- Email: [email protected]
Wear estimation of ceramic and coated carbide tools in turning of Inconel 625: 3D FE analysis
Author’s Accepted Manuscript Wear Estimation of Ceramic and Coated Carbide Tools in Turning of Inconel 625: 3D FE Analysis Mohammad Lotfi, Mehran Jahanbakhsh, Ali Akhavan Farid www.elsevier.com/locate/jtri
PII: DOI: Reference:
S0301-679X(16)30002-0 http://dx.doi.org/10.1016/j.triboint.2016.03.008 JTRI4114
To appear in: Tribiology International Received date: 2 December 2015 Revised date: 8 March 2016 Accepted date: 10 March 2016 Cite this article as: Mohammad Lotfi, Mehran Jahanbakhsh and Ali Akhavan Farid, Wear Estimation of Ceramic and Coated Carbide Tools in Turning of Inconel 625: 3D FE Analysis, Tribiology International, http://dx.doi.org/10.1016/j.triboint.2016.03.008 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 galley proof before it is published in its final citable 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 Estimation of Ceramic and Coated Carbide Tools in Turning of Inconel 625: 3D FE Analysis a
b
Mohammad Lotfi , Mehran Jahanbakhsh , Ali Akhavan Farid
c,d,
*
a
Department of Manufacturing, Faculty of Mechanical Engineering, University of Kashan, Kashan, Iran Department of Mechanical Engineering, Najafabad Branch, Islamic Azad University, Isfahan, Iran c Faculty of Engineering and Technology, Multimedia University, Jalan Ayer Keroh Lama, 75450 Melaka, Malaysia d Department of Manufacturing and Industrial Engineering, Faculty of Mechanical Engineering, Universiti Teknologi Malaysia, Johor, Malaysia b
* Corresponding author Email: [email protected] Telephone: 0060129337063 R2018- Faculty of Engineering and Technology Multimedia University 75450 Melaka Malaysia
Abstract Examination of cutting tool wear by experimental approaches is too costly. The main objective of this work is to develop an accurate 3D finite element model to predict the tool wear of PVD-TiAlN coated carbide and ceramic inserts in turning of Inconel 625. Thus, the cutting tools with complex geometries are modeled. Usui wear rate model is used to estimate wear rate where its constant parameters are achieved based on the cutting tools and workpiece material. The verification tests showed that the predicted values are in good agreement with the experiments. Moreover, among the cutting parameters, the increase of depth of cut were found as the most effective factor on the generation of temperature and stresses on the tool faces. Keywords: Machining; Wear; Ceramic; Finite-element method
1. Introduction Nickel-based superalloys are being used widely in the aerospace field and production of aircraft engine components particularly. The numerous usage is due to the existence of some considerable characteristics of nickel-based superalloys, namely, creep resistance, high-temperature corrosion resistance, and oxidation resistance [1]. In general, various types of machining operation, turning in particular, are being used in order to generate the desired shape of mechanical components. But turning of superalloys (as known difficult-to-cut materials) always faced with some difficulties in which rapid tool wear is the main one due to the rapid strain hardening, high strength as well as poor thermal conductivity [2, 3]. To overcome this problem, a proper comprehension of the metal cutting process and identification of effective cutting parameters could be helpful. Therefore, different theoretical, empirical, and simulation studies have been performed by researchers [4-6]. Among these methods, finite element (FE) simulation of turning operation can be more applicable because of the high cost of superalloys in experiments and complexity of metal cutting process that may not be modeled adequately in theoretical works. However, some theoretical models (e.g., friction, flow stress, wear) with the accurate elastic, plastic, and thermal parameters of materials are needed to be defined based on the experimental data
1
whereby simulation of turning operation and tool wear would be feasible. Thus, some studies [7-20] focused on the evaluation of aforementioned models are reviewed. Kagnaya et al. [7] analyzed the influence of flank and crater wear of uncoated carbide tools on cutting forces and temperature in turning of AISI 1045 steel by applying 2D FE simulation. After running the experiments, the parameters of flank and crater wear have been measured to model the worn tool in different configuration based on the cutting time. The results showed that the cutting forces and the temperature on the rake face increased as a result of increment of flank wear and crater wear, respectively. Yadav et al. [8] presented a 3D turning simulation of Inconel 718 superalloy using a flatfaced coated carbide insert in order to estimate flank wear and material removal rate. The tool and the workpiece material were defined from the library of DEFORM 3D. They used constant parameters of Usui’s [9] wear rate model obtained in the earlier research conducted by Yen et al. [10] in turning of AISI 1045 steel with uncoated carbide tool. However, it seems that this couldn’t be a right approach since there is a great difference between mechanical and thermal properties of Inconel 718 and AISI 1045 which directly affects on the tool wear rate. This means that constant parameters of a wear model obtained from experiments for a specific material cannot be used for another material. In another work, Lorentzon and Jarvstrat [11] developed an empirical tool wear model in machining of Inconel 718 with cemented carbide tool. Firstly, they attained tool wear rate experimentally. Besides, 2D simulation in the same condition was carried out to get the values of interface pressure, sliding velocity, and interface temperature for Usui’s wear rate model and absolute temperature for Takeyama and Murata [12] wear rate model. Then the constants of these models were computed by regression analysis. At the end, it was concluded that Usui’s empirical wear rate model gives good experimental agreement. Binder et al. [13] used finite element method to simulate crater wear in turning operation of AISI 1045 carbon steel with uncoated and PVD-TiAlN coated carbide. A modified Usui’s wear rate model has been applied where absolute maximal shear stress ( ) was used instead of normal stress ( ). The results indicate that the predicted crater wear is in good agreement with the experiments. However, the model underestimated the crater width. By implementation of different values of modified Coulomb friction model in orthogonal metal cutting, Shi et al. [14] found that the higher coefficient of friction model increases the cutting force, the tool-chip contact length and the maximum temperature at the tool edge. Attanasio et al.[15] considered both abrasive and diffusive wear phenomenon by combining Usui and Takeyama wear rate models in order to predict crater wear in a three-dimensional method. They used uncoated carbide tools in turning of AISI 1045 steel, and it was claimed that the simulation results are in good agreement with the experiments. Kurt et al. [16] studied the effect of tool-chip contact length on ceramic tool stresses generated in finish turning of AISI H13 hardened steel. At first, they calculated the tool-chip contact length by the model presented by Toropov and Ko [17]. Then they simulated the turning operation to estimate tensile and compression stresses on the rake and flank faces of the tool. As a result, it was stated that the highest tensile and compression stresses have been observed on the flank face and rake face, respectively. Furthermore, they concluded that these stresses increased by a reduction in tool-chip contact length. However, this conclusion could not be reliable due to the misinterpretation and substitution of depth of cut value of experiments instead of feed rate parameter as undeformed chip thickness in Toprove's equation. Ozel [18] examined a 3D simulation to investigate the effect of tool-edge geometry on tool stresses and tool wear in turning of AISI 4340 alloy steel using PCBN tools. He exerted shear friction model, Johnson-Cook flow stress model, and Usui’s wear rate model. He also defined the Usui constants by 2
trial-and-error with simulations. It was observed that the best stress distribution together with the lowest tool wear belongs to the variable edge geometry. Haddag et al. [19] analyzed the tribological behavior of tool-chip interface and associated tool wear in turning of a mild-steel part by coated carbide insert. Beside the experiments, they developed a 3D FE model to predict tool wear and heat generation on tool rake face. Finally, the comparison of results showed that the predicted values are reproduced well by the FE model. Hokka et al. [20] tried to achieve the constant values of Johnson-Cook model for Inconel 625 super alloy and Titanium-6246 by using a Split Hopkinson Pressure Bar (SHPB) method in order to model flow stress in simulation, where it was verified by 2D turning simulation and comparison of predicted cutting forces with experimental results. As it was reviewed, neither an experimental nor a simulation work has been reported regarding the tool wear investigation in machining of Inconel 625 superalloy. Therefore, the aim of present study is to estimate the tool wear in turning of Inconel 625 by using 3D finite element simulation and experimental validation. In this study, PVD-TiAlN coated carbide and ceramic inserts were used and all strengths such as accurate wear models and weaknesses such as simplifications and inaccurate model parameters of previous works explained in this section are taken into account. Furthermore, the effect of cutting parameters on the tool wear, temperature and stress distribution on the tool faces are studied.
