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Fatigue failure of coated carbide tool and its influence on cutting performance in face milling SKD11 hardened steel
Int. Journal of Refractory Metals and Hard Materials 64 (2017) 27–34
Contents lists available at ScienceDirect
Int. Journal of Refractory Metals and Hard Materials journal homepage: www.elsevier.com/locate/IJRMHM
Fatigue failure of coated carbide tool and its influence on cutting performance in face milling SKD11 hardened steel Feng Gong, Jun Zhao ⁎, Yiwei Jiang, Haiwang Tao, Zuoli Li, Jian Zang Key Laboratory of High Efficiency and Clean Mechanical Manufacture of MOE, School of Mechanical Engineering, Shandong University, Jinan 250061, China
a r t i c l e
i n f o
Article history: Received 9 November 2016 Received in revised form 3 January 2017 Accepted 8 January 2017 Available online 10 January 2017 Keywords: Coated tool Failure patterns Fatigue morphologies Cutting forces Surface roughness
a b s t r a c t Failure patterns of coated carbide tool were investigated by high-speed face milling of the hardened steel SKD11. Tool failure surface morphology, cutting force and machined surface roughness were also analyzed to reveal the failure mechanisms. The results indicated that the dominant failure pattern of coated carbide tool was breakage. The primary mechanism of tool breakage was fatigue fracture. Under different cutting speeds, the distinctive morphologies of fatigue fracture were presented on the failure surfaces. At low cutting speeds, many fatigue sources were observed on the rake face. The distance between fatigue sources and tool nose was approximately two times of the depth of cut. With the increase of cutting speed, the fatigue striations and riven patterns were observed at the fracture surface. In addition, the fatigue steps and crack deflection were found under high cutting speeds. The main fracture mode was intergranular fracture at lower cutting speeds. However, it was transgranular fracture at higher cutting speeds. Furthermore, the irregular fracture surfaces at low cutting speeds and at high cutting speeds contribute to a larger cutting force increment compared with the medium cutting speeds. The increment of surface roughness in the initial and severe wear stages was lower than that in the steady wear stage, while the deviation of surface roughness was relatively large. © 2017 Elsevier Ltd. All rights reserved.
1. Introduction Hardened steels have been widely used in the mould manufacturing industry because of their high hardness, high strength and good wear resistance. The traditional solution to finishing hardened steel parts is grinding. However, the grinding has many drawbacks, including low production rate, surface burning and high cost. Thus, the hard machining, which has lots of benefits, such as high metal removal rate, elimination of coolants in most cases and low machine tool investment, comes into being [1]. However, hard machining will contribute to high cutting forces and cutting temperature, which puts out a higher requirement of the performance for cutting tools [2]. The coated carbide tools have outstanding physical and mechanical properties, such as high surface hardness, low friction factor, low thermal conductivity and stable chemical properties. In addition, compared with the ceramic and CBN tools, the coated carbide tools have better economic applicability, and have been widely used in the hard machining. Many scholars have conducted considerable studies to investigate the cutting performance of coated carbide tool in hard machining. Urbanski et al. [3] reported the tool wear by using indexable inserts of ball end mill to machine hardened AISI H13 steel with different tool material, and found that TiCN coated tools had the highest tool life. Ghani et ⁎ Corresponding author. E-mail address: [email protected] (J. Zhao).
http://dx.doi.org/10.1016/j.ijrmhm.2017.01.001 0263-4368/© 2017 Elsevier Ltd. All rights reserved.
al. [4] also investigated the wear mechanisms of TiN coated tool under different cutting parameters for end milling of hardened AISI H13 steel. They observed that flank wear was the dominating tool failure mechanism at vc = 224–355 m/min. Cui and Zhao [5] also studied the wear mechanisms of coated cemented carbide tool in face milling of H13 hardened steel. It was found that the abrasion wear was the tool main failure mechanism in symmetric milling. Wang et al. [6] used two kinds of coated tools for milling of hardened steel to explore the tool failure mechanisms. The experimental results indicated that rack face wear, flank wear, breakage and micro-chipping were the tool dominant failure patterns. It can be concluded that the abrasion and adhesion wear are the primary failure mechanisms of coated carbide tool in hard machining. On the other hand, the crack growth and fatigue behavior on the tool surface are also investigated by other scholars. Yin et al. [7,8] used threepoint bending test to analyze the fatigue behavior of Al2O3/TiC micro– nano-composite ceramic tool materials. The results showed that the slow crack growth zone of specimens all originated from the tensile surface. In addition, they also confirmed that the maximum tension stress was at the tool nose through the finite element simulation. Ramírez et al. [9] compared the fatigue behavior of TiN coating and WC: H thin film coating. They found that the WC: H thin film coating tool has a better performance compared with the TiN-coated tool. Gui et al. [10] also studied the slow crack growth behavior of Si3N4 ceramics in ambient and the cryogenic environment by dynamic fatigue testing. Zheng et
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F. Gong et al. / Int. Journal of Refractory Metals and Hard Materials 64 (2017) 27–34
al. [11] analyzed the crack extension paths by the fractal geometry method. The results showed that the crack paths in the Si3N4 composite had a very obvious fractal characteristic. However, the fatigue fracture of the coated tool in the high-speed hard milling was rarely reported. And the fatigue characteristic is not studied adequately in the practical machining. Although the fractography was widely applied to the observation of the fatigue crack propagation behavior in the metallic materials [12–14], it is rarely used in the analysis of the tool fatigue behavior. In this paper, the hardened steel SKD11was selected as the workpiece to study the tool fatigue behavior under different cutting speeds. The fractography was used for observing the fracture surface of the tool to analyze the fatigue crack propagation process. The location of fatigue source was investigated by the three dimensional (3D) laser microscope. In addition, the influence of tool damage on the cutting forces and machined surface roughness were also introduced. 2. Experimental procedure 2.1. Cutting tool and workpiece
Fig. 2. Microstructure of the coated cement carbide tool.
