile strength of Ti-6Al-4V alloy decreases sharply at high temperatures of 500 °C or more, whereas the elongation increases to as much as 120% [23]. The elongation was, however, particularly small with respect to the structure of the adhered material observed on the tool surface after cutting. Figure 12 shows the results of selected-area diffraction (SAD) (140 nm) and nanobeam diffraction (NBD) at the interface between the adhered material and the TiN coating. In this paper, the hcp structure miller index is defined by three indices {hkl}, which are identical to {hkil}, i = -{h + k}. The coating (analysis area a) was observed to have the pattern of a single crystal. In contrast, the adhered material (analysis area b) had a polycrystalline pattern that included a large number of fine crystals, which were non-uniformly orientated. The results of the nanobeam diffraction analysis of the crystal grains of the adhered material (analysis points 1, 2, and 3) could be indexed as -Ti (hcp). Furthermore, the particles of the adhered material exhibited different diffraction patterns and were randomly orientated. It has been suggested that the local cutting temperature in °C is 1.2–1.4 times the average temperature in the contact area between the chip and the rake face [24,25], based on which the highest local cutting temperature at the cutting edge during the machining for the Ti-6Al-4V alloy in the presented study was estimated to be 808–943 °C. The estimation is, however, considered to be higher than the actual highest cutting temperature. This is because the turning was performed under the wet condition. Based on the  transformation point of Ti-6Al-4V alloy, which is

| 11

940–1010 °C [26], the actual highest cutting temperature is estimated to be within the stable region of the phase of the alloy. In addition, the TiN coating (analysis area a in Fig.11) could be indexed as TiN (fcc). Figure 13 shows a high-magnification dark-field STEM image of the interface between the adhered material and the TiN coating. The observed area corresponds to Fig. 11-F. The adhered material was observed to have an equiaxed crystal morphology, with the size of the material particles in the microcrystals being approximately 50 nm. Fig. 14 shows the results of the nanobeam diffraction (NBD) analysis and the blight field TEM image of the interface between the adhered material particles and the TiN coating. The diffraction patterns were obtained from the numbered positions in Fig. 13. The crystal structure of the adhered material at points 1, 6, 7 and 8, which were placed about 10 to 50 nm apart from the interface, could be indexed as -Ti (hcp). Points 2 and 3 in Fig. 11, which were placed 200 nm apart from the interface, were also indexed as -Ti. The crystal orientation relationship between the adhered material and the TiN coating was examined by analyzing the crystal orientation based on each of the diffraction patterns shown in Fig. 14(a). The diffraction pattern at the interface (point 5) was a superposition of TiN (point 4) and -Ti (point 6). This shows that 020TiN and 100-Ti spots were closely placed as indicated circle in the figure. Hence, {020}TiN were roughly parallel to {100}-Ti, but approximately 6-7° tilted. When the positions were shifted to points 7 and 8, 002-Ti spot could always be observed. Hence, {200}TiN and {002}-Ti were almost parallel. The crystal orientation relationship can thus be summarized as follows: {200}TiN and {002}-Ti are almost parallel with 6-7° tilted and rotated around [100]TiN or [001]-Ti axis. The above observations indicate that the α-Ti particles at analysis points 6, 7, and 8 were attached with a slight rotation about the [001]-Ti axis, indicating a crystal orientation relationship between the coating and the particles of the adhesive material just above it. Similar analysis results have been obtained in other analyzed areas. The continuities of the lattice fringes of {100}-Ti and {010}TiN were also observed in the