2. FE modelling and experimental setup A SiC whisker-reinforced Al2O3 ceramic insert with geometry of RNGN1907 and a PVD-TiAlN coated carbide inserts (SNMG 190616) with the complex chip breaker (QM) are selected for this study. These tool materials and geometries are recommended by Sandvik Coromant catalogue for machining of nickel-based superalloys [21]. In order to model coated carbide insert with a complex chip breaker geometry which has great effect on chip formation, a 3D-scanner device was used to obtain cloud of points of cutting insert, and then it was modeled by using a CAD software. Owing to the non-complexity in the geometry of ceramic insert, it was directly designed in the CAD software. Besides, the workpieces were also modeled based on the cutting conditions and tool geometries. Afterward, the STL files of the workpieces and the cutting inserts were imported to DEFORM-3D [22] (Fig. 1).
3
Fig. 1.a) the cutting inserts, b) the modeled inserts, c) the meshed inserts and workpieces, and d) the boundary condition
The material properties of cutting inserts were exerted from DEFORM’s library. Whereas, the workpiece material properties which is Inconel 625 were imported to the software. Accordingly, a new material was defined and all elastic, plastic, and thermal properties of Inconel 625 were imported to DEFORM-3D. Graphs in Fig. 2 which express mechanical properties of Inconel 625 in the range of operating temperature are constructed based on the cloud of points collected from literatures [20, 23-25].
4
Fig. 2. Inconel 625 material properties and its Johnson-Cook constant parameters.
In this work, DEFORM-3D software with Sparse solver was applied to simulate turning operation. Simulation process was carried out in 3500 steps. Each time-step included 1.3e-5 second which was proportional to 0.0165 mm of cutting length for each stroke-step. Totally, 60 mm length was determined for the workpiece and approximately 0.04s was simulated in each particular cutting condition. An updated Lagrangian formulation has been used to simulate chip formation and to remesh the workpiece when the elements of the mesh are too distorted. The workpiece was meshed based on feed value with tetrahedron mesh type. It means that one fourth of feed value in each particular cutting conditions was defined as minimum element size, in which the size ratio was 7. In this method, the number of elements has been varied from 10000 to 12000 regarding feed value. Moreover, higher mesh density with the size of 5e-4 compared to other areas was applied in the cutting zone due to large gradients of strain, strain rate, and temperature in this area (see Fig.1-c). While, the cutting tool was defined as a rigid body, the heat transfer mode was activated to analyze temperature on the tool faces. Also, the workpiece was defined as a plastic object where the Johnson-Cook flow stress model (Eq. (1)) was used to represent the workpiece material constitutive behavior. ̅
[
( )̅ ] [
̅̇
( ̅̇ )] [
(
) ]
(1)
where A, B, C, n, and m are the yield strength, the hardening modulus, the strain rate sensitivity, the strain-hardening, and the thermal softening exponent, respectively (given in Fig. 2). Also ̅ is the equivalent flow stress, ̅ is the equivalent strain, ̇ ̅ is the plastic strain rate, ̅̇ is the reference of plastic 5
strain rate, T is the temperature, Tm is the melting temperature, and Tr is the reference temperature (20°C) [26]. To model the friction in the tool-workpiece contact, the constant shear model was utilized (Eq. (2)). In this equation, k is the shear flow stress of the working material at the tool-chip interface and m is the constant shear friction factor which defined equal to 0.6. However, an iterative procedure should be carried out if the predicted values are not in good agreement with the experiments [27]. (2) There are several wear rate models to predict tool wear, as reviewed in the literature [9,12,13]. In this work, Usui’s wear rate model was used to calculate wear rate on the tool surfaces. According to Eq. (3), Usui model is established based on the normal stress ( ), sliding velocity (Vs), interface temperature (T), and two constant parameters A and B [28]. (
⁄ )
(3)
The variation in cutting parameters can be translated to the variation of model variables where changes in cutting speed results in sliding velocity alteration, and normal stress will be altered by changes in feed rate and depth of cut. An accurate value for constants of A and B, applicable to all cutting parameters, can be found if there is a strong relationship between cutting parameters and wear rate. This can be measured by a coefficient of determination that will be attained in a set of cutting experiments. For this purpose, experiments were designed based on the Central Composite Design (CCD) that allows fitting a polynomial curve on the values of wear rate. CCD is a fractional factorial design with five level. The design includes 2k factorial experiments, 2k axial points where k is the number of variables [29]. Table 1 shows the levels of factorial (+1 and -1) and axial point (-alpha, +alpha) where α=√ is selected to attain the rotatable condition in the design. Consequently, constants of A and B should be calibrated in an iteration comparison of the wear rate obtained in both simulation and experimental cutting process (Fig.3). The iteration process will be stopped when the constants of A and B result in the prediction error of less than ten percent in all cutting parameters. Table 1. The level of cutting parameters. Type Parameter
Unit
(Vc) (f) (ap)
Coated carbide
Ceramic
Levels
Levels
-
-1
0
+1
-
-1
0
+1
(m/min) (mm/rev)
32.96
50
75
100
117.04
65.91
100
150
200
234.09
0.1
0.14
0.2
(mm)
0.32
0.8
1.5
0.26
0.3
0.07
0.1
0.15
0.2
0.23
2.2
2.68
0.46
0.7
1.05
1.4
1.64
6
Fig. 3. Flow chart for determination of the Usui constant values.
Both Flank and crater wear have been investigated in this study. However, flank wear (VB) was selected as tool wear criterion. In order to assign the cutting length, some trial-and-error tests were conducted considering ISO-standard defined 0.6 mm for maximum flank wear (VB-max). Therefore, 70 and 50 mm cutting length were determined in turning by using coated carbide and ceramic tools, respectively. Fig. 4 shows a schematic tool wear profile. Tool wear profile commonly divided into three zones. Zone C is related to the tool nose. Zone N is a quarter of length b in the worn cutting edge. Zone B is the remained part that VB and VB-max are measured in this region. Furthermore, tool nose and notch wear bandwidth are named with VC and VN, respectively [30].
7
Fig. 4. Tool wear profile.
The experimental tests were conducted on a TC35 CNC machine with the transversal speed range from 1 to 3500 rpm. Also, an optical microscope and a scanning electron microscope (SEM) used to measure and validate tool wear estimation. The rake ( ) and clearance ( ) angles of both ceramic and coated carbide inserts mounted on tool holder were -6 and +6 degrees regardless of their chip breaker angles. Furthermore, the lead angle ( ) of coated carbide insert was 45°, while it was variable for ceramic insert due to its round shape. The turning tests were performed on a bar with 100 mm diameter. Equipment applied in experiments are illustrated in Fig. 5.
Fig. 5.Applied instruments.