All the tests were conducted on a vertical CNC milling machine DAEWOOACE-V500 (South Korea) in dry machining. The spindle rotational speed of the machine tool ranges from 80 to 10,000 rpm. The photograph of the experimental setup is shown in Fig. 1. The coated cemented carbide inserts (Model SEEX 09T3AFTN-D09, type MP1500) produced by Seco Tool were used for the milling test. Fig. 2 shows the microstructures of the coated cement tool. The tool is CVD coated with double layer of Al2O3/Ti(C, N). The face milling cutter R220.53-012509-8C (Seco, Sweden) with the diameter of 125 mm was used in the test, which can hold eight inserts. To avoid the influence caused by the difference among the inserts, only one of tooth was used in all the tests. The major cutting edge angle, radial rake angle and axial rake angle of the tool were 45°, −5°and 20°, respectively. The hardened steel SKD11, which was widely applied in the production of die, precision gauges, spindle, jigs and fixtures due to its high toughness and great wear resistance, was selected as workpiece material. In this experiment, the rectangular block of SKD11 with a dimension of 105 mm × 75 mm × 50 mm was applied. The chemical composition and hardness, and physical properties of SKD11 are given in Table 1 and Table 2.
SKD11. The feed per tooth fz, axial depth of cut ap, and radial depth of cut ae were fixed at 0.05 mm/z, 0.10 mm and 75 mm as shown in Table 3. The symmetric milling was adopted in the present test. In this test, the insert condition was examined periodically by the tool microscope (AM413ZT, Taiwan) after each pass. The Kistler threecomponent piezoelectric dynamometer (type 9257A) was used to measure cutting forces. The machined surface roughness was measured by the TR200 portable surface roughness tester (China, scanning length was 0.8 mm). Each machined surface was measured for three times to calculate the average machined surface roughness value. In order to reveal the tool failure patterns and mechanisms, the tool failure surface morphologies were analyzed by a 3D laser microscope (KEYENCE VKX200K, Japan) and energy-dispersive spectroscopy (SEM and EDS, JSM-6510 LV). When the flank wear value reaches 0.3 mm or severe breakage occurs, according to the ISO standards, the tool was considered as a failure. 3. Results and discussions 3.1. Tool failure mechanisms
2.2. Cutting test A single factor experiment design was employed to explore the influence of cutting speeds (vc) ranging from 90 m/min to 180 m/min on the failure patterns of the tool when face milling of hardened steel
Fig. 3 shows the SEM micrographs and EDS analysis of worn tool at vc = 90 m/min, from which it can be seen that the tool damage region is located at the rake face. The tool failure mechanisms include fatigue fracture, chipping and wear, as can be seen in Fig. 3b. The multi-fatigue
Fig. 1. Experimental setup and the tool geometry.
F. Gong et al. / Int. Journal of Refractory Metals and Hard Materials 64 (2017) 27–34
29
Table 1 Chemical composition and hardness of SKD11 [15]. Material
Element (%)
SKD11
Hardness (HRC)
C
Si
Mn
Cr
Mo
V
P
1.4– 1.6
≤0.6
≤0.6
11.0– 13.0
0.7– 1.2
0.15–0.3
b0.03
crack sources generate at the rake face because of the high cutting force and repeated impact load. The riven patterns are observed at the fatigue propagation zone, as shown in Fig. 3a. It can be seen that the crack propagation mode is the I–III mixed mode cracks. The coating peeling occurs at rake face. In addition, the ratcheting mark [17] is observed in Fig. 3c. It is attributed to the fact that multiple cracks nucleate at different points and join together on the fracture surface. In Fig. 3f, the fatigue cracks adjacent to the ratcheting marks are found. The EDS analysis of region A in Fig. 3e shows that Fe element (78.70 at.%) and a small amount of O element (7.41 at.%) are detected in the adhesion zone, which indicates that oxidation and abrasion wear occur during the low speed milling of hardened steel. Fig. 4 exhibits the SEM micrographs and EDS analysis of worn tool at vc = 150 m/min, from which it can be seen that the main failure mechanism is fatigue fracture. A fan-shaped fatigue propagation region is discovered at the rake face. The fatigue characteristics at vc = 150 m/min are remarkably different from that at vc = 90 m/min. The fatigue striations, which are the typical fatigue characteristics, are observed at the tool nose as shown in Fig. 4a. During the milling process, the fatigue crack growth rate is unstable because of the tool damage. This induces the irregularity of fatigue striation width. The propagation direction of the fatigue crack, which is indicated by the red arrow mark (Fig. 4a), is perpendicular to the fatigue striations. Besides that, the river patterns are also observed on the fracture surface away from tool nose (Fig. 4c). The propagation direction of fatigue crack is consistent with river patterns. It can be inferred that the fatigue source is located on the area between the fatigue striations and river patterns. In addition, as can be seen from Fig. 4f, the main fracture mode is the intergranular fracture at vc = 150 m/min. The adhesion and oxidation wear also occur at the flank face, as shown in Fig. 4d and e. The bulk of sawtooth chips are bonded on the tool nose due to the high cutting forces and cutting temperature. The SEM micrographs and EDS analysis of worn tool at vc = 180 m/min are presented in Fig. 5. At the higher cutting speed, the appearance of fracture surface (Fig. 5b) is irregular compared with the cases at other cutting speeds (Figs. 3b and 4b). The small cutting force and rapid failure of tool make that only one fatigue source is observed at the rake face (Fig. 5a). Because the internal stress far away from the tool nose falls off rapidly, the crack deflection labeled by red arrows is found in Fig. 5c. Moreover, due to the asymmetrical stress, the fatigue steps with the secondary crack are also observed on the fatigue crack propagation zone. The fatigue steps can increase the roughness of fracture surface, which has a negative influence on the cutting forces and machined surface roughness. The transgranular fracture is the main fracture mode, as can be seen from Fig. 5f. It can be revealed that the grain interior is weaker than grain boundaries at high temperature. Because of the high cutting temperature, the adhesion wear and severer oxidation wear occur at the tool nose (Fig. 5d). Fig. 6 shows the two-dimensional micrographs of the rake face at vc = 90 m/min, vc = 150 m/min and vc = 180 m/min. The two-dimensional profiles, which cross the fatigue source (point 2), are plotted, as Table 2 Physical properties of SKD11 [16]. Density (kg/m3)
