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An investigation on wear mechanism of high-speed turning of free-cutting steel AISI 1215


Int J Adv Manuf Technol DOI 10.1007/s00170-012-4502-8

ORIGINAL ARTICLE

An investigation on wear mechanism of high-speed turning of free-cutting steel AISI 1215 using uncoated and multi-layer coated tools
Jinyang Xu & Zhiqiang Liu & Guoqiang Guo & Ming Chen

Received: 25 February 2012 / Accepted: 5 September 2012 # Springer-Verlag London Limited 2012

Abstract AISI 1215 is a new kind of green and nontoxic free-cutting steel with minimum environmental pollution and excellent machinability, which receives wide promotion, investigation, and application in manufacturing industries. In machining of AISI 1215 steel, tool wear has a close relation with the presence of manganese sulfide lubricant zone formed on the tool surface. In this work, with the aid of cutting temperature and tool Von Mises stress simulations, tool wear analysis on the uncoated and multi-layer (Al2O3/TiCN) coated carbide tools was performed in high-speed turning operation. Wear pattern and wear mechanisms were studied through the experimental results. The main findings showed that the uncoated tool suffered high cutting temperature and severe tool wear and was not conducive to form a manganese sulfide lubricant zone in the turning operation. In contrast, the multi-layer coated tool could form a manganese sulfide lubricant zone on the chip–tool contact area. The beneficial roles of the manganese sulfide lubricant zone formed on the coated tool surface can be summarized as lubrication and diffusion blocking. The main wear mechanisms of the uncoated tool were crater wear, oxidation wear, adhesive wear, and abrasive wear, whereas for the multi-layer coated tool, they were crater wear, adhesive wear, and abrasive wear. Keywords High-speed turning . Free-cutting steel . Multi-layer coated tool . Cutting temperature . Tool wear

J. Xu : Z. Liu : G. Guo : M. Chen (*) School of Mechanical Engineering, Shanghai Jiao Tong University, Shanghai 200240, People’s Republic of China e-mail: mchen@sjtu.edu.cn

Nomenclature vc Cutting speed (in meters per minute) f Feed rate (in millimeters per revolution) ap Depth of cut (in millimeters) g(εp) Strain hardening : *?"? Strain rate sensitivity Θ(T) Thermal softening ΔT Temperature increment f1 Work–heat convection factor f2 Conversion efficiency factor σ Effective stress (in megapascals) @"pl Effective plastic strain increment ρ Material density (in grams per cubic centimeter) Cp Specific heat σ0 Initial yield stress of material (in megapascals) p ε Plastic strain "p Reference plastic strain 0 "p Cutoff strain cut n Strain hardening exponent ? " Strain rate ? "0 Reference plastic strain rate ? "t Strain rate where the transition between low and high strain rate sensitivity occurs m1 Low strain rate sensitivity coefficient m2 High strain rate sensitivity coefficient c0~c5 Constants for polynomial fit T Cutting temperature (in degrees Celsius) Tcut Linear cutoff temperature (in degrees Celsius) Tmelt Melting temperature (in degrees Celsius) VB Average flank wear land width (in millimeters) hs Surface wear rate (in micrometers/103 sm2) hr Current radial wear (in micrometers) hr–i Initial radial wear (in micrometers) li Initial length of tool path (in millimeters) lt Total length of tool path (in millimeters) tun Uncut chip thickness

Int J Adv Manuf Technol

dcw wa Wa t c

Depth of cold working Abrasive wear rate Mass loss due to abrasion of the tool Cutting time Constant