| 12

blight field TEM image near point 5 (shown in Fig. 14(b)). The results indicate the existence of a crystal orientation relationship between the adhered material and the TiN coating. Although the interplaner spacing of {001}-Ti is different from that of the {001}TiN coating, the observed crystal orientation relationship between {001}-Ti and {001}TiN was as indicated in Fig. 14. In the case of the cold rolling of -Ti and Ti alloys, the {001}-Ti normal is oriented parallel to the stress direction [27, 28]. Because the cutting temperature in the present study was lower than the  transition temperature, an orientation relationship similar to that of cold rolling may be expected; that is, the {001}-Ti normal would be parallel to the stress direction. Further study is required to explain the relationship between the strong interfacial bond and the crystal orientation. The observations indicated that there is crystal orientation relationship of coating surface and adhered material, and thus suggested strong adhesion with orientation relationship. The crystal orientation relationship between the adhered material and the coating has not been observed in the cutting of alloy 718 [14,17], and is thus considered to be unique to the cutting of Ti-6Al-4V alloy. The presence or absence of the crystal orientation relationship is likely to affect the damage to the cutting tool. Figure 15 shows the cross-sectional TEM and STEM images of the damaged portion of the coating in Fig. 10-E. A gap (crack) was observed along the grain boundaries of the columnar crystals of the TiN coating in the TEM image. The tiny adhered material was also observed at the beginning of the crack due to high magnification TEM image (Fig. 10(a)-E). From the surface and cross-sectional observations described above, the crack was due to the workpiece material adhering to the grain boundary on the damaged surface of the coating. The chip flow caused the adhered material to exert a force along the surface, and this initiated the crack along the grain boundary. The adhered material was also compressed by the cutting force and acted as a wedge that promoted the cracking. Table 3 presents all the results of the EDS analysis of each point in Figs. 11, 13, and 15.

| 13

Each analysis result shows the composition of the adhered material or coating. This implies that the enrichment of specific elements or the existence of a clear reaction layer at the interface was not confirmed. The construction of the dislocation cell and recovery of the strain induced atomic rearrangement just above the coating, recrystallization while did not occur in the adhered materials, when the adhered material was substantially strained and heated by the cutting process. The atomic rearrangement began on the coating surface and produced the crystal orientation relationship at the interface. The subdivision of the grain and recovery of the adhered material occurred at the same time, and the grain size thus remained about 50 nm. The above results of Ti-6Al-4V alloy machining suggest that the adhesion interface between the coating of the cutting tool and the adhered workpiece material is relatively strong owing to the existence of the crystal orientation relationship. The formation of this high-strength interface causes the preferential damage of a ceramic (or any other brittle). This is due to the force exerted by the adhered material on the grain boundary on the damaged surface of the coating. Moreover, there is a difference between the damage rates of a ductile cemented carbide and a brittle coating. However, the reason why there is hardly a crystal orientation relationship between certain adhered materials and coatings is presently unclear.

3.3 Comparison of tool coating damages during machining of Ti-6Al-4V alloy and alloy 718 Figure 16 compares the damage models of a TiN-coated cutting tool for the turning of Ti-6Al-4V alloy and alloy 718. The models depict the observed coating damages as discussed in Section 3.2. From the observation of the surface and cross section of the cutting edge, it is believed that the adhered material directly above the damaged portion of the coating maintained a constant thickness during the cutting. It is therefore assumed that the chip sequentially generated material flow over the adhered material. Based on

| 14

this model, diagrams of the chip, adhered material, and coating were drawn. Fig. 16(a) shows the initial adhered conditions on the coating surface during the machining of Ti-6Al-4V alloy and alloy 718, while Fig. 16(b) shows the corresponding progressions of the coating damage. At the beginning of the adhesion shown in Fig. 16(a), both the Ti-6Al-4V alloy and alloy 718 are exposed to a cutting temperature of about 680 °C, and the adhered material is introduced to enormous strain. This causes grain subdivision and subsequent crystal recovery, with the fine grains of the adhered material measuring about several ten nanometers. In the case of the machining of Ti-6Al-4V alloy, because the structure of the adhered material is recovered by atomic rearrangement under the effect of the coating surface, the crystal orientation relationship at the interface is observed as a feature of the coating surface. It is considered that the crystal orientation relationship at the interface is induced by crystal recovery in the adhesive material when reorganization also occurs in the damaged part of the coating. It is inferred that the interfacial strength in the presence of a crystal orientation relationship is higher than when there is no such relationship. However, it has been reported that there is no crystal orientation relationship at the interface between the adhered material and the TiN coating during the machining of alloy 718 [14, 17]. It is therefore deduced that the difficulty of the adhered material to slide over the coating surface during the cutting of Ti-6Al-4V alloy is partly due to the crystal orientation relationship, while the adhered material is able to slide during the cutting of alloy 718 due to the absence of a crystal orientation relationship at the interface. In addition, the adhered material, which directly interacts with the coating surface during the cutting, is characterized by a fine grain structure. This makes it necessary to take the properties of the load applied by the adhered material into consideration. As discussed in Section 3.2, the adhered material during the machining of Ti-6Al-4V alloy was subjected to enormous strain and was observed to exhibit a high yield stress and high work hardening, but low fracture strain. Conversely, although the adhered material during the machining of alloy 718 initially exhibited significant work hardening, the work hardening in the high strain region was low. It is considered that the adhered material was stable in