Young's modulus (GPa)
Poisson's ratio
Thermal conductivity (W/m·K)
Specific heat (J/kg °C)
8400
208
0.3
20.5
461
62
shown in Fig. 6b, d and f, respectively. From Fig. 6b, the chipping is found at the tool nose due to the large cutting forces. The fatigue source is located in the valley of the profile. It can be concluded that the fatigue cracks originate from the defect of tool substrate (point A in Fig. 2). The distance between the tool nose and fatigue source (point 1 to point 2) is about 200 μm, which is two times of the depth of cut. In addition, the fatigue crack path is also investigated (point 2 to point 3). Because the tool is subjected to severe mechanical stress at low speed, the fatigue crack path is relatively rough. However, as can be seen from Fig. 6d and f, the fatigue crack path is relatively smooth at vc = 150 m/min and vc = 180 m/min. It is attributed to the fact that the high cutting temperature decreases the resistance to crack propagation. 3.2. Cutting forces In order to investigate the influence of tool damage on the cutting forces, the cutting forces and flank wear value were measured when the tool goes through each pass of the workpiece surface. Fig. 7 presents the cutting force components and resultant cutting force for face milling SKD11 hardened steel at vc = 90 m/min, from which it can be seen that resultant cutting force reaches the peak when the tool cut into the workpiece. So the cutting forces are calculated based on the values extracted from the tool cut-into workpiece period. The average cutting force components are calculated by Eq. (1): Fx ¼
1 N 1 N 1 N ∑i¼1 Fxi ; Fy ¼ ∑i¼1 Fyi ; Fz ¼ ∑i¼1 Fzi N N N
ð1Þ
And the resultant cutting force Fr is calculated by the Eq. (2): qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Fri ¼ F2xi þF2yi þF2zi
ð2Þ
where Fxi, Fyi and Fzi are cutting force components in x, y, and z directions, respectively. Fig. 8 shows the flank wear and resultant cutting force versus volume of removed metal in milling SKD11. The flank wear rate at vc = 90 m/min is faster than that under the speed of 120 m/min in the initial tool wear stage (Fig. 8a), because of the larger cutting force at vc = 90 m/min as shown in Fig. 8b. It is known that the cutting temperature becomes higher when the cutting speed increases [18]. Therefore, it can be concluded that the cutting force has a larger influence on the tool wear than cutting temperature in the initial tool wear stage. After that, the coating peeling occurs at the rake face, which results in increase in both cutting force and cutting temperature. The cutting force at vc = 90 m/min is still larger than that at vc = 120 m/min. However, tool wear rate at vc = 90 m/min is slower than that under 120 m/min. It is inferred that the cutting temperature has a larger influence on the tool wear after the coating peeling. Fig. 9 depicts the cutting force components and resultant cutting force when face milling SKD11 in the initial wear stage. Obviously, it can be seen that the axial cutting force Fz are much larger than other cutting force components due to high hardness of workpiece. Furthermore, Table 3 Cutting parameters of face milling experiments. Cutting parameters
vc (m/min)
fz (mm/z)
ap (mm)
ae (mm)
90, 120, 150, 180
0.05
0.10
75
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Fig. 3. SEM micrographs and EDS analysis of worn tool (vc = 90 m/min, f = 0.05 mm/z, ap = 0.1 mm, ae = 75 mm).
the resultant cutting force firstly decreases and then increases with the increase of cutting speed from 90 m/min to 150 m/min. At last, it decreases when the cutting speed is above 150 m/min. The decrease in cutting force with the increasing of cutting speed up to 120 m/min can be explained by the decreasing of the friction coefficient between the tool and workpiece and the increasing shear angle with increasing cutting speed. However, when the speed exceeds 150 m/min, the thermal softening effect has a huge influence on the cutting forces, which also results in cutting force reduction.
To explore the influence of the tool damage on the cutting force, the growth rate of cutting force is defined as shown in Eq. (3) in the present study: δ¼
ð Ffin−FiniÞ 100% Fini
ð3Þ
where Fini and Ffin are the cutting forces when the tool goes through the first and last pass of workpiece.
Fig. 4. SEM micrographs and EDS analysis of worn tool (vc = 150 m/min, f = 0.05 mm/z, ap = 0.1 mm, ae = 75 mm).
F. Gong et al. / Int. Journal of Refractory Metals and Hard Materials 64 (2017) 27–34
Fig. 5. SEM micrographs and EDS analysis of worn tool (vc = 180 m/min, f = 0.05 mm/z, ap = 0.1 mm, ae = 75 mm).
Fig. 6. Two-dimensional micrographs of the rake face at (a) vc = 90 m/min (c) vc = 150 m/min (e) vc = 180 m/min.
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Fig. 7. Cutting force components and resultant cutting force in face milling of SKD11 at 90 m/min.
Fig. 10 shows the growth rate of the cutting force components and resultant cutting force. As shown in Fig. 10, with the increase of cutting speed, the cutting force firstly decreases until a minimum value reached at the cutting speed of 120 m/min, then increases when the cutting speed is beyond that. The irregular fracture surface at vc = 90 m/min (Fig. 3b) and 180 m/min (Fig. 5b) makes the ploughing rather than cutting occurs between the workpiece and tool [19], which leads to high growth rate of cutting force. 3.3. Machined surface roughness Fig. 11 shows the tool flank wear, resultant cutting force and surface roughness in milling of SKD11 at vc = 120 m/min. It can be seen that
variation trends of flank wear, cutting force and machined roughness value are identical. In the initial wear stage, the flank wear and cutting force are both stable, however, the roughness value has a jump increase. Although the value of flank wear is relatively low, the condition of rake face constantly deteriorates because of repeated impact load. The friction between the tool and workpiece is unstable, which results in the jumping change of surface roughness. In the steady wear stage, the flank wear and cutting forces increase rapidly, leading to the rapid increase in surface roughness. At last, when the flank wear reaches 0.31 mm, the machined surface roughness Ra rises from 0.15 μm to 0.86 μm. In addition, the deviation of roughness values is relatively large in the initial and severe wear stage. In the first pass, the tool flank wear
Fig. 8. (a) flank wear and (b) resultant cutting force versus volume of removed metal in milling of hardened steel SKD11.