3. Results The cutting parameters of experimental runs designed based on CCD and flank wear value of each run is listed in Table 2. The experimental cutting times (Eq. (4)) were used to compute the flank wear rates (dw/dt) for all cutting conditions in order to make results compatible with simulation outputs. 8
Table 2. Design of experiment and results. Type Run 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15
Coated carbide
Ceramic
Vc
f
ap
VB
Vc
f
ap
VB
(m/min)
(mm/rev)
(mm)
(mm)
(m/min)
(mm/rev)
(mm)
(mm)
50 100 50 100 50 100 50 100 32.96 117.04 75 75 75 75 75
0.14 0.14 0.26 0.26 0.14 0.14 0.26 0.26 0.2 0.2 0.1 0.3 0.2 0.2 0.2
0.8 0.8 0.8 0.8 2.2 2.2 2.2 2.2 1.5 1.5 1.5 1.5 0.32 2.68 1.5
0.125 0.271 0.163 0.292 0.245 0.725 0.187 0.604 0.063 0.500 0.388 0.312 0.030 0.542 0.343
100 200 100 200 100 200 100 200 65.91 234.09 150 150 150 150 150
0.1 0.1 0.2 0.2 0.1 0.1 0.2 0.2 0.15 0.15 0.07 0.23 0.15 0.15 0.15
0.7 0.7 0.7 0.7 1.4 1.4 1.4 1.4 1.05 1.05 1.05 1.05 0.46 1.64 1.05
0.396 0.334 0.478 0.291 0.459 0.167 0.627 0.479 0.564 0.847 0.589 0.311 0.231 0.459 0.312
(4) where t (s) is machining time, l (mm) is cutting length, n (rpm) is rotational speed of spindle, and f (mm/rev) is the feed rate. The coefficient of determination of experimental results for carbide and ceramic inserts is R2=0.954 and R2=0.913, respectively. This reveals that there is a strong correlation between independent and dependent variables where any changes in cutting parameters are always accompanied by changes in the flank wear rate. Therefore, it is evident that experimental results are reliable to be utilized in the iterative calibration method explained in section 2 in order to achieve the accurate constants of A and B as follow:
(
) ,
for TiAlN-coated carbide
(
) ,
for SiC-reinforced-Al2O3
Fig.6 compares the predicted and actual values of flank wear rate obtained in simulation and experiments, respectively. As can be seen, the predicted values are close to the experiments which verify the accuracy of A and B constants and the adequacy of finite element model developed.
9
Fig. 6. Comparison of predicted and experimental valuesof flank wear rate.
4. Discussions Based on the experimental and simulation results, depth of cut is the most influential factor on the flank wear compared to the other cutting parameters in turning of Inconel 625 superalloy. As can be seen in Fig. 7, the size of flank wear increased by an increment of depth of cut for both ceramic and coated carbide inserts, while the effect of feed rate increment on the flank wear for carbide tools is almost insignificant. This condition can be helpful to enhance productivity by utilizing higher range of feed rate which results in higher material removal rate. Also, the influence of cutting speed on the flank wear is similar to the effect of depth of cut when coated carbide tool was used. But this factor is not so effective on the flank wear of ceramic tool. The lower friction coefficient of ceramic tools compared to the carbide tools, their anti-adhesion characteristic, and chemical stability under 1200 0C made them suitable for applying in high cutting speeds [31].
Fig. 7. Effect of cutting parameters on the flank wear (VB).
The effect of depth of cut increment on the propagation of flank wear was investigated by considering generated temperature and stresses on tool faces shown in Figs. 8 and 9. In the machining operation, the energy of plastic deformation in the primary cutting zone is converted into the heat, and then with progress of cutting process, frictional heat is generated at the tool-chip interface (secondary deformed zone)[32]. A higher range of depth of cut corresponds with the temperature rise in tool faces due to the 10
higher plastic work requirement. Owing to the poor thermal conductivity of Inconel 625 and then low heat dissipation, temperature rises rapidly in the cutting zone and tool which finally results in faster tool wear on the tool surfaces (Figs. 8c and 9c). Furthermore, effective stress increased on the edge and faces of inserts by an increase of depth of cut which is completely distributed at the tool-chip contact zone (Figs. 8d and 9d). Accordingly, tool wear rate increased at the areas that more stress concentration is seen (illustrated in Figs. 8e and 9e). It should be noted that in order to compute stress-effective in simulations, the cutting tools were defined as elastic objects and the turning simulations were carried out under the same conditions described earlier. Besides, the severe chipping of cutting edge in the ceramic tool at the highest depth of cut can be seen in Fig.9g-3 which is due to the sudden engagement and impact imposed on that. Ceramic tools are not able to withstand against the impact of cutting forces due to their low fracture toughness that negative rake angles often applied to avoid chipping phenomenon [33].
11
Fig. 8.(a)The settings of coated carbide tool, (b) deformed chips, (c) the temperature distribution, (d) the stress distribution, (e) tool wear rate, (f) the experimental tool wear, ((g-1), (g-2) and (g-3)) FEA, optical and SEM images of flank wear (f=0.2 mm/rev, VC=75 m/min).
12
Fig. 9. (a) The settings of ceramic tool, (b) deformed chips, (c) the temperature distribution, (d) the stress distribution, (e) tool wear rate, (f) the experimental tool wear, ((g-1), (g-2) and (g-3)) FEA, optical and SEM images of flank wear (f=0.15 mm/rev, VC=150 m/min).
13
Crater wear is another important tool failure mode. However, as it was observed that flank wear decreases with an increase of feed rate, the increase of this parameter in order to enhance productivity regardless of crater wear concerns could be an unconsidered plan. Thus, the effect of feed rate variations on the crater wear is studied (Fig. 10). Usually, length of tool-chip contact ( ) is taken into account in order to analysis of crater wear [17, 34]. Undeformed chip thickness ( ) which is directly related to the feed rate (Eq. (5)), is the most important factor in the definition of tool-chip contact length regarding the Eq. (6) presented by Toropov and Ko. (5) (
(
where x, , and
)
)
(6)
are the cutting edge angle, tool rake angle and shear angle, respectively.
Accordingly, the increase of undeformed chip thickness as a result of feed rate increment caused further tool-chip contact length, seen in Fig. 10-b.This condition intensified wear depth at higher level of feed rate for both ceramic and coated carbide inserts. Fig. 10-c indicates generated tool crater wear in the same step at three levels of feed rate. Furthermore, Fig. 10-d shows optical microscope images of crater wear at tool rake face in the experiment. Eventually, the investigation of the effect of cutting parameters on the flank and crater wear by using FE simulation and experimental work revealed that low level of depth of cut and cutting speed with middle level of feed rate is a proper condition for coated carbide insert in turning of Inconel 625. Also the low depth of cut, high cutting speed together with the middle level of feed rate represents an optimum tool wear condition for the ceramic tool.
14
Fig. 10.(a)Turning operation at different feed rates (middle level of depth of cut and cutting speed), (b) deformed chip thickness, (c) predicted crater wear, and (d) experimental crater wear.
15
4. Concluding remarks In this study, a 3D finite element simulation method has been applied to predict tool wear in machining of Inconel 625 superalloy. Accordingly, Usui’s wear rate model was exerted and its constant parameters were calibrated by experiments carried out on PVD-TiAlN coated carbide and ceramic inserts. The main conclusions of this work are listed as follow:
Comparison of tool wear rate shows predicted and experimental values are in good agreement. However, it should be noted that constant A and B are valid in the range of cutting parameters selected in experiments, and might need to be calibrated again if a higher range of cutting parameters are desired. The simulation results indicate that tool wear rate is higher in the areas in which temperature and stresses are concentrated. The values of temperature and stresses increase by an increment in the level of depth of cut. The increase of depth of cut severely propagates the flank wear for both tools. The effect of cutting speed on the flank wear of ceramic tool is insignificant due to its heat resistance characteristic. However, the increase of cutting speed affects tool flank wear of coated carbide as the depth of cut does. Variation of feed rate is almost ineffective on the flank wear, while it intensifies crater wear due to the increase of undeformed chip thickness resulting the increase of tool-chip contact pressure. It can be concluded that low depth of cut and cutting speed with middle level of feed rate are appropriate cutting conditions for PVD-TiAlN coated carbide insert, while ceramic insert should be applied under condition of low depth of cut, middle level of feed rate and high cutting speed.
Acknowledgements The authors wish to thank the Multimedia University for funding this project under Mini-fund scheme (MMUI/160043) and the Production Lab of University of Kashan for their technical supports. Assistance of manufacturing lab-UTM is also greatly appreciated which provided the DEFORM license.