| 15

the high strain region and that the deformation stress did not vary with the conditions [18]. Generally, the application of enormous strain to a workpiece degrades the ductility, especially when the uniform elongation is very small. It has also been reported that the yield strength is significantly increased by ultra-miniaturization of the grain, while the work hardening property is not increased, and that this leads to plastic instability (the development of a constriction during a tensile test) at an early stage [21]. In the cutting of Ti-6Al-4V alloy as shown in Fig. 16(b), the movement of the sequentially generated chips produces frictional force on the top of the adhered material. The slide of the adhered material over the coating is made difficult by the crystal orientation relationship, and it is believed that a shear force acts on the adhered material sandwiched between the chip and the coating. The yield strength of the coating is estimated to be higher than that of the adhered material, and the fracture strain of the material is therefore considered to be small due to the miniaturization of the crystal grains. Hence, the adhered material first fractures, and the fracture becomes the starting point of a crack that propagates through the coating. This is probably because of the stronger interfacial bond resulting from the crystal orientation relationship. The adhered material finally falls off when the coating fractures. In the case of the machining of alloy 718, there is no crystal orientation relationship at the interface between the adhered material and the coating surface. In addition, the fracture strain of the adhered material is large and the material exhibits high fracture strength at the elevated temperatures. The adhered material is therefore capable of sliding on the coating surface without fracturing. Meanwhile, sufficient heat flows into the tool, and the sliding of the high-strength adhered material applies a frictional force on the tool coating. Plastic deformation and microfracture of the coating is thus presumed to occur. 3.4 Wear progress of uncoated cutting tool Figure 17 shows the SEM images of the rake faces of the coated and uncoated cutting tools after machining with a more practical cutting speed V of 60 m/min. Another cutting test conducted using a cutting speed of 40 m/min revealed the same morphology. In these

| 16

cutting tests, the coating was not confirmed to have a wear suppressing effect. Figure 18 shows enlarged images of the material adhesion to the cutting edge of the uncoated tool shown in Fig. 17(a). In the composition BEI image (Fig. 18(b)), the adhered material shown in dark gray adheres in the vicinity of the cutting edge. The adhered material was observed over the entire contact area of the chip on the rake face. This suggests that the chip slides over the area containing the adhered material, as in the case of the coated tool shown in Fig. 16. Incidentally, particles considered to be wear particles, were observed on the adhered material. The qualitative results of a WDS analysis of the particles revealed the presence of cemented carbide components. Figure 19 shows the mapping of the area near the wear particles in Fig. 18(a). It was confirmed that W and Co were present at the same point. Although C can be used to obtain information about contamination, its detectability is low and it was not necessarily detected at the same point as W. Moreover, WC-Co and O were not detected at the same point, and it is therefore believed that the observed falling off of the WC-Co particles was not due to oxidation. Furthermore, because the particles measured several microns, it is believed that the agglomerated WC or WC-Co particles in the cemented carbide dropped out. It is considered that the generation of the wear particles is due to the softening of the Co phase by the cutting heat and the consequent reduction of the holding force of the WC particles. It is thus important to enhance the cooling to suppress the production of the wear particles. Although it has been reported that the reaction phases such as TiC and the

-phase (Co2W4C) are produced at the interface between the cemented carbide and the adhered Ti-6Al-4V alloy [11,13], within the level of the analysis accuracy of the present study, the formation of these phases were not confirmed. Further study is necessary to investigate the production of the reaction phases and the abrasion of the reaction layer. It is concluded that Ti alloys are suitably machined using WC-Co cemented carbide tools. Cooling of the cutting tool is also strongly recommended. A brittle material such as a TiN coating with defects that induce stress concentration is unsuitable. 4. Conclusions