F. Gong et al. / Int. Journal of Refractory Metals and Hard Materials 64 (2017) 27–34
Fig. 9. Cutting force components and resultant cutting force when face milling of SKD11 in the initial wear stage.
changes in a wide range (rising from 0 to 0.05 mm) which results in a large deviation of roughness value. In the steady wear stage, the surface of cutting tool matches the cutting tracks of last pass, leading to the low deviation of roughness value. In the severe wear stage, the cutting tool has a serious breakage and is extremely prone to break. Therefore, the deviation of roughness value apparently increases. 4. Conclusion The failure patterns and failure mechanisms of coated carbide tool were investigated by high-speed face milling of the hardened steel SKD11. In addition, the influence of tool damage on cutting forces and surface roughness were also analyzed. The following conclusions can be drawn from these investigations: (1) The dominating failure pattern of coated tool when face milling of the hardened steel SKD11 was fatigue fracture accompanied by chipping. At vc = 90 m/min, the fracture caused by multi-fatigue crack propagation occurred on the rake face. With the increase of cutting speed, the fatigue striations and river patterns were observed at the fatigue crack propagation zone. At vc = 180 m/min, the fatigue steps with the secondary crack were also observed on the fatigue propagation area. The main fracture mode was intergranular fracture at lower cutting speeds, while it was transgranular fracture at higher cutting speeds. The adhesion and oxidation wear were presented at the tool nose at vc = 90 m/min–180 m/min.
Fig. 10. Growth rate of the cutting force components and resultant cutting force.
33
Fig. 11. Tool flank wear, resultant cutting force and surface roughness in milling of SKD11 at vc = 120 m/min.
(2) The fatigue crack propagation paths were investigated for different cutting speeds. The fatigue cracks originated from the defect of tool substrate. The distance between the tool nose and fatigue source was approximately two times of the depth of cut. At low cutting speeds, the fatigue path was rough duo to severe mechanical stress. At high cutting speeds, because high cutting temperature decreased the resistance to crack propagation, the fatigue crack path was relatively smooth. (3) In the initial wear stage, the cutting force had a larger influence on tool life. However, when the tool breakage and coating peeling occurred, the cutting temperature was the dominating factor for tool life. In addition, the irregular fracture surfaces at vc = 90 m/min and 180 m/min contributed to a larger cutting force increment. (4) The machined surface roughness was significantly affected by the tool damage. With the flank wear and cutting forces increased, the value of machined surface roughness constantly increased. However, the deviation of roughness was relatively larger in the initial and severe wear stages, and it became smaller in the steady wear stage. Acknowledgements The authors would like to acknowledge the financial support from the National Natural Science Foundation of China (51475273). References [1] J.P. Davim, Machining of Hard Materials, Springer, London, 2011. [2] B. Wang, Z.Q. Liu, Cutting performance of solid ceramic end milling tools in machining hardened AISI H13 steel, Int. J. Refract. Met. Hard Mater. 55 (2016) 24–32. [3] J.P. Urbanski, P. Koshy, R.C. Dewes, D.K. Aspinwall, High speed machining of moulds and dies for net shape manufacture, Mater. Des. 21 (2000) 395–402. [4] J.A. Ghani, I.A. Choudhury, H.H. Masjuki, Wear mechanism of TiN coated carbide and uncoated cermets tools at high cutting speed applications, J. Mater. Process. Technol. 153-154 (2004) 1067–1073. [5] X.B. Cui, J. Zhao, Cutting performance of coated carbide tools in high-speed face milling of AISI H13 hardened steel, Int. J. Adv. Manuf. Technol. 71 (2014) 1811–1824. [6] C.Y. Wang, Y.X. Xie, Z. Qin, H.S. Lin, Y.H. Yuan, Q.M. Wang, Wear and breakage of TiAlN- and TiSiN-coated carbide tools during high-speed milling of hardened steel, Wear 336–337 (2015) 29–42. [7] Z.B. Yin, C.Z. Huang, B. Zou, H.L. Liu, H.T. Zhu, J. Wang, Dynamic fatigue behavior of Al2O3/TiC micro–nano-composite ceramic tool materials at ambient and high temperatures, Mater. Sci. Eng. A 593 (2014) 64–69. [8] Z.B. Yin, C.Z. Huang, J.T. Yuan, B. Zou, H.L. Liu, H.T. Zhu, Cutting performance and life prediction of an Al2O3/TiC micro–nano-composite ceramic tool when machining austenitic stainless steel, Ceram. Int. 41 (2015) 7059–7065. [9] G. Ramírez, E. Jiménez-Piqué, A. Mestra, M. Vilaseca, D. Casellas, L. Llanes, A comparative study of the contact fatigue behavior and associated damage micromechanisms of TiN-and WC:H-coated cold-work tool steel, Tribol. Int. 88 (2015) 263–270. [10] J.Y. Gui, S. Wei, Z.P. Xie, Slow crack growth behavior of silicon nitride ceramics in cryogenic environment, Ceram. Int. 42 (2016) 3687–3691.
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[11] G.M. Zheng, J. Zhao, L. Li, X. Cheng, M. Wang, A fractal analysis of the crack extension paths in a Si3N4 ceramic tool composite, Int. J. Refract. Met. Hard Mater. 51 (2015) 160. [12] J. Schijve, The application of small overloads for fractography of small fatigue cracks initiated under constant-amplitude loading, Int. J. Fatigue 70 (2015) 63–72. [13] X.H. Zhao, Y.J. Fan, Y. Liu, H.Y. Wang, P. Dong, Evaluation of fatigue fracture mechanism in a flash butt welding joint of a U75V type steel for railroad applications, Eng. Fail. Anal. 55 (2015) 26–38. [14] E. Merson, A.V. Kudrya, V.A. Trachenko, D. Merson, V. Danilov, A. Vinogradov, Quantitative characterization of cleavage and hydrogen-assisted quasi-cleavage fracture surfaces with the use of confocal laser scanning microscopy, Mater. Sci. Eng. A 665 (2016) 35–46.