References [1] J. L. Cantero, J. Díaz-Álvarez, M.H. Miguélez, N.C.Marín, Analysis of tool wear patterns in finishing turning of Inconel 718, Wear, 297(1)(2013) 885-894. [2] A. Bhatt, H. Attia, R. Vargas, V. Thomson, Wear mechanisms of WC coated and uncoated tools in finish turning of Inconel 718. Tribology International, 43(5) (2010) 1113-1121. [3] C. Courbon, F. Pusavec, F. Dumont, J. Rech, and J. Kopac. Tribological behaviour of Ti6Al4V and Inconel718 under dry and cryogenic conditions—Application to the context of machining with carbide tools. Tribology International 66 (2013): 72-82. [4] A. Devillez, G. Le Coz, S. Dominiak, D. Dudzinski, Dry machining of Inconel 718, workpiece surface integrity. Journal of Materials Processing Technology 211(10) (2011) 1590-1598.
16
[5] Z. Hao, D. Gao, Y. Fan, R. Han, New observations on tool wear mechanism in dry machining Inconel718. International Journal of Machine Tools and Manufacture, 51(12) (2011) 973-979. [6] Zhuang, K., Zhu, D., Zhang, X., Ding, H. Notch wear prediction model in turning of Inconel 718 with ceramic tools considering the influence of work hardened layer, Wear, 313(1) (2014) 63-74. [7]T. Kagnaya, L. Lambert, M. Lazard, C.Boher, T. Cutard, Investigation and FEA-based simulation of tool wear geometry and metal oxide effect on cutting process variables, Simulation Modelling Practice and Theory, 42 (2014) 84-97. [8] R.K. Yadav, K. Abhishek, S.S. Mahapatra, A simulation approach for estimating flank wear and material removal rate in turning of Inconel 718, Simulation Modelling Practice and Theory, 52 (2015) 1-14. [9] E. Usui, T. Shirakashi, T. Kitagawa, Analytical prediction of cutting tool wear, Wear, 100(1) (1984) 129-151. [10] Y.C. Yen, J. Söhner, B. Lilly, T. Altan, Estimation of tool wear in orthogonal cutting using the finite element analysis, J. Mater. Proc. Technol. 146(1) (2004) 82-91. [11] J. Lorentzon, N. Järvstråt, Modelling tool wear in cemented-carbide machining alloy 718, Int. J. Mach. Tools Manuf. 48(10) (2008) 1072-1080. [12] H. Takeyama, R. Murata, Basic investigation of tool wear, J. Manuf. Sci. Eng. 85(1) (1963) 33-37. [13] M. Binder, F. Klocke, D. Lung, Tool wear simulation of complex shaped coated cutting tools, Wear, 330 (2015) 600607. [14] G. Shi, X. Deng, C. Shet, A finite element study of the effect of friction in orthogonal metal cutting, Finite Elements in Analysis and Design, 38(9) (2002) 863-883. [15] A. Attanasio, E. Ceretti, A. Fiorentino, C. Cappellini, C. Giardini, Investigation and FEM-based simulation of tool wear in turning operations with uncoated carbide tools, Wear, 269(5) (2010) 344-350. [16] A. Kurt, B. Yalçin, N. Yilmaz, The cutting tool stresses in finish turning of hardened steel with mixed ceramic tool, Int. J. Adv. Manuf. Technol. (2015) 1-11. [17] A. Toropov, S.L. Ko, Prediction of tool-chip contact length using a new slip-line solution for orthogonal cutting, Int. J. Mach. Tools Manuf. 43(12) (2003) 1209-1215. [18] T. Özel, Computational modelling of 3D turning: Influence of edge micro-geometry on forces, stresses, friction and tool wear in PcBN tooling, J. Mater. Proc. Technol. 209(11) (2009) 5167-5177. [19] B. Haddag, H. Makich, M. Nouari, and J. Dhers. Tribological behaviour and tool wear analyses in rough turning of large-scale parts of nuclear power plants using grooved coated insert. Tribology International 80 (2014) 58-70. [20] M. Hokka, D. Gomon, A. Shrot, T. Leemet, M. Bäker, V.T. Kuokkala, Dynamic Behavior and High Speed Machining of Ti-6246 and Alloy 625 Super alloys: Experimental and Modeling Approaches, Experimental Mechanics, 54(2) (2014) 199-210. [21] Technical Committee, Coro Guide. In Coromant, S., Ed. Sandvik Coromant, 2014.
17
[22] DEFORM-3D [Computer software]. (2012). http:// www.deform.com [23] A. Vesel, A. Drenik, K. Elersic, M. Mozetic, J. Kovac, T. Gyergyek, J. Stockel, J. Varju, R. Panek, M. Balat-Pichelin, Oxidation of Inconel 625 super alloy upon treatment with oxygen or hydrogen plasma at high temperature, Applied Surface Science, 305 (2014) 674-682. [24] http://www.specialmetals.com/inconel-alloy-625 [25] http://www.jacquet.biz/JACQUET/USA/files/JCQusa-alloy-625.pdf [26] T. Özel, The influence of friction models on finite element simulations of machining, Int. J. Mach. Tools Manuf. 46(5) (2006) 518-530. [27] M. Lotfi, A.A. Farid, H. Soleimanimehr, The effect of chip breaker geometry on chip shape, bending moment, and cutting force: FE analysis and experimental study, Int. J. Adv. Manuf. Technol. 78(5-8) (2015) 917-925. [28] L.J.Xie, J. Schmidt, C. Schmidt, F. Biesinger, 2D FEM estimate of tool wear in turning operation, Wear, 258(10) (2005) 1479-1490. [29] A. A. Farid, S. Sharif, S. Alizadeh Ashrafi, M. H. Idris, Statistical Analysis, Modeling, and Optimization of Thrust Force and Surface Roughness in High-Speed Drilling of Al–Si Alloy. Proc. Inst. Mech. Eng. Part B-J. Eng. Manuf. 227 (6) (2013) 808-20. [30] B. Li, A review of tool wear estimation using theoretical analysis and numerical simulation technologies. International Journal of Refractory Metals and Hard Materials 35 (2012) 143-151. [31] A. Altin, M. Nalbant, A. Taskesen, The effects of cutting speed on tool wear and tool life when machining Inconel 718 with ceramic tools, Materials & design, 28(9) (2007) 2518-2522. [32] E. Muthu, K. Senthamarai, S. Jayabal, Finite element simulation in machining of Inconel 718 nickel based super alloy, Int. J of Adv. Eng. Applic,1(3) (2012) 22-27. [33]http://www.tech.plym.ac.uk/sme/mfmt201/cuttingtoolmats.htm [34] M. Lotfi, A.A. Farid, H. Soleimanimehr, A new hybrid model based on the radius ratio for prediction of effective cutting limit of chip breakers, Proc. IMechE Part B: J. of Eng. Manuf. (2015), 0954405415586549.
18
Table captions: Table 1. The level of cutting parameters. Table 2. Design of experiment and results.
Figure captions: Fig. 1. a) The cutting inserts, b) the modeled inserts, c) the meshed inserts and workpieces, and d) the boundary condition. Fig. 2. Inconel 625 material properties and its Johnson-Cook constant parameters. Fig. 3. Flow chart for determination of the Usui constant values. Fig. 4. Tool wear profile. Fig. 5.Applied instruments. Fig. 6. Comparison of predicted and experimental values of flank wear rate. Fig. 7. Effect of cutting parameters on the flank wear (VB). Fig. 8. (a) The settings of coated carbide tool, (b) deformed chips, (c) the temperature distribution, (d) the stress distribution, (e) tool wear rate, (f) the experimental tool wear, ((g-1), (g-2) and (g-3)) FEA, optical and SEM images of flank wear (f=0.2 mm/rev, VC=75 m/min). Fig. 9. (a) The settings of ceramic tool, (b) deformed chips, (c) the temperature distribution, (d) the stress distribution, (e) tool wear rate, (f) the experimental tool wear, ((g-1), (g-2) and (g-3)) FEA, optical and SEM images of flank wear (f=0.15 mm/rev, VC=150 m/min). Fig. 10. (a)Turning operation at different feed rates (middle level of depth of cut and cutting speed), (b) deformed chip thickness, (c) predicted crater wear, and (d) experimental crater wear.