Abbreviations SEM Scanning electron microscope EDS Energy dispersive X-ray spectroscopy XPS X-ray photoelectron spectroscopy 1 Introduction The excellent machinability of free-cutting steels has awarded the materials such wide application in machinery, energy, metallurgy, aerospace, and other fields since the steels are first developed in the USA in the 1920s [1]. The so-called free-cutting steel refers to a specific kind of steel by adding some individual or composite elements (such as sulfur, phosphorus, lead, selenium, etc.) to improve its cutting performance and to satisfy the development demands of automatic processing [2, 3]. The machinability of free-cutting steel is mainly related to the beneficial role of inclusions in its matrix. These advantageous inclusions can form stress concentration source, suppress crack growth, resist friction, and play lubricant effect in the machining process. As a result, it leads to the improvement of machinability with much lower tool flank wear and crater wear. Based on the excellent machinability, various freecutting steels have been researched and developed such as the leaded, sulfured, selenium, and tellurium-treated freecutting steels [4, 5]. However, most of the traditional freecutting steels are often toxic and contain heavy metals, so the smelting and hot processes usually cause serious environmental pollution and great harm to human health. Therefore, developing environment-friendly freecutting steel as a replacement for toxic free-cutting steel is of strategic significance. AISI 1215 is a newly developed sulfured free-cutting steel which contains no heavy metal elements, so it will not cause environmental pollution or health hazard in the smelting, hot, and machining process. In addition, AISI 1215 is also supposed to have excellent cutting performance owing to the sulfide inclusions in its matrix. These sulfide inclusions distribute in the steel in the form of manganese sulfide [6], and they have a dominant effect in reducing tool wear and improving the machinability of the steel. In metal cutting, tool wear is an important factor directly affecting the surface quality of the machined parts and an important parameter in evaluating the performance of the cutting tools [7, 8].

To better improve tool life, various machining researches had been conducted to investigate the wear mechanism of free-cutting steels. Jiang et al. [9] studied the effects of sulfide inclusions on the machinability of resulfurized freecutting steel. It had been observed that tool flank wear was affected markedly by the behavior of sulfide inclusions in the third shear zone. The microvoids initiated by sulfide inclusions with higher shape factor greatly reduced the flank wear of the tool. Fang and Zhang [10] found that the adhering layer formed on both tool rake and flank faces had a protective effect on reducing tool wear and thus prolonged the tool life when machining Ca–S free-cutting stainless steel. Li and Wu [5] used cemented carbide tools to investigate the effect of free-cutting additives on machining characteristics. Results showed that the machinability of free-cutting steel was visibly improved by adding freecutting additives, such as S, Mn, Bi, etc. Lane et al. [11] studied the tool wear of AISI 1215 using diamond tool. They found that wear rate was significantly higher for the AISI 1215 than for Al6061 ascribed to the higher rates of chemical wear near the high-temperature cutting zone. Xu et al. [12] studied the wear mechanism of free-cutting AISI 12L14 using chemical vapor deposition (CVD)-coated tools, and found that the MnS inclusions in the steel matrix formed a lubricant zone on the tool rake face, which greatly reduced the tool wear. Hashimura et al. [13] investigated the effect of the distribution and size of MnS grains in steels on the machinability of freecutting steels and compared to conventional steels. It was found that the fine MnS grains which kept supplied on the deformation zone around the cutting edge played the lubricant effect in improving the machinability of the steel and reducing tool wear. However, the previous researches presented few studies on the effects of different tools (especially the uncoated tool and coated tool) on the formation of the manganese sulfide lubricant zone when machining sulfured free-cutting steels. Based on this, the paper focused on the fundamental wear mechanism obtained in high-speed turning of AISI 1215 by using uncoated tool and CVD-coated carbide tool with the same geometry. The coating of the CVD-coated tool refers to typical multi-layer coating materials which are widely employed to extend the tool life and cutting performance due to their advanced wear resistance and superior perforTable 1 Chemical composition of free-cutting steel AISI 1215 (in weight percent) Elements Contents C 0.08 Mn 1.05 S 0.28 P 0.055 Si 0.03 Fe Balance

Int J Adv Manuf Technol Fig. 1 Microstructure of freecutting steel AISI 1215 (×200). a Cross section perpendicular to the rolling direction. b Longitudinal section along the rolling direction

(a)
Pearlite MnS inclusions Carbon inclusions Pearlite

(b)

Ferrite

mance under corrosive or high-temperature conditions [14–17]. Finite element method (FEM) simulations of cutting temperature and tool Von Mises stress were applied to assist the analysis of tool wear. Both the tool wear rate and wear mechanism were investigated, and their results were discussed correspondingly.