| 17

The wear mechanisms of PVD-coated cutting tools during the continuous turning of Ti-6Al-4V alloy were investigated to clarify the damage factors. The wear modes of the tool coating were examined by various analytical techniques, and the property requirements of the coating were established. The main conclusions of the study are summarized as follows: (1) The damage rate of the coating is higher than that of the uncoated cemented carbide tool, with the coating on the rake face disappearing through brittle fracture at an early stage of the machining. After exposure of the cemented carbide, the wear rate reduces. (2) The densely adhered workpiece material on the coating surface has fine grains measuring less than 100 nm, and there is a crystal orientation relationship at some adhesion interfaces. However, the enrichment of specific elements or the formation of a reaction layer at the interfaces was not confirmed. (3) Plastic deformation of the coating at the damaged point was not observed, although there was cracking in parts of the columnar crystal grain boundary. This suggests that the damage of the coating was due to the force exerted by the adhered materials on the grain boundary on the previously damaged coating surface. (4) Comparison of the damage modes of the coating during the machining of Ti-6Al-4V alloy and alloy 718 revealed a difference. It is considered that the damage of the coating is influenced by the fracture strength of the adhered material, as well as the interfacial strength (adhesion strength) with respect to the presence or absence of a crystal orientation relationship between the adhered material and the coating. Based on this consideration, we proposed a model for explaining the coating damage for specific workpiece materials. (5) The wear of the cemented carbide resulted from the falling off of WC-Co particles due to the softening of the Co phase at elevated temperatures by the cutting heat.

| 18

References [1] Ezugwu EO, Wang ZM, Machado AR. Titanium alloys and their machinability - A review. J Mater Process Technol 1997;68;262–274. | 19 [2] Motonishi S, Isoda S, Yoshida K, Itoh H, Tsumori Y, Terada Y. Machining study on titanium alloys. Kobe Steel Eng Rep 1885;35(3)61–64. [3] Obikawa T, Usui E. FEM analysis on serrated chip formation of titanium alloy (2nd Report) – Chip formation mechanism and cutting properties of titanium alloy. J Jpn Soc Precis Eng 1993;59(6):933–938. [4] Köning W. AGARD Conf Proc 1979;256;1.1–1.2. [5] Usui E, Obikawa T, Shirakashi T. Study on machining process of difficult-to-machine materials (2nd Report) – Analysis of stress and temperature distributions by visio-plasticity technique and finite difference method. J Jpn Soc Precis Eng 1986;52(9);1623–1630. [6] Ducobu F, Lorphèvre ER, Filippi E. Numerical contribution to the comprehension of saw-toothed Ti6Al4V chip formation in orthogonal cutting. Int J Mech Sci 2014;81;77–87. [7] Usuki H, Sato K, Furuya S. High speed dry milling of titanium alloys with coated carbide tool. J Jpn Soc Precis Eng 2005;71(4);491–495. [8] Hartung PD, Kramer BM. Tool wear in titanium machining. Annuals of the CIRP 1982;31(1);75–80. [9] Maekawa K, Ohshima I, Suzuki K. Improvements in cutting efficiency of Ti-6Al-2Sn titanium alloy (2nd Report) – Investigations for reducing tool tip temperature. J Jpn Soc Precis Eng 1993;59(6)927–932. [10] Jawaid A, Sharif S, Koksal S. Evaluation of wear mechanisms of coated carbide tools when face milling titanium alloy. J Mater Process Technol 2000;99;266–274. [11] Ikuta A, Shinozaki K, Masuda H, Kuroki H, Yamane Y, Fukaya Y. Bondability between various titanium alloy and cemented carbide during cutting process - Study