Highlights:
A 3D finite element simulation of tool wear in turning of Inconel 625 is carried out. PVD-TiAlN coated carbide and ceramic inserts are used. The effect of cutting parameters on the temperature, stress, and tool wear are studied. The increment in the level of depth of cut intensifies the flank wear propagation. The increase of tool-chip contact length increases crater wear of cutting tools.
19
PII: DOI: Reference:
S0301-679X(16)30002-0 http://dx.doi.org/10.1016/j.triboint.2016.03.008 JTRI4114
To appear in: Tribiology International Received date: 2 December 2015 Revised date: 8 March 2016 Accepted date: 10 March 2016 Cite this article as: Mohammad Lotfi, Mehran Jahanbakhsh and Ali Akhavan Farid, Wear Estimation of Ceramic and Coated Carbide Tools in Turning of Inconel 625: 3D FE Analysis, Tribiology International, http://dx.doi.org/10.1016/j.triboint.2016.03.008 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 galley proof before it is published in its final citable 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 Estimation of Ceramic and Coated Carbide Tools in Turning of Inconel 625: 3D FE Analysis a
b
Mohammad Lotfi , Mehran Jahanbakhsh , Ali Akhavan Farid
c,d,
*
a
Department of Manufacturing, Faculty of Mechanical Engineering, University of Kashan, Kashan, Iran Department of Mechanical Engineering, Najafabad Branch, Islamic Azad University, Isfahan, Iran c Faculty of Engineering and Technology, Multimedia University, Jalan Ayer Keroh Lama, 75450 Melaka, Malaysia d Department of Manufacturing and Industrial Engineering, Faculty of Mechanical Engineering, Universiti Teknologi Malaysia, Johor, Malaysia b
* Corresponding author Email: [email protected] Telephone: 0060129337063 R2018- Faculty of Engineering and Technology Multimedia University 75450 Melaka Malaysia
Abstract Examination of cutting tool wear by experimental approaches is too costly. The main objective of this work is to develop an accurate 3D finite element model to predict the tool wear of PVD-TiAlN coated carbide and ceramic inserts in turning of Inconel 625. Thus, the cutting tools with complex geometries are modeled. Usui wear rate model is used to estimate wear rate where its constant parameters are achieved based on the cutting tools and workpiece material. The verification tests showed that the predicted values are in good agreement with the experiments. Moreover, among the cutting parameters, the increase of depth of cut were found as the most effective factor on the generation of temperature and stresses on the tool faces. Keywords: Machining; Wear; Ceramic; Finite-element method
1. Introduction Nickel-based superalloys are being used widely in the aerospace field and production of aircraft engine components particularly. The numerous usage is due to the existence of some considerable characteristics of nickel-based superalloys, namely, creep resistance, high-temperature corrosion resistance, and oxidation resistance [1]. In general, various types of machining operation, turning in particular, are being used in order to generate the desired shape of mechanical components. But turning of superalloys (as known difficult-to-cut materials) always faced with some difficulties in which rapid tool wear is the main one due to the rapid strain hardening, high strength as well as poor thermal conductivity [2, 3]. To overcome this problem, a proper comprehension of the metal cutting process and identification of effective cutting parameters could be helpful. Therefore, different theoretical, empirical, and simulation studies have been performed by researchers [4-6]. Among these methods, finite element (FE) simulation of turning operation can be more applicable because of the high cost of superalloys in experiments and complexity of metal cutting process that may not be modeled adequately in theoretical works. However, some theoretical models (e.g., friction, flow stress, wear) with the accurate elastic, plastic, and thermal parameters of materials are needed to be defined based on the experimental data
1
whereby simulation of turning operation and tool wear would be feasible. Thus, some studies [7-20] focused on the evaluation of aforementioned models are reviewed. Kagnaya et al. [7] analyzed the influence of flank and crater wear of uncoated carbide tools on cutting forces and temperature in turning of AISI 1045 steel by applying 2D FE simulation. After running the experiments, the parameters of flank and crater wear have been measured to model the worn tool in different configuration based on the cutting time. The results showed that the cutting forces and the temperature on the rake face increased as a result of increment of flank wear and crater wear, respectively. Yadav et al. [8] presented a 3D turning simulation of Inconel 718 superalloy using a flatfaced coated carbide insert in order to estimate flank wear and material removal rate. The tool and the workpiece material were defined from the library of DEFORM 3D. They used constant parameters of Usui’s [9] wear rate model obtained in the earlier research conducted by Yen et al. [10] in turning of AISI 1045 steel with uncoated carbide tool. However, it seems that this couldn’t be a right approach since there is a great difference between mechanical and thermal properties of Inconel 718 and AISI 1045 which directly affects on the tool wear rate. This means that constant parameters of a wear model obtained from experiments for a specific material cannot be used for another material. In another work, Lorentzon and Jarvstrat [11] developed an empirical tool wear model in machining of Inconel 718 with cemented carbide tool. Firstly, they attained tool wear rate experimentally. Besides, 2D simulation in the same condition was carried out to get the values of interface pressure, sliding velocity, and interface temperature for Usui’s wear rate model and absolute temperature for Takeyama and Murata [12] wear rate model. Then the constants of these models were computed by regression analysis. At the end, it was concluded that Usui’s empirical wear rate model gives good experimental agreement. Binder et al. [13] used finite element method to simulate crater wear in turning operation of AISI 1045 carbon steel with uncoated and PVD-TiAlN coated carbide. A modified Usui’s wear rate model has been applied where absolute maximal shear stress ( ) was used instead of normal stress ( ). The results indicate that the predicted crater wear is in good agreement with the experiments. However, the model underestimated the crater width. By implementation of different values of modified Coulomb friction model in orthogonal metal cutting, Shi et al. [14] found that the higher coefficient of friction model increases the cutting force, the tool-chip contact length and the maximum temperature at the tool edge. Attanasio et al.[15] considered both abrasive and diffusive wear phenomenon by combining Usui and Takeyama wear rate models in order to predict crater wear in a three-dimensional method. They used uncoated carbide tools in turning of AISI 1045 steel, and it was claimed that the simulation results are in good agreement with the experiments. Kurt et al. [16] studied the effect of tool-chip contact length on ceramic tool stresses generated in finish turning of AISI H13 hardened steel. At first, they calculated the tool-chip contact length by the model presented by Toropov and Ko [17]. Then they simulated the turning operation to estimate tensile and compression stresses on the rake and flank faces of the tool. As a result, it was stated that the highest tensile and compression stresses have been observed on the flank face and rake face, respectively. Furthermore, they concluded that these stresses increased by a reduction in tool-chip contact length. However, this conclusion could not be reliable due to the misinterpretation and substitution of depth of cut value of experiments instead of feed rate parameter as undeformed chip thickness in Toprove's equation. Ozel [18] examined a 3D simulation to investigate the effect of tool-edge geometry on tool stresses and tool wear in turning of AISI 4340 alloy steel using PCBN tools. He exerted shear friction model, Johnson-Cook flow stress model, and Usui’s wear rate model. He also defined the Usui constants by 2
trial-and-error with simulations. It was observed that the best stress distribution together with the lowest tool wear belongs to the variable edge geometry. Haddag et al. [19] analyzed the tribological behavior of tool-chip interface and associated tool wear in turning of a mild-steel part by coated carbide insert. Beside the experiments, they developed a 3D FE model to predict tool wear and heat generation on tool rake face. Finally, the comparison of results showed that the predicted values are reproduced well by the FE model. Hokka et al. [20] tried to achieve the constant values of Johnson-Cook model for Inconel 625 super alloy and Titanium-6246 by using a Split Hopkinson Pressure Bar (SHPB) method in order to model flow stress in simulation, where it was verified by 2D turning simulation and comparison of predicted cutting forces with experimental results. As it was reviewed, neither an experimental nor a simulation work has been reported regarding the tool wear investigation in machining of Inconel 625 superalloy. Therefore, the aim of present study is to estimate the tool wear in turning of Inconel 625 by using 3D finite element simulation and experimental validation. In this study, PVD-TiAlN coated carbide and ceramic inserts were used and all strengths such as accurate wear models and weaknesses such as simplifications and inaccurate model parameters of previous works explained in this section are taken into account. Furthermore, the effect of cutting parameters on the tool wear, temperature and stress distribution on the tool faces are studied.