2 Experimental procedures 2.1 Testing material The workpiece used in the turning experiments was an AISI 1215 bar with the size of Ф70×220 mm. The chemical composition (in weight percent) of the steel is given in Table 1. As a typical kind of hypo-eutectoid steels, the microstructures of AISI 1215 are ferrite, pearlite, manganese sulfide (MnS) inclusions, and carbon inclusions as shown in Fig. 1. Ferrite is the primary phase of AISI 1215 steel. Pearlite distributes
Table 2 Tool geometric parameters of used inserts

around the boundary of a few ferrites, and small inclusions are observed in the internal ferrite. As can be seen in Fig. 1b, several spindle-shaped manganese sulfide inclusions are clearly evident around the ferrite boundary from the longitudinal section along the rolling direction. In addition, pearlite distributing along the rolling direction is more concentrated, which forms strips of pearlite in the steel matrix. These strips of pearlite will be the reinforced phase which makes the matrix strength in the rolling direction much higher than that perpendicular to the rolling direction. Therefore, for easy cutting, the feed direction was set perpendicular to the rolling direction in the turning operation. 2.2 Turning tests Turning trails were carried out on an INDEX G200 CNC Centre Lathe. The travel range of the machine tool is defined as follows: X—105 mm, Z—400 mm, four axes (X-, Z-, C-, and Y-axis); spindle speed range, 20–6,000 rpm; positioning

Size (mm) Tool model Tool type Insert photo Coating morphology iC s
r

l

CNMG120408

Uncoated carbide insert

12.7

4.7625

0.8

12

CNMG120408

Multi-layer coated insert

12.7

4.7625

0.8

12

Int J Adv Manuf Technol

Fig. 2 Schematic diagram of the tool temperature measurement setup in the turning operation

accuracy, 0.01 mm. Both uncoated insert and CVD-coated (Al2O3/TiCN) carbide insert were adopted in the following experiments which had the same rake angle of ?6°, clearance angle of 0°, and inclination angle of 0°. Tool geometric parameters of the used inserts are presented in Table 2. It should be highlighted that the thick and low-stress Al2O3 coating layer provides maximum thermal and chemical protection while the MTCVD TiCN coating supplies the best possible mechanical wear resistance and adherence to the substrate. In addition, the gradient substrate also gives optimized deformation resistance and toughness for turning operation. The range of cutting parameters for tool wear rate analysis was selected as follows: cutting speed (vc) of 100 and 500 m/min, feed rate (f) of 0.1, 0.2, 0.3, and 0.4 mm/rev, and depth of cut (ap) maintaining a constant of 1 mm. For wear mechanism analysis, cutting parameters (vc 0500 m/min, f00.2 mm/rev, ap 01 mm) were selected for the usage. Average flank wear land width, VB00.4 mm, was selected as the wear criterion of the cutting tools, and cutting fluid was not used in the operation. Tool edge morphology was recorded by a Nikon toolmaker's microscope with 30 magnification and 1 μm resolution. During the turning process, tool flank wear was measured several times by self-developed software according to ISO 3685 standard (1993) with VB used for sintered carbide tools. Tool temperature distribution was measured by using the CCD sensor-based near-infrared (380–780 nm) imaging technique covering a temperature range of 100– 1,000 °C, and a suitable calibration method enables the measurement of tool temperature with reasonably good
Fig. 3 Schematic diagram of the composition and workflow of the used software AdvantEdge?