on the adhesion mechanism during machining process (Report 1). Quart J JWS 2000;18(2);280–287. [12] Jianxin D, Yousheng L, Wenlong S. Diffusion wear in dry cutting of Ti–6Al–4V with WC/Co carbide tools. Wear 2008;265;1776–1783. [13] Dukuzaki K, Ikuta A, Masuda H, Yamane Y, Kuroki H, Aritoshi M, Fukaya Y. Fundamental study on adhesion mechanism of difficult-to-machine materials during cutting (1st report) - Estimation on properties of adhering titanium alloys to cemented carbide tool. J Jpn Soc Precis Eng 2000;66;224–228. [14] Koseki S, Inoue K, Usuki H. Damage of physical vapor deposition coatings of cutting tools during alloy 718 turning. Prec Eng 2016;44;41–54. [15] Kitagawa T, Shirakashi T, Usui E. Characteristic equation of crater wear - Study on analytical prediction of cutting tool life (1st report). J Jpn Soc Precis Eng 1976;42;1178–1183. [16] Zorev NN. Interrelationship between shear processes occurring along tool face and on shear plane in metal cutting. Int Res Product Eng in New York 1963;9;42–49. [17] Koseki S, Inoue K, Morito S, Ohba T, Usuki H. Comparison of TiN-coated tools using CVD and PVD processes during continuous cutting of Ni-based superalloys. Surf Coat Tech 2015;283;353–363. [18] Obikawa T, Shirakashi T, Usui E. Study on machining process of difficult-to-machine materials (1st report) – Flow stress characteristics and material properties of titanium alloy, stainless steel, and nickel base superalloy. J Jpn Soc Precis Eng 1986;52(1);127–133. [19] Koseki S, Inoue K, Uehara K, Usuki H, Yoshinobu M, Tanaka R, Hagino M. Damage of PVD-coated cutting tools due to interrupted cutting for alloy 718. Key Eng Mater 2015;656–657;191–197. [20] Umemoto M. Nanocrystallization of steels by severe plastic deformation. Mater Trans 2003;44(10);1900–1911.

| 20

[21] Maki T, Furuhara T, Tsuji N, Morito S, Miyamoto G, Shibata A. Thermomechanical processing

of

steel-Past,

present

and

future.

Tetsu-to-Hagane

2014;100(9);1062–1075. [22] Tsuji N. Ultrafine grained steels. Tetsu-to-Hagane 2002;88(7);359–369. [23] Inoue M. J Jpn Soc Tech Plast 1980;21;128. [24] Chao BT, Trigger KJ. Temperature distribution at the tool-chip interface in metal cutting. Trans ASME 1995;77;1107–1121. [25] Hirao M, Sata T. Measurement of local temperature of the tool face in cutting. J Jpn Soc Precis Eng 1974;40;156–161. [26] Fukuhara Y. Heat treatment of titanium alloys. J Jpn Soc Heat Treat 1986;26;163–169. [27] Matsumoto H, Chiba A, Gejima F, Hanada S. Crystallographic orientation and mechanical properties of α' martensite Ti-V alloy. J JSTP 2009;50; 249–255. [28] Nagashima S. Texture of titanium. ISIJ 1986;72; 314–320.

| 21

Figure captions Fig. 1 Observed microstructure of Ti-6Al-4V alloy.

Fig. 2 Cross-sectional images of the tool materials.

Fig. 3 Diagram of a sample observed cross section of a tool cutting edge.

Fig. 4 Images of the rake face of the cutting edge of a TiN-coated tool after the turning of Ti-6Al-4V alloy. V = 40 m/min, f = 0.2 mm/rev, d = 0.5 mm, cutting length = 10 m.)

Fig. 5 Images of the rake face after the cutting of Ti-6Al-4V alloy (areas A and B of Fig. 4(a)). Upper figure: SEM image; lower figure: compositional image.

Fig. 6 Images of the flank face of the TiN-coated tool after the cutting of Ti-6Al-4V alloy. Upper figure: low magnification; lower figure: high magnification. V = 40 m/min, f = 0.2 mm/rev, d = 0.5 mm, cutting length = 10 m.

Fig. 7 WDS mapping of rake and flank faces of the TiN-coated cutting tool after the turning of Ti-6Al-4V. The analyzed areas correspond to those shown in Figs. 4 and 6(b).

Fig. 8 Cross-sectional SEM images of the cutting edge (cutting line shown in Fig. 4) of the TiN-coated cutting tool after the cutting of Ti-6Al-4V alloy. (a) Overall cutting edge, (b) Cross-sectional image of a large area, (c) Cross-sectional image of a narrow area. V = 40 m/min, f = 0.2 mm/rev, d = 0.5 mm, cutting length = 10 m. The chipping shown in (c) occurred during the preparation of the specimen.

Fig. 9 WDS mapping of a cross section of the cutting edge of the TiN-coated cutting tool after the turning of Ti-6Al-4V. The analyzed area corresponds to the areas shown in Fig.

| 22

8(c)-C. Although slightly green color are shown in doted closed area in the result of Ti, this comes from perpendicular to this figure.