2. FE modelling and experimental setup A SiC whisker-reinforced Al2O3 ceramic insert with geometry of RNGN1907 and a PVD-TiAlN coated carbide inserts (SNMG 190616) with the complex chip breaker (QM) are selected for this study. These tool materials and geometries are recommended by Sandvik Coromant catalogue for machining of nickel-based superalloys [21]. In order to model coated carbide insert with a complex chip breaker geometry which has great effect on chip formation, a 3D-scanner device was used to obtain cloud of points of cutting insert, and then it was modeled by using a CAD software. Owing to the non-complexity in the geometry of ceramic insert, it was directly designed in the CAD software. Besides, the workpieces were also modeled based on the cutting conditions and tool geometries. Afterward, the STL files of the workpieces and the cutting inserts were imported to DEFORM-3D [22] (Fig. 1).
3
Fig. 1.a) the cutting inserts, b) the modeled inserts, c) the meshed inserts and workpieces, and d) the boundary condition
The material properties of cutting inserts were exerted from DEFORM’s library. Whereas, the workpiece material properties which is Inconel 625 were imported to the software. Accordingly, a new material was defined and all elastic, plastic, and thermal properties of Inconel 625 were imported to DEFORM-3D. Graphs in Fig. 2 which express mechanical properties of Inconel 625 in the range of operating temperature are constructed based on the cloud of points collected from literatures [20, 23-25].
4
Fig. 2. Inconel 625 material properties and its Johnson-Cook constant parameters.
In this work, DEFORM-3D software with Sparse solver was applied to simulate turning operation. Simulation process was carried out in 3500 steps. Each time-step included 1.3e-5 second which was proportional to 0.0165 mm of cutting length for each stroke-step. Totally, 60 mm length was determined for the workpiece and approximately 0.04s was simulated in each particular cutting condition. An updated Lagrangian formulation has been used to simulate chip formation and to remesh the workpiece when the elements of the mesh are too distorted. The workpiece was meshed based on feed value with tetrahedron mesh type. It means that one fourth of feed value in each particular cutting conditions was defined as minimum element size, in which the size ratio was 7. In this method, the number of elements has been varied from 10000 to 12000 regarding feed value. Moreover, higher mesh density with the size of 5e-4 compared to other areas was applied in the cutting zone due to large gradients of strain, strain rate, and temperature in this area (see Fig.1-c). While, the cutting tool was defined as a rigid body, the heat transfer mode was activated to analyze temperature on the tool faces. Also, the workpiece was defined as a plastic object where the Johnson-Cook flow stress model (Eq. (1)) was used to represent the workpiece material constitutive behavior. ̅
[
( )̅ ] [
̅̇
( ̅̇ )] [
(
) ]
(1)
where A, B, C, n, and m are the yield strength, the hardening modulus, the strain rate sensitivity, the strain-hardening, and the thermal softening exponent, respectively (given in Fig. 2). Also ̅ is the equivalent flow stress, ̅ is the equivalent strain, ̇ ̅ is the plastic strain rate, ̅̇ is the reference of plastic 5
strain rate, T is the temperature, Tm is the melting temperature, and Tr is the reference temperature (20°C) [26]. To model the friction in the tool-workpiece contact, the constant shear model was utilized (Eq. (2)). In this equation, k is the shear flow stress of the working material at the tool-chip interface and m is the constant shear friction factor which defined equal to 0.6. However, an iterative procedure should be carried out if the predicted values are not in good agreement with the experiments [27]. (2) There are several wear rate models to predict tool wear, as reviewed in the literature [9,12,13]. In this work, Usui’s wear rate model was used to calculate wear rate on the tool surfaces. According to Eq. (3), Usui model is established based on the normal stress ( ), sliding velocity (Vs), interface temperature (T), and two constant parameters A and B [28]. (
⁄ )
(3)
The variation in cutting parameters can be translated to the variation of model variables where changes in cutting speed results in sliding velocity alteration, and normal stress will be altered by changes in feed rate and depth of cut. An accurate value for constants of A and B, applicable to all cutting parameters, can be found if there is a strong relationship between cutting parameters and wear rate. This can be measured by a coefficient of determination that will be attained in a set of cutting experiments. For this purpose, experiments were designed based on the Central Composite Design (CCD) that allows fitting a polynomial curve on the values of wear rate. CCD is a fractional factorial design with five level. The design includes 2k factorial experiments, 2k axial points where k is the number of variables [29]. Table 1 shows the levels of factorial (+1 and -1) and axial point (-alpha, +alpha) where α=√ is selected to attain the rotatable condition in the design. Consequently, constants of A and B should be calibrated in an iteration comparison of the wear rate obtained in both simulation and experimental cutting process (Fig.3). The iteration process will be stopped when the constants of A and B result in the prediction error of less than ten percent in all cutting parameters. Table 1. The level of cutting parameters. Type Parameter
Unit
(Vc) (f) (ap)
Coated carbide
Ceramic
Levels
Levels
-
-1
0
+1
-
-1
0
+1
(m/min) (mm/rev)
32.96
50
75
100
117.04
65.91
100
150
200
234.09
0.1
0.14
0.2
(mm)
0.32
0.8
1.5
0.26
0.3
0.07
0.1
0.15
0.2
0.23
2.2
2.68
0.46
0.7
1.05
1.4
1.64
6
Fig. 3. Flow chart for determination of the Usui constant values.
Both Flank and crater wear have been investigated in this study. However, flank wear (VB) was selected as tool wear criterion. In order to assign the cutting length, some trial-and-error tests were conducted considering ISO-standard defined 0.6 mm for maximum flank wear (VB-max). Therefore, 70 and 50 mm cutting length were determined in turning by using coated carbide and ceramic tools, respectively. Fig. 4 shows a schematic tool wear profile. Tool wear profile commonly divided into three zones. Zone C is related to the tool nose. Zone N is a quarter of length b in the worn cutting edge. Zone B is the remained part that VB and VB-max are measured in this region. Furthermore, tool nose and notch wear bandwidth are named with VC and VN, respectively [30].
7
Fig. 4. Tool wear profile.
The experimental tests were conducted on a TC35 CNC machine with the transversal speed range from 1 to 3500 rpm. Also, an optical microscope and a scanning electron microscope (SEM) used to measure and validate tool wear estimation. The rake ( ) and clearance ( ) angles of both ceramic and coated carbide inserts mounted on tool holder were -6 and +6 degrees regardless of their chip breaker angles. Furthermore, the lead angle ( ) of coated carbide insert was 45°, while it was variable for ceramic insert due to its round shape. The turning tests were performed on a bar with 100 mm diameter. Equipment applied in experiments are illustrated in Fig. 5.
Fig. 5.Applied instruments.