accuracy (±5 °C) and better spatial resolution (4.5 μm). The schematic diagram of tool temperature measurement setup in turning operation is presented in Fig. 2. In addition, scanning electron microscopy (SEM) and energy dispersive spectrometer (EDS) analyses were all performed on the tool surfaces to understand the wear behavior by using a JSM7600F SEM manufactured by JEOL company with the following specifications: SEI resolution, 1.5 nm (1 kV) in GB mode, 1.0 nm (15 kV); magnification, ×25 to ×1,000,000; and accelerating voltage, 0.1 to 30 kV. After that, X-ray photoelectron spectroscopy (XPS) of the chemical composition on the worn tool surfaces was conducted by using the Thermo Scientific K-Alpha X-ray photoelectron spectrometer. The K-Alpha analyzer is a fully integrated X-ray photoelectron spectrometer using a monochromatic Al Kα source (1,486.6 eV) with the following technical indexes: angleresolved XPS angle ≤1 °C, adjustable beam from 30 to 400 μm, and energy resolution ≤0.5 eV. FEM software named AdvantEdge? was applied to simulate the cutting temperature and tool Von Mises stress in the turning operation. 2.3 FEM model The heat generation and Von Mises stress distribution during metal cutting process strongly influence tool life, and FEM simulation is a good method to observe them [18, 19]. Based on this, FEM software named AdvantEdge? has been introduced to model the current dry turning of AISI 1215 steel. This software uses a Lagrangian method to account for dynamic effects, heat conduction, and full thermo-mechanical coupling [20]. Figure 3 shows the composition and workflow of the used software. It is notable that the AdvantEdge? software package has three main components. The first is simulation setup interface which allows users to set up the entire simulation, including defining material condition, tool geometries, and machining parameters. When the simulation setup is finished, the AdvantEdge Engine performs all the calculations from the setup inputs. Finally, the Results Viewer enables users to extract the necessary simulation results, including cutting forces, tool temperature, Von Mises stress, etc. The FE simulation models were built for high-speed turning of AISI 1215 steel with uncoated and multi-layer coated cutting tools. The differences of the machining process between uncoated tool and coated tool were obtained through the observation of the simulation results. Dynamic
Cutting forces

Material condition Tool temperature Tool geometries Machining parameters Von Mises stress Steady state analysis

Int J Adv Manuf Technol Fig. 4 Detailed simulation setup for AISI 1215 turning modeling
Work material: AISI 1215 Tool material: carbide P Tool geometry: -6o rake angle, 0o clearance angle and -6o inclination angle Cutting parameters: vc=500m/min, f=0.2mm/r, ap=1mm Cutting condition: dry turning; frictional coefficient equal to 0.4

Inp

ut
Output Cutting temperature AdvantEdgeTM FEM software

Work material: AISI 1215 Tool material: carbide P; Al2O3 (6 m) and TiCN (6 m ) Tool geometry: -6o rake angle, 0o clearance angle and -6o inclination angle Cutting parameters: vc=500m/min, f=0.2mm/r, ap=1mm Cutting condition: dry turning; frictional coefficient equal to 0.3

Output

u In p

t

Von Mises stress

effects, heat conduction, and full thermo-mechanical coupling were calculated with Lagrangian method in FEM simulation.

In the current FE model, the power law is used to model work material by Advantage?, which can be expressed as: ?  ?  σ "p ; "; T ? g ?"p ? ? * " ? D?T ? ?1? ? where g(εp) is strain hardening, * " is strain rate sensitivity, and Θ(T ) is thermal softening. The constant Coulomb frictional coefficient of 0.4 was set for the uncoated tool [18]. Since (Ti, Al) N coatings could reduce heat radiation and friction, the coefficient of tool–chip friction of multi-layer coated tool was set as 0.3. The majority of the heat generated in the metal cutting process comes from plastic deformation and tool–chip interface friction. The temperature increment associated with the heat generation can be expressed as follows [21, 22]: $T ? f1 ? f2 ? σ ? @"pl ρCp ?2?

(a)
2.5

Y (mm) 1.5 2.0

Heat dissipation into the rake face Elastic-plastic deformation induced heat

1.0

8.0

8.5

9.0

9.5 X (mm)

10.0

10.5

(b)
Localized peak temperature zone

where ΔT is temperature increment, f1 is the work–heat convection factor, and f2 is the conversion efficiency factor. Both f1 and f2 are taken as 0.9 as most of the workpiece deformation is converted to thermal energy. σ is the effective stress, @"pl is
800 700
Simulated-uncoated tool Simulated-multi-layer coated tool Measured-uncoated tool Measured-multi-layer coated tool

1.5

Cutting temperature ( C)
7.0 7.5 X (mm) 8.0 8.5

600 500 400 300 200 100 0 0.0 0.2

0.5

Y (mm) 1.0

Distance from the starting point of the ray (mm)
Fig. 6 Comparison of simulated and measured tool temperature distribution of the tool bodies

0.4

0.6

0.8

Fig. 5 Temperature distribution of used cutting tools at the steady cutting stage. a Uncoated tool. b Multi-layer coated tool

Int J Adv Manuf Technol

(a)
2.6

Von Mises Stress (MPa)

the effective plastic strain increment, and ρ and Cp are material density and specific heat, respectively. The strain hardening function g(εp) is defined as:  1 "p n p g?" ? ? σ0 1 ? p ; if " p < "cut ; "0
p

2.4

?3?