Fig. 10 Cross-sectional TEM images of the TiN coating on the rake face of the tool after the cutting of (a) Ti-6Al-4V alloy (V = 40 m/min), and (b) alloy 718 (V = 30 m/min) [14]. f = 0.2 mm/rev, d = 0.5 mm, cutting length = 10 m.

Fig. 11 High-magnification bright-field TEM image of the interface between the adhered material and the TiN coating, corresponding to the area in Fig. 10(a)-D. The selected area diffraction patterns at several positions in this area are shown in Fig. 12.

Fig. 12 Selected-area diffraction (SAD) and nanobeam diffraction (NBD) patterns at the interface between the adhered material and the TiN coating. The diffraction patterns were obtained from the numbered positions in Fig. 11

Fig. 13 High-magnification dark-field STEM image of the interface between the adhered material and the TiN coating. The observed area corresponds to Fig. 11-F.

Fig. 14 Nano-beam diffraction (NBD) patterns and blight field TEM image of the interface between the adhered material particles and the TiN coating. The diffraction patterns were obtained from the numbered positions in Fig. 13.

Fig. 15 High-magnification TEM and STEM images of the interface around the fracture face. The observed area corresponds to Fig. 10 (a)-E.

Fig. 16 Comparison of the damage models of the TiN-coated cutting tool during the turning of Ti-6Al-4V alloy and alloy 718 [14].

| 23

Fig. 17 SEM images of the cutting edge of the uncoated and TiN-coated cutting tool after the cutting of Ti-6Al-4V alloy. V = 60 m/min, f = 0.2 mm/rev, d = 0.5 mm, cutting length = 10 m. | 24 Fig. 18 SEM and compositional images of the cutting edge of the uncoated cutting tool after the cutting of Ti-6Al-4V alloy. Upper image: low magnification; lower image: high magnification. The wear particles are shown in the high-magnification images. V = 60 m/min, f = 0.2 mm/rev, d = 0.5 mm, cutting length = 10 m.

Fig. 19 WDS mapping of the rake face of the uncoated tool edge after the turning of Ti-6Al-4V alloy. The analyzed area corresponds to that shown in Fig. 18(a).

Fig 1

20 µm

Fig 2

TiN coating

cemented carbide (a) TiN coating

1 μm

2 μm (b) cemented carbide

Fig 3

cross-sectional observation area cutting edge

(at a distance of about 100 μm from the cutting edge)

rake face chip contact area

flank wear

end flank face cutting line

Fig 4

cut for cross-sectional observation (shown in Fig. 8) B A

share stress

maximum wear depth adhered material

width of chip contact area 100 μm

100 μm (a) SEM image

highest temperature

(b) compositional image

normal stress

contact area

distance maximum depth cutting edge (c) diagram of stress distribution

Fig 5

cemented carbide

TiN coating

adhered material

TiN coating

adhered material 10 μm

SEM image

adhered material 10 μm SEM image fracture morphology adhered material

10 μm

10 μm

compositional image compositional image (a) high-magnification at area A (b) high-magnification at area B

Fig 6

adhered material

fracture morphology 50 μm

50 μm

adhered material

fracture of coating

10 μm

(a) tip of the cutting edge

10 μm (b) boundary wear potion

Fig 7

SEM

Al

O

Ti

W

N

low high (a) WDS mapping of rake face

SEM

Al

O

Ti

W

N

low high (b) WDS mapping of flank face

200 μm

50 μm

Fig 8

(a)

rake face

(b)

maximum crater depth position

damage flank face

(c)

cutting edge

20 μm

chipping

exposed area of cemented carbide 30 μm adhered material C

crack

TiN coaing

wear of the coating cemented carbide

10 μm

Fig 9

COMP

Al

O

Ti

W

N

low

high

5 μm

Fig 10

direction of chip flow

direction of chip flow

fracture material fracture of adhered material adhered material fracture of TiN coating

adhered material

D TiN coating TiN coating

E

plastic deformation

no plastic deformation

cemented carbide (a) after cutting of Ti-6Al-4V alloy

(b) after cutting of alloy 718

0.5μm

Fig 11

2

adhered material

area b 3 1 F area a TiN coating

200 nm

Fig 12

area a (SAD)