3. Results The cutting parameters of experimental runs designed based on CCD and flank wear value of each run is listed in Table 2. The experimental cutting times (Eq. (4)) were used to compute the flank wear rates (dw/dt) for all cutting conditions in order to make results compatible with simulation outputs. 8
Table 2. Design of experiment and results. Type Run 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15
Coated carbide
Ceramic
Vc
f
ap
VB
Vc
f
ap
VB
(m/min)
(mm/rev)
(mm)
(mm)
(m/min)
(mm/rev)
(mm)
(mm)
50 100 50 100 50 100 50 100 32.96 117.04 75 75 75 75 75
0.14 0.14 0.26 0.26 0.14 0.14 0.26 0.26 0.2 0.2 0.1 0.3 0.2 0.2 0.2
0.8 0.8 0.8 0.8 2.2 2.2 2.2 2.2 1.5 1.5 1.5 1.5 0.32 2.68 1.5
0.125 0.271 0.163 0.292 0.245 0.725 0.187 0.604 0.063 0.500 0.388 0.312 0.030 0.542 0.343
100 200 100 200 100 200 100 200 65.91 234.09 150 150 150 150 150
0.1 0.1 0.2 0.2 0.1 0.1 0.2 0.2 0.15 0.15 0.07 0.23 0.15 0.15 0.15
0.7 0.7 0.7 0.7 1.4 1.4 1.4 1.4 1.05 1.05 1.05 1.05 0.46 1.64 1.05
0.396 0.334 0.478 0.291 0.459 0.167 0.627 0.479 0.564 0.847 0.589 0.311 0.231 0.459 0.312
(4) where t (s) is machining time, l (mm) is cutting length, n (rpm) is rotational speed of spindle, and f (mm/rev) is the feed rate. The coefficient of determination of experimental results for carbide and ceramic inserts is R2=0.954 and R2=0.913, respectively. This reveals that there is a strong correlation between independent and dependent variables where any changes in cutting parameters are always accompanied by changes in the flank wear rate. Therefore, it is evident that experimental results are reliable to be utilized in the iterative calibration method explained in section 2 in order to achieve the accurate constants of A and B as follow:
(
) ,
for TiAlN-coated carbide
(
) ,
for SiC-reinforced-Al2O3
Fig.6 compares the predicted and actual values of flank wear rate obtained in simulation and experiments, respectively. As can be seen, the predicted values are close to the experiments which verify the accuracy of A and B constants and the adequacy of finite element model developed.
9
Fig. 6. Comparison of predicted and experimental valuesof flank wear rate.
4. Discussions Based on the experimental and simulation results, depth of cut is the most influential factor on the flank wear compared to the other cutting parameters in turning of Inconel 625 superalloy. As can be seen in Fig. 7, the size of flank wear increased by an increment of depth of cut for both ceramic and coated carbide inserts, while the effect of feed rate increment on the flank wear for carbide tools is almost insignificant. This condition can be helpful to enhance productivity by utilizing higher range of feed rate which results in higher material removal rate. Also, the influence of cutting speed on the flank wear is similar to the effect of depth of cut when coated carbide tool was used. But this factor is not so effective on the flank wear of ceramic tool. The lower friction coefficient of ceramic tools compared to the carbide tools, their anti-adhesion characteristic, and chemical stability under 1200 0C made them suitable for applying in high cutting speeds [31].
Fig. 7. Effect of cutting parameters on the flank wear (VB).
The effect of depth of cut increment on the propagation of flank wear was investigated by considering generated temperature and stresses on tool faces shown in Figs. 8 and 9. In the machining operation, the energy of plastic deformation in the primary cutting zone is converted into the heat, and then with progress of cutting process, frictional heat is generated at the tool-chip interface (secondary deformed zone)[32]. A higher range of depth of cut corresponds with the temperature rise in tool faces due to the 10
higher plastic work requirement. Owing to the poor thermal conductivity of Inconel 625 and then low heat dissipation, temperature rises rapidly in the cutting zone and tool which finally results in faster tool wear on the tool surfaces (Figs. 8c and 9c). Furthermore, effective stress increased on the edge and faces of inserts by an increase of depth of cut which is completely distributed at the tool-chip contact zone (Figs. 8d and 9d). Accordingly, tool wear rate increased at the areas that more stress concentration is seen (illustrated in Figs. 8e and 9e). It should be noted that in order to compute stress-effective in simulations, the cutting tools were defined as elastic objects and the turning simulations were carried out under the same conditions described earlier. Besides, the severe chipping of cutting edge in the ceramic tool at the highest depth of cut can be seen in Fig.9g-3 which is due to the sudden engagement and impact imposed on that. Ceramic tools are not able to withstand against the impact of cutting forces due to their low fracture toughness that negative rake angles often applied to avoid chipping phenomenon [33].
11
Fig. 8.(a)The settings of coated carbide tool, (b) deformed chips, (c) the temperature distribution, (d) the stress distribution, (e) tool wear rate, (f) the experimental tool wear, ((g-1), (g-2) and (g-3)) FEA, optical and SEM images of flank wear (f=0.2 mm/rev, VC=75 m/min).
12
Fig. 9. (a) The settings of ceramic tool, (b) deformed chips, (c) the temperature distribution, (d) the stress distribution, (e) tool wear rate, (f) the experimental tool wear, ((g-1), (g-2) and (g-3)) FEA, optical and SEM images of flank wear (f=0.15 mm/rev, VC=150 m/min).
13
Crater wear is another important tool failure mode. However, as it was observed that flank wear decreases with an increase of feed rate, the increase of this parameter in order to enhance productivity regardless of crater wear concerns could be an unconsidered plan. Thus, the effect of feed rate variations on the crater wear is studied (Fig. 10). Usually, length of tool-chip contact ( ) is taken into account in order to analysis of crater wear [17, 34]. Undeformed chip thickness ( ) which is directly related to the feed rate (Eq. (5)), is the most important factor in the definition of tool-chip contact length regarding the Eq. (6) presented by Toropov and Ko. (5) (
(
where x, , and
)
)
(6)
are the cutting edge angle, tool rake angle and shear angle, respectively.
Accordingly, the increase of undeformed chip thickness as a result of feed rate increment caused further tool-chip contact length, seen in Fig. 10-b.This condition intensified wear depth at higher level of feed rate for both ceramic and coated carbide inserts. Fig. 10-c indicates generated tool crater wear in the same step at three levels of feed rate. Furthermore, Fig. 10-d shows optical microscope images of crater wear at tool rake face in the experiment. Eventually, the investigation of the effect of cutting parameters on the flank and crater wear by using FE simulation and experimental work revealed that low level of depth of cut and cutting speed with middle level of feed rate is a proper condition for coated carbide insert in turning of Inconel 625. Also the low depth of cut, high cutting speed together with the middle level of feed rate represents an optimum tool wear condition for the ceramic tool.
14
Fig. 10.(a)Turning operation at different feed rates (middle level of depth of cut and cutting speed), (b) deformed chip thickness, (c) predicted crater wear, and (d) experimental crater wear.
15
4. Concluding remarks In this study, a 3D finite element simulation method has been applied to predict tool wear in machining of Inconel 625 superalloy. Accordingly, Usui’s wear rate model was exerted and its constant parameters were calibrated by experiments carried out on PVD-TiAlN coated carbide and ceramic inserts. The main conclusions of this work are listed as follow:
Comparison of tool wear rate shows predicted and experimental values are in good agreement. However, it should be noted that constant A and B are valid in the range of cutting parameters selected in experiments, and might need to be calibrated again if a higher range of cutting parameters are desired. The simulation results indicate that tool wear rate is higher in the areas in which temperature and stresses are concentrated. The values of temperature and stresses increase by an increment in the level of depth of cut. The increase of depth of cut severely propagates the flank wear for both tools. The effect of cutting speed on the flank wear of ceramic tool is insignificant due to its heat resistance characteristic. However, the increase of cutting speed affects tool flank wear of coated carbide as the depth of cut does. Variation of feed rate is almost ineffective on the flank wear, while it intensifies crater wear due to the increase of undeformed chip thickness resulting the increase of tool-chip contact pressure. It can be concluded that low depth of cut and cutting speed with middle level of feed rate are appropriate cutting conditions for PVD-TiAlN coated carbide insert, while ceramic insert should be applied under condition of low depth of cut, middle level of feed rate and high cutting speed.
Acknowledgements The authors wish to thank the Multimedia University for funding this project under Mini-fund scheme (MMUI/160043) and the Production Lab of University of Kashan for their technical supports. Assistance of manufacturing lab-UTM is also greatly appreciated which provided the DEFORM license.
References [1] J. L. Cantero, J. Díaz-Álvarez, M.H. Miguélez, N.C.Marín, Analysis of tool wear patterns in finishing turning of Inconel 718, Wear, 297(1)(2013) 885-894. [2] A. Bhatt, H. Attia, R. Vargas, V. Thomson, Wear mechanisms of WC coated and uncoated tools in finish turning of Inconel 718. Tribology International, 43(5) (2010) 1113-1121. [3] C. Courbon, F. Pusavec, F. Dumont, J. Rech, and J. Kopac. Tribological behaviour of Ti6Al4V and Inconel718 under dry and cryogenic conditions—Application to the context of machining with carbide tools. Tribology International 66 (2013): 72-82. [4] A. Devillez, G. Le Coz, S. Dominiak, D. Dudzinski, Dry machining of Inconel 718, workpiece surface integrity. Journal of Materials Processing Technology 211(10) (2011) 1590-1598.