[15] C.Y. Wang, Y.X. Xie, L.J. Zheng, Z. Qin, D.W. Tang, Y.X. Song, Research on the Chip Formation Mechanism during the high-speed milling of hardened steel, Int. J. Mach. Tools Manuf. 79 (2014) 31–48. [16] C.Y. Wang, F. Ding, D.W. Tang, L.J. Zheng, S.Y. Li, Y.X. Xie, Modeling and simulation of the high-speed milling of hardened steel SKD11 (62 HRC) based on SHPB technology, Int. J. Mach. Tools Manuf. 108 (2016) 13–26. [17] G.E. Totten, Fatigue crack propagation, Adv. Mater. Process. 166 (2008) 39–41. [18] M. Kikuchi, The use of cutting temperature to evaluate the machinability of titanium alloys, Acta Biomater. 5 (2009) 770–775. [19] G.L. Liu, B. Zou, C.Z. Huang, X.Y. Wang, J. Wang, Z.Q. Liu, Tool damage and its effect on the machined surface roughness in high-speed face milling the 17-4PH stainless steel, Int. J. Adv. Manuf. Technol. 83 (2015) 257–264.
Contents lists available at ScienceDirect
Int. Journal of Refractory Metals and Hard Materials journal homepage: www.elsevier.com/locate/IJRMHM
Fatigue failure of coated carbide tool and its influence on cutting performance in face milling SKD11 hardened steel Feng Gong, Jun Zhao ⁎, Yiwei Jiang, Haiwang Tao, Zuoli Li, Jian Zang Key Laboratory of High Efficiency and Clean Mechanical Manufacture of MOE, School of Mechanical Engineering, Shandong University, Jinan 250061, China
a r t i c l e
i n f o
Article history: Received 9 November 2016 Received in revised form 3 January 2017 Accepted 8 January 2017 Available online 10 January 2017 Keywords: Coated tool Failure patterns Fatigue morphologies Cutting forces Surface roughness
a b s t r a c t Failure patterns of coated carbide tool were investigated by high-speed face milling of the hardened steel SKD11. Tool failure surface morphology, cutting force and machined surface roughness were also analyzed to reveal the failure mechanisms. The results indicated that the dominant failure pattern of coated carbide tool was breakage. The primary mechanism of tool breakage was fatigue fracture. Under different cutting speeds, the distinctive morphologies of fatigue fracture were presented on the failure surfaces. At low cutting speeds, many fatigue sources were observed on the rake face. The distance between fatigue sources and tool nose was approximately two times of the depth of cut. With the increase of cutting speed, the fatigue striations and riven patterns were observed at the fracture surface. In addition, the fatigue steps and crack deflection were found under high cutting speeds. The main fracture mode was intergranular fracture at lower cutting speeds. However, it was transgranular fracture at higher cutting speeds. Furthermore, the irregular fracture surfaces at low cutting speeds and at high cutting speeds contribute to a larger cutting force increment compared with the medium cutting speeds. The increment of surface roughness in the initial and severe wear stages was lower than that in the steady wear stage, while the deviation of surface roughness was relatively large. © 2017 Elsevier Ltd. All rights reserved.
1. Introduction Hardened steels have been widely used in the mould manufacturing industry because of their high hardness, high strength and good wear resistance. The traditional solution to finishing hardened steel parts is grinding. However, the grinding has many drawbacks, including low production rate, surface burning and high cost. Thus, the hard machining, which has lots of benefits, such as high metal removal rate, elimination of coolants in most cases and low machine tool investment, comes into being [1]. However, hard machining will contribute to high cutting forces and cutting temperature, which puts out a higher requirement of the performance for cutting tools [2]. The coated carbide tools have outstanding physical and mechanical properties, such as high surface hardness, low friction factor, low thermal conductivity and stable chemical properties. In addition, compared with the ceramic and CBN tools, the coated carbide tools have better economic applicability, and have been widely used in the hard machining. Many scholars have conducted considerable studies to investigate the cutting performance of coated carbide tool in hard machining. Urbanski et al. [3] reported the tool wear by using indexable inserts of ball end mill to machine hardened AISI H13 steel with different tool material, and found that TiCN coated tools had the highest tool life. Ghani et ⁎ Corresponding author. E-mail address: [email protected] (J. Zhao).
http://dx.doi.org/10.1016/j.ijrmhm.2017.01.001 0263-4368/© 2017 Elsevier Ltd. All rights reserved.
al. [4] also investigated the wear mechanisms of TiN coated tool under different cutting parameters for end milling of hardened AISI H13 steel. They observed that flank wear was the dominating tool failure mechanism at vc = 224–355 m/min. Cui and Zhao [5] also studied the wear mechanisms of coated cemented carbide tool in face milling of H13 hardened steel. It was found that the abrasion wear was the tool main failure mechanism in symmetric milling. Wang et al. [6] used two kinds of coated tools for milling of hardened steel to explore the tool failure mechanisms. The experimental results indicated that rack face wear, flank wear, breakage and micro-chipping were the tool dominant failure patterns. It can be concluded that the abrasion and adhesion wear are the primary failure mechanisms of coated carbide tool in hard machining. On the other hand, the crack growth and fatigue behavior on the tool surface are also investigated by other scholars. Yin et al. [7,8] used threepoint bending test to analyze the fatigue behavior of Al2O3/TiC micro– nano-composite ceramic tool materials. The results showed that the slow crack growth zone of specimens all originated from the tensile surface. In addition, they also confirmed that the maximum tension stress was at the tool nose through the finite element simulation. Ramírez et al. [9] compared the fatigue behavior of TiN coating and WC: H thin film coating. They found that the WC: H thin film coating tool has a better performance compared with the TiN-coated tool. Gui et al. [10] also studied the slow crack growth behavior of Si3N4 ceramics in ambient and the cryogenic environment by dynamic fatigue testing. Zheng et
28
F. Gong et al. / Int. Journal of Refractory Metals and Hard Materials 64 (2017) 27–34
al. [11] analyzed the crack extension paths by the fractal geometry method. The results showed that the crack paths in the Si3N4 composite had a very obvious fractal characteristic. However, the fatigue fracture of the coated tool in the high-speed hard milling was rarely reported. And the fatigue characteristic is not studied adequately in the practical machining. Although the fractography was widely applied to the observation of the fatigue crack propagation behavior in the metallic materials [12–14], it is rarely used in the analysis of the tool fatigue behavior. In this paper, the hardened steel SKD11was selected as the workpiece to study the tool fatigue behavior under different cutting speeds. The fractography was used for observing the fracture surface of the tool to analyze the fatigue crack propagation process. The location of fatigue source was investigated by the three dimensional (3D) laser microscope. In addition, the influence of tool damage on the cutting forces and machined surface roughness were also introduced. 2. Experimental procedure 2.1. Cutting tool and workpiece
Fig. 2. Microstructure of the coated cement carbide tool.