TiN (cubic)

point 1 (NBD)x

a-Ti(hexagonal)

area b (SAD)

a-Ti(hexagonal)

point 2 (NBD)

a-Ti(hexagonal)

point 3 (NBD)

a-Ti(hexagonal)

Fig 13

adhered material

6 5

7

8

4 TiN coating

50 nm

Fig 14

point 4 (NBD)

point 5 (NBD)

point 6 (NBD)

a-Ti hexagonal

{100}a-Ti TiN (cubic)

superposition of point 4 and 6

point 7 (NBD)

a-Ti(hexagonal)

area pointb8(SAD) (NBD)

{010}TiN

a-Ti(hexagonal)

a-Ti(hexagonal)

(a) nano-beam diffraction (NBD) patterns

TiN cubic

2 nm

(b) blight field TEM image near the point 5

Fig 15

adhered material 10 9 crack TiN coating

columnar grain boundary 50 nm

200 nm

chip

chip flow

fine crystallization ⇒ high yield stress and low fracture strain crystal orientation relationship

fine crystallization ⇒ high yield stress and high fracture strain no crystal orientation relationship

Ti-6Al4V cutting

alloy 718 cutting

coating

chip flow

adhered material

Fig 16

coating

adhered material

chip

(a) Initial adhered condition on coating surface friction force

friction force

stable at high temperature thermal softening high fracture strength low fracture strength ⇒fracture of adhered material ⇒sliding with coating debris with coating difficult to slide micro fracturing crack no plastic deformation Ti-6Al-4V cutting

plastic deformation alloy 718 cutting

(b) Progression of coating damage

Fig 17

wear area of TiN coating adhered material

adhered material

100 μm

100 μm

(a) uncoated tool

(b) TiN-coated tool

Fig 18

thin adhesion area

adhered material

thick adhered material

50 μm

50 μm

wear particles shown in Fig. 19

10 μm (a) SEM image

10 μm (b) compositional image

Fig 19

WC-Co particles

SEM

Ti

O

W

Co

C

low

high

10μm

Table captions Table 1 Machining conditions. Workpiece material

Machining tool

Machining conditions

Ti-6Al-4V, annealed, 715 C × 1 h → A.C., 100 mm CNMG120408 Rake back and side angle = 13° Relief back and side angle = 0° Geometry After fixing in the tool holder: Rake back and side angle = 7° Relief back and side angle = 6° Cemented carbide Substrate (JIS K05 grade) Coating composition TiN-coated or Uncoated Coating method PVD Arc Coating thickness 5 μm Cutting speed 40 m/min 60 m/min Cutting length 10 m Feed 0.2 mm/rev Cutting depth 0.5 mm Coolant Wet

| 25

Table 2 Employed analysis methods and conditions. Method Device name Analytical conditions SEM Hitachi, N3700 Acceleration voltage: 15 kV Area analysis EPMA Acceleration voltage = 15 kV JEOL, / WDS Probe current = 0.3 μA JXA-8500F / BSI Beam diameter, rake face = 1 μm Flank face = 0 μm Selected-area diffraction (SAD): Acceleration voltage = 200 kV Diffraction area = 140 nm TEM Nano-beam diffraction (NBD): JEOL, / STEM Acceleration voltage = 200 kV JEM-2010F / EDS Beam diameter = 1–3 nm Energy dispersive X-ray spectrometry (EDS): Acceleration voltage = 200 kV Beam diameter = 1 nm

Table 3 TEM-EDS results for the adhered material and the interface between the adhered material and the TiN coating. The analyzed area corresponds to the areas shown in Figs. 11, 13, and 15. Al V W Fe Ti O Analysis point (at.%) (at.%) (at.%) (at.%) (at.%) (at.%) Point 1 14 3 1 0 81 0 Adhered material Point 2 16 4 0 1 79 0 Point 3 14 4 0 1 98 0 Coating Point 4 1 0 0 0 98 0 Adhered material / coating Point 5 11 4 0 2 84 0 Point 6 13 7 1 3 77 0 Adhered material Point 7 16 3 0 0 80 0 Point 8 12 11 1 5 72 0 Coating Point 9 0 0 0 0 99 0 Adhered material Point 10 13 4 1 0 82 0