16
[5] Z. Hao, D. Gao, Y. Fan, R. Han, New observations on tool wear mechanism in dry machining Inconel718. International Journal of Machine Tools and Manufacture, 51(12) (2011) 973-979. [6] Zhuang, K., Zhu, D., Zhang, X., Ding, H. Notch wear prediction model in turning of Inconel 718 with ceramic tools considering the influence of work hardened layer, Wear, 313(1) (2014) 63-74. [7]T. Kagnaya, L. Lambert, M. Lazard, C.Boher, T. Cutard, Investigation and FEA-based simulation of tool wear geometry and metal oxide effect on cutting process variables, Simulation Modelling Practice and Theory, 42 (2014) 84-97. [8] R.K. Yadav, K. Abhishek, S.S. Mahapatra, A simulation approach for estimating flank wear and material removal rate in turning of Inconel 718, Simulation Modelling Practice and Theory, 52 (2015) 1-14. [9] E. Usui, T. Shirakashi, T. Kitagawa, Analytical prediction of cutting tool wear, Wear, 100(1) (1984) 129-151. [10] Y.C. Yen, J. Söhner, B. Lilly, T. Altan, Estimation of tool wear in orthogonal cutting using the finite element analysis, J. Mater. Proc. Technol. 146(1) (2004) 82-91. [11] J. Lorentzon, N. Järvstråt, Modelling tool wear in cemented-carbide machining alloy 718, Int. J. Mach. Tools Manuf. 48(10) (2008) 1072-1080. [12] H. Takeyama, R. Murata, Basic investigation of tool wear, J. Manuf. Sci. Eng. 85(1) (1963) 33-37. [13] M. Binder, F. Klocke, D. Lung, Tool wear simulation of complex shaped coated cutting tools, Wear, 330 (2015) 600607. [14] G. Shi, X. Deng, C. Shet, A finite element study of the effect of friction in orthogonal metal cutting, Finite Elements in Analysis and Design, 38(9) (2002) 863-883. [15] A. Attanasio, E. Ceretti, A. Fiorentino, C. Cappellini, C. Giardini, Investigation and FEM-based simulation of tool wear in turning operations with uncoated carbide tools, Wear, 269(5) (2010) 344-350. [16] A. Kurt, B. Yalçin, N. Yilmaz, The cutting tool stresses in finish turning of hardened steel with mixed ceramic tool, Int. J. Adv. Manuf. Technol. (2015) 1-11. [17] A. Toropov, S.L. Ko, Prediction of tool-chip contact length using a new slip-line solution for orthogonal cutting, Int. J. Mach. Tools Manuf. 43(12) (2003) 1209-1215. [18] T. Özel, Computational modelling of 3D turning: Influence of edge micro-geometry on forces, stresses, friction and tool wear in PcBN tooling, J. Mater. Proc. Technol. 209(11) (2009) 5167-5177. [19] B. Haddag, H. Makich, M. Nouari, and J. Dhers. Tribological behaviour and tool wear analyses in rough turning of large-scale parts of nuclear power plants using grooved coated insert. Tribology International 80 (2014) 58-70. [20] M. Hokka, D. Gomon, A. Shrot, T. Leemet, M. Bäker, V.T. Kuokkala, Dynamic Behavior and High Speed Machining of Ti-6246 and Alloy 625 Super alloys: Experimental and Modeling Approaches, Experimental Mechanics, 54(2) (2014) 199-210. [21] Technical Committee, Coro Guide. In Coromant, S., Ed. Sandvik Coromant, 2014.
17
[22] DEFORM-3D [Computer software]. (2012). http:// www.deform.com [23] A. Vesel, A. Drenik, K. Elersic, M. Mozetic, J. Kovac, T. Gyergyek, J. Stockel, J. Varju, R. Panek, M. Balat-Pichelin, Oxidation of Inconel 625 super alloy upon treatment with oxygen or hydrogen plasma at high temperature, Applied Surface Science, 305 (2014) 674-682. [24] http://www.specialmetals.com/inconel-alloy-625 [25] http://www.jacquet.biz/JACQUET/USA/files/JCQusa-alloy-625.pdf [26] T. Özel, The influence of friction models on finite element simulations of machining, Int. J. Mach. Tools Manuf. 46(5) (2006) 518-530. [27] M. Lotfi, A.A. Farid, H. Soleimanimehr, The effect of chip breaker geometry on chip shape, bending moment, and cutting force: FE analysis and experimental study, Int. J. Adv. Manuf. Technol. 78(5-8) (2015) 917-925. [28] L.J.Xie, J. Schmidt, C. Schmidt, F. Biesinger, 2D FEM estimate of tool wear in turning operation, Wear, 258(10) (2005) 1479-1490. [29] A. A. Farid, S. Sharif, S. Alizadeh Ashrafi, M. H. Idris, Statistical Analysis, Modeling, and Optimization of Thrust Force and Surface Roughness in High-Speed Drilling of Al–Si Alloy. Proc. Inst. Mech. Eng. Part B-J. Eng. Manuf. 227 (6) (2013) 808-20. [30] B. Li, A review of tool wear estimation using theoretical analysis and numerical simulation technologies. International Journal of Refractory Metals and Hard Materials 35 (2012) 143-151. [31] A. Altin, M. Nalbant, A. Taskesen, The effects of cutting speed on tool wear and tool life when machining Inconel 718 with ceramic tools, Materials & design, 28(9) (2007) 2518-2522. [32] E. Muthu, K. Senthamarai, S. Jayabal, Finite element simulation in machining of Inconel 718 nickel based super alloy, Int. J of Adv. Eng. Applic,1(3) (2012) 22-27. [33]http://www.tech.plym.ac.uk/sme/mfmt201/cuttingtoolmats.htm [34] M. Lotfi, A.A. Farid, H. Soleimanimehr, A new hybrid model based on the radius ratio for prediction of effective cutting limit of chip breakers, Proc. IMechE Part B: J. of Eng. Manuf. (2015), 0954405415586549.
18
Table captions: Table 1. The level of cutting parameters. Table 2. Design of experiment and results.
Figure captions: Fig. 1. a) The cutting inserts, b) the modeled inserts, c) the meshed inserts and workpieces, and d) the boundary condition. Fig. 2. Inconel 625 material properties and its Johnson-Cook constant parameters. Fig. 3. Flow chart for determination of the Usui constant values. Fig. 4. Tool wear profile. Fig. 5.Applied instruments. Fig. 6. Comparison of predicted and experimental values of flank wear rate. Fig. 7. Effect of cutting parameters on the flank wear (VB). Fig. 8. (a) The settings of coated carbide tool, (b) deformed chips, (c) the temperature distribution, (d) the stress distribution, (e) tool wear rate, (f) the experimental tool wear, ((g-1), (g-2) and (g-3)) FEA, optical and SEM images of flank wear (f=0.2 mm/rev, VC=75 m/min). Fig. 9. (a) The settings of ceramic tool, (b) deformed chips, (c) the temperature distribution, (d) the stress distribution, (e) tool wear rate, (f) the experimental tool wear, ((g-1), (g-2) and (g-3)) FEA, optical and SEM images of flank wear (f=0.15 mm/rev, VC=150 m/min). Fig. 10. (a)Turning operation at different feed rates (middle level of depth of cut and cutting speed), (b) deformed chip thickness, (c) predicted crater wear, and (d) experimental crater wear.
Highlights:
A 3D finite element simulation of tool wear in turning of Inconel 625 is carried out. PVD-TiAlN coated carbide and ceramic inserts are used. The effect of cutting parameters on the temperature, stress, and tool wear are studied. The increment in the level of depth of cut intensifies the flank wear propagation. The increase of tool-chip contact length increases crater wear of cutting tools.
19