All the tests were conducted on a vertical CNC milling machine DAEWOOACE-V500 (South Korea) in dry machining. The spindle rotational speed of the machine tool ranges from 80 to 10,000 rpm. The photograph of the experimental setup is shown in Fig. 1. The coated cemented carbide inserts (Model SEEX 09T3AFTN-D09, type MP1500) produced by Seco Tool were used for the milling test. Fig. 2 shows the microstructures of the coated cement tool. The tool is CVD coated with double layer of Al2O3/Ti(C, N). The face milling cutter R220.53-012509-8C (Seco, Sweden) with the diameter of 125 mm was used in the test, which can hold eight inserts. To avoid the influence caused by the difference among the inserts, only one of tooth was used in all the tests. The major cutting edge angle, radial rake angle and axial rake angle of the tool were 45°, −5°and 20°, respectively. The hardened steel SKD11, which was widely applied in the production of die, precision gauges, spindle, jigs and fixtures due to its high toughness and great wear resistance, was selected as workpiece material. In this experiment, the rectangular block of SKD11 with a dimension of 105 mm × 75 mm × 50 mm was applied. The chemical composition and hardness, and physical properties of SKD11 are given in Table 1 and Table 2.
SKD11. The feed per tooth fz, axial depth of cut ap, and radial depth of cut ae were fixed at 0.05 mm/z, 0.10 mm and 75 mm as shown in Table 3. The symmetric milling was adopted in the present test. In this test, the insert condition was examined periodically by the tool microscope (AM413ZT, Taiwan) after each pass. The Kistler threecomponent piezoelectric dynamometer (type 9257A) was used to measure cutting forces. The machined surface roughness was measured by the TR200 portable surface roughness tester (China, scanning length was 0.8 mm). Each machined surface was measured for three times to calculate the average machined surface roughness value. In order to reveal the tool failure patterns and mechanisms, the tool failure surface morphologies were analyzed by a 3D laser microscope (KEYENCE VKX200K, Japan) and energy-dispersive spectroscopy (SEM and EDS, JSM-6510 LV). When the flank wear value reaches 0.3 mm or severe breakage occurs, according to the ISO standards, the tool was considered as a failure. 3. Results and discussions 3.1. Tool failure mechanisms
2.2. Cutting test A single factor experiment design was employed to explore the influence of cutting speeds (vc) ranging from 90 m/min to 180 m/min on the failure patterns of the tool when face milling of hardened steel
Fig. 3 shows the SEM micrographs and EDS analysis of worn tool at vc = 90 m/min, from which it can be seen that the tool damage region is located at the rake face. The tool failure mechanisms include fatigue fracture, chipping and wear, as can be seen in Fig. 3b. The multi-fatigue
Fig. 1. Experimental setup and the tool geometry.
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Table 1 Chemical composition and hardness of SKD11 [15]. Material
Element (%)
SKD11
Hardness (HRC)
C
Si
Mn
Cr
Mo
V
P
1.4– 1.6
≤0.6
≤0.6
11.0– 13.0
0.7– 1.2
0.15–0.3
b0.03
crack sources generate at the rake face because of the high cutting force and repeated impact load. The riven patterns are observed at the fatigue propagation zone, as shown in Fig. 3a. It can be seen that the crack propagation mode is the I–III mixed mode cracks. The coating peeling occurs at rake face. In addition, the ratcheting mark [17] is observed in Fig. 3c. It is attributed to the fact that multiple cracks nucleate at different points and join together on the fracture surface. In Fig. 3f, the fatigue cracks adjacent to the ratcheting marks are found. The EDS analysis of region A in Fig. 3e shows that Fe element (78.70 at.%) and a small amount of O element (7.41 at.%) are detected in the adhesion zone, which indicates that oxidation and abrasion wear occur during the low speed milling of hardened steel. Fig. 4 exhibits the SEM micrographs and EDS analysis of worn tool at vc = 150 m/min, from which it can be seen that the main failure mechanism is fatigue fracture. A fan-shaped fatigue propagation region is discovered at the rake face. The fatigue characteristics at vc = 150 m/min are remarkably different from that at vc = 90 m/min. The fatigue striations, which are the typical fatigue characteristics, are observed at the tool nose as shown in Fig. 4a. During the milling process, the fatigue crack growth rate is unstable because of the tool damage. This induces the irregularity of fatigue striation width. The propagation direction of the fatigue crack, which is indicated by the red arrow mark (Fig. 4a), is perpendicular to the fatigue striations. Besides that, the river patterns are also observed on the fracture surface away from tool nose (Fig. 4c). The propagation direction of fatigue crack is consistent with river patterns. It can be inferred that the fatigue source is located on the area between the fatigue striations and river patterns. In addition, as can be seen from Fig. 4f, the main fracture mode is the intergranular fracture at vc = 150 m/min. The adhesion and oxidation wear also occur at the flank face, as shown in Fig. 4d and e. The bulk of sawtooth chips are bonded on the tool nose due to the high cutting forces and cutting temperature. The SEM micrographs and EDS analysis of worn tool at vc = 180 m/min are presented in Fig. 5. At the higher cutting speed, the appearance of fracture surface (Fig. 5b) is irregular compared with the cases at other cutting speeds (Figs. 3b and 4b). The small cutting force and rapid failure of tool make that only one fatigue source is observed at the rake face (Fig. 5a). Because the internal stress far away from the tool nose falls off rapidly, the crack deflection labeled by red arrows is found in Fig. 5c. Moreover, due to the asymmetrical stress, the fatigue steps with the secondary crack are also observed on the fatigue crack propagation zone. The fatigue steps can increase the roughness of fracture surface, which has a negative influence on the cutting forces and machined surface roughness. The transgranular fracture is the main fracture mode, as can be seen from Fig. 5f. It can be revealed that the grain interior is weaker than grain boundaries at high temperature. Because of the high cutting temperature, the adhesion wear and severer oxidation wear occur at the tool nose (Fig. 5d). Fig. 6 shows the two-dimensional micrographs of the rake face at vc = 90 m/min, vc = 150 m/min and vc = 180 m/min. The two-dimensional profiles, which cross the fatigue source (point 2), are plotted, as Table 2 Physical properties of SKD11 [16]. Density (kg/m3)
Young's modulus (GPa)
Poisson's ratio
Thermal conductivity (W/m·K)
Specific heat (J/kg °C)
8400
208
0.3
20.5
461
62
shown in Fig. 6b, d and f, respectively. From Fig. 6b, the chipping is found at the tool nose due to the large cutting forces. The fatigue source is located in the valley of the profile. It can be concluded that the fatigue cracks originate from the defect of tool substrate (point A in Fig. 2). The distance between the tool nose and fatigue source (point 1 to point 2) is about 200 μm, which is two times of the depth of cut. In addition, the fatigue crack path is also investigated (point 2 to point 3). Because the tool is subjected to severe mechanical stress at low speed, the fatigue crack path is relatively rough. However, as can be seen from Fig. 6d and f, the fatigue crack path is relatively smooth at vc = 150 m/min and vc = 180 m/min. It is attributed to the fact that the high cutting temperature decreases the resistance to crack propagation. 3.2. Cutting forces In order to investigate the influence of tool damage on the cutting forces, the cutting forces and flank wear value were measured when the tool goes through each pass of the workpiece surface. Fig. 7 presents the cutting force components and resultant cutting force for face milling SKD11 hardened steel at vc = 90 m/min, from which it can be seen that resultant cutting force reaches the peak when the tool cut into the workpiece. So the cutting forces are calculated based on the values extracted from the tool cut-into workpiece period. The average cutting force components are calculated by Eq. (1): Fx ¼
1 N 1 N 1 N ∑i¼1 Fxi ; Fy ¼ ∑i¼1 Fyi ; Fz ¼ ∑i¼1 Fzi N N N
ð1Þ
And the resultant cutting force Fr is calculated by the Eq. (2): qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Fri ¼ F2xi þF2yi þF2zi
ð2Þ
where Fxi, Fyi and Fzi are cutting force components in x, y, and z directions, respectively. Fig. 8 shows the flank wear and resultant cutting force versus volume of removed metal in milling SKD11. The flank wear rate at vc = 90 m/min is faster than that under the speed of 120 m/min in the initial tool wear stage (Fig. 8a), because of the larger cutting force at vc = 90 m/min as shown in Fig. 8b. It is known that the cutting temperature becomes higher when the cutting speed increases [18]. Therefore, it can be concluded that the cutting force has a larger influence on the tool wear than cutting temperature in the initial tool wear stage. After that, the coating peeling occurs at the rake face, which results in increase in both cutting force and cutting temperature. The cutting force at vc = 90 m/min is still larger than that at vc = 120 m/min. However, tool wear rate at vc = 90 m/min is slower than that under 120 m/min. It is inferred that the cutting temperature has a larger influence on the tool wear after the coating peeling. Fig. 9 depicts the cutting force components and resultant cutting force when face milling SKD11 in the initial wear stage. Obviously, it can be seen that the axial cutting force Fz are much larger than other cutting force components due to high hardness of workpiece. Furthermore, Table 3 Cutting parameters of face milling experiments. Cutting parameters
vc (m/min)
fz (mm/z)
ap (mm)
ae (mm)
90, 120, 150, 180
0.05
0.10
75
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F. Gong et al. / Int. Journal of Refractory Metals and Hard Materials 64 (2017) 27–34
Fig. 3. SEM micrographs and EDS analysis of worn tool (vc = 90 m/min, f = 0.05 mm/z, ap = 0.1 mm, ae = 75 mm).
the resultant cutting force firstly decreases and then increases with the increase of cutting speed from 90 m/min to 150 m/min. At last, it decreases when the cutting speed is above 150 m/min. The decrease in cutting force with the increasing of cutting speed up to 120 m/min can be explained by the decreasing of the friction coefficient between the tool and workpiece and the increasing shear angle with increasing cutting speed. However, when the speed exceeds 150 m/min, the thermal softening effect has a huge influence on the cutting forces, which also results in cutting force reduction.
To explore the influence of the tool damage on the cutting force, the growth rate of cutting force is defined as shown in Eq. (3) in the present study: δ¼
ð Ffin−FiniÞ 100% Fini
ð3Þ
where Fini and Ffin are the cutting forces when the tool goes through the first and last pass of workpiece.
Fig. 4. SEM micrographs and EDS analysis of worn tool (vc = 150 m/min, f = 0.05 mm/z, ap = 0.1 mm, ae = 75 mm).
F. Gong et al. / Int. Journal of Refractory Metals and Hard Materials 64 (2017) 27–34
Fig. 5. SEM micrographs and EDS analysis of worn tool (vc = 180 m/min, f = 0.05 mm/z, ap = 0.1 mm, ae = 75 mm).
Fig. 6. Two-dimensional micrographs of the rake face at (a) vc = 90 m/min (c) vc = 150 m/min (e) vc = 180 m/min.
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Fig. 7. Cutting force components and resultant cutting force in face milling of SKD11 at 90 m/min.
Fig. 10 shows the growth rate of the cutting force components and resultant cutting force. As shown in Fig. 10, with the increase of cutting speed, the cutting force firstly decreases until a minimum value reached at the cutting speed of 120 m/min, then increases when the cutting speed is beyond that. The irregular fracture surface at vc = 90 m/min (Fig. 3b) and 180 m/min (Fig. 5b) makes the ploughing rather than cutting occurs between the workpiece and tool [19], which leads to high growth rate of cutting force. 3.3. Machined surface roughness Fig. 11 shows the tool flank wear, resultant cutting force and surface roughness in milling of SKD11 at vc = 120 m/min. It can be seen that
variation trends of flank wear, cutting force and machined roughness value are identical. In the initial wear stage, the flank wear and cutting force are both stable, however, the roughness value has a jump increase. Although the value of flank wear is relatively low, the condition of rake face constantly deteriorates because of repeated impact load. The friction between the tool and workpiece is unstable, which results in the jumping change of surface roughness. In the steady wear stage, the flank wear and cutting forces increase rapidly, leading to the rapid increase in surface roughness. At last, when the flank wear reaches 0.31 mm, the machined surface roughness Ra rises from 0.15 μm to 0.86 μm. In addition, the deviation of roughness values is relatively large in the initial and severe wear stage. In the first pass, the tool flank wear
Fig. 8. (a) flank wear and (b) resultant cutting force versus volume of removed metal in milling of hardened steel SKD11.
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Fig. 9. Cutting force components and resultant cutting force when face milling of SKD11 in the initial wear stage.
changes in a wide range (rising from 0 to 0.05 mm) which results in a large deviation of roughness value. In the steady wear stage, the surface of cutting tool matches the cutting tracks of last pass, leading to the low deviation of roughness value. In the severe wear stage, the cutting tool has a serious breakage and is extremely prone to break. Therefore, the deviation of roughness value apparently increases. 4. Conclusion The failure patterns and failure mechanisms of coated carbide tool were investigated by high-speed face milling of the hardened steel SKD11. In addition, the influence of tool damage on cutting forces and surface roughness were also analyzed. The following conclusions can be drawn from these investigations: (1) The dominating failure pattern of coated tool when face milling of the hardened steel SKD11 was fatigue fracture accompanied by chipping. At vc = 90 m/min, the fracture caused by multi-fatigue crack propagation occurred on the rake face. With the increase of cutting speed, the fatigue striations and river patterns were observed at the fatigue crack propagation zone. At vc = 180 m/min, the fatigue steps with the secondary crack were also observed on the fatigue propagation area. The main fracture mode was intergranular fracture at lower cutting speeds, while it was transgranular fracture at higher cutting speeds. The adhesion and oxidation wear were presented at the tool nose at vc = 90 m/min–180 m/min.
Fig. 10. Growth rate of the cutting force components and resultant cutting force.
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Fig. 11. Tool flank wear, resultant cutting force and surface roughness in milling of SKD11 at vc = 120 m/min.
(2) The fatigue crack propagation paths were investigated for different cutting speeds. The fatigue cracks originated from the defect of tool substrate. The distance between the tool nose and fatigue source was approximately two times of the depth of cut. At low cutting speeds, the fatigue path was rough duo to severe mechanical stress. At high cutting speeds, because high cutting temperature decreased the resistance to crack propagation, the fatigue crack path was relatively smooth. (3) In the initial wear stage, the cutting force had a larger influence on tool life. However, when the tool breakage and coating peeling occurred, the cutting temperature was the dominating factor for tool life. In addition, the irregular fracture surfaces at vc = 90 m/min and 180 m/min contributed to a larger cutting force increment. (4) The machined surface roughness was significantly affected by the tool damage. With the flank wear and cutting forces increased, the value of machined surface roughness constantly increased. However, the deviation of roughness was relatively larger in the initial and severe wear stages, and it became smaller in the steady wear stage. Acknowledgements The authors would like to acknowledge the financial support from the National Natural Science Foundation of China (51475273). References [1] J.P. Davim, Machining of Hard Materials, Springer, London, 2011. [2] B. Wang, Z.Q. Liu, Cutting performance of solid ceramic end milling tools in machining hardened AISI H13 steel, Int. J. Refract. Met. Hard Mater. 55 (2016) 24–32. [3] J.P. Urbanski, P. Koshy, R.C. Dewes, D.K. Aspinwall, High speed machining of moulds and dies for net shape manufacture, Mater. Des. 21 (2000) 395–402. [4] J.A. Ghani, I.A. Choudhury, H.H. Masjuki, Wear mechanism of TiN coated carbide and uncoated cermets tools at high cutting speed applications, J. Mater. Process. Technol. 153-154 (2004) 1067–1073. [5] X.B. Cui, J. Zhao, Cutting performance of coated carbide tools in high-speed face milling of AISI H13 hardened steel, Int. J. Adv. Manuf. Technol. 71 (2014) 1811–1824. [6] C.Y. Wang, Y.X. Xie, Z. Qin, H.S. Lin, Y.H. Yuan, Q.M. Wang, Wear and breakage of TiAlN- and TiSiN-coated carbide tools during high-speed milling of hardened steel, Wear 336–337 (2015) 29–42. [7] Z.B. Yin, C.Z. Huang, B. Zou, H.L. Liu, H.T. Zhu, J. Wang, Dynamic fatigue behavior of Al2O3/TiC micro–nano-composite ceramic tool materials at ambient and high temperatures, Mater. Sci. Eng. A 593 (2014) 64–69. [8] Z.B. Yin, C.Z. Huang, J.T. Yuan, B. Zou, H.L. Liu, H.T. Zhu, Cutting performance and life prediction of an Al2O3/TiC micro–nano-composite ceramic tool when machining austenitic stainless steel, Ceram. Int. 41 (2015) 7059–7065. [9] G. Ramírez, E. Jiménez-Piqué, A. Mestra, M. Vilaseca, D. Casellas, L. Llanes, A comparative study of the contact fatigue behavior and associated damage micromechanisms of TiN-and WC:H-coated cold-work tool steel, Tribol. Int. 88 (2015) 263–270. [10] J.Y. Gui, S. Wei, Z.P. Xie, Slow crack growth behavior of silicon nitride ceramics in cryogenic environment, Ceram. Int. 42 (2016) 3687–3691.
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