Y (mm)

2.2


Localized peak stress

g ?" ? ? σ0
p

2.0

"p 1 ? cut p "0

1 n

p ; if " p ! "cut

9.0

9.2

9.4 9.6 X (mm)

9.8

10.0

10.2

p where σ0 is the initial yield stress, εp is the plastic strain, "0 p is the reference plastic strain, "cut is the cutoff strain, and n is the strain hardening exponent.   ? The rate sensitivity function " is expressed as:

1.8

(b)
2.2

Von Mises Stress (MPa)

0 1 m1 1 ? ? "A ? ; if " * " ? @1 ? ? p "0

";
t

?

?4?

Y (mm) 2.0

0 1 m1 ! m1 ? m1 1 ? ? ? "A "t 1 2 ? ? @1 ? ? * " ? 1? ? ; if " > " t p "0 "0
Localized peak stress

where " is the strain rate, " 0 is the reference plastic strain rate, "t is the strain rate where the transition between low and high strain rate sensitivity occurs, m1 is the low strain rate sensitivity coefficient, and m2 is the high strain rate sensitivity coefficient. The thermal softening function Θ (T) is defined as:
D?T ? ? c0 ? c1 T ? c2 T 2 ? c3 T 3 ? c4 T 4 ? c5 T 5 ; if T < Tcut ;   T ? Tcut ; if T ! Tcut D?T ? ? D?Tcut ? 1 ? Tmelt ? Tcut
?

?

?

1.8

7.0

7.2

7.4 X (mm)

7.6

7.8

Fig. 7 Tool Von Mises stress distribution of used cutting tools at the steady cutting stage. a Uncoated tool. b Multi-layer coated tool

0.6
Flank wear, VB (mm)

?5?
Multi-layer coated tool Uncoated tool

0.5 0.4 0.3

VB=0.4 mm

l=2056m

where c0~c5 are coefficients for the polynomial fit, T is the cutting temperature, Tcut is the linear cutoff temperature, and Tmelt is the melting temperature. In the finite element model, the coating layers (Al2O3 and TiCN) were represented by separate thin layers of their

(a)
0.2 0.1 0.0 0 6 12 18 24 30 36 2 Cutting length, l (x10 m) 42 48

(b)
Adhesive chip

Crater wear
Fig. 9 Tool morphology of uncoated tool, l0465 m (×30). a Rake face. b Flank face

Fig. 8 Tool flank wear VB curve versus cutting length (l)

Int J Adv Manuf Technol

(a)

(b)

Wear land expanding
hs (?m/10 sm )
2

15 12 9 6
vc=100m/min
3

Multi-layer coated tool vc=500m/min Uncoated tool

vc=500m/min

Crater wear exacerbating
Fig. 10 Tool morphology of uncoated tool, l02,056 m (×30). a Rake face. b Flank face

3 0.1

vc=100m/min

0.2

respective thickness (Al2O3 (6 μm) and TiCN (6 μm)) on the top of the carbide substrate. The detailed simulation setup for AISI 1215 turning modeling is presented in Fig. 4. The thermo-mechanical coupling works through the selected finite elements. These elements are naturally thermo-mechanical coupled so the strain generates as well as the heat when these elements deform.

0.3 f (mm/rev)

0.4

Fig. 12 Influence of feed rate (f) and cutting speed (vc) on tool wear rate of the used cutting tools

3 Experimental results and discussion 3.1 Analysis on cutting temperature and tool Von Mises stress It is well known that the tool wear mechanism has close relation with the cutting temperature and tool Von Mises stress. The high cutting temperature generated on the tool surface often aggravates the crater wear, adhesive wear, and oxidation wear. Von Mises stress is an equivalent stress according to the fourth strength theory and an important indicator to measure the stress level of cutting tools in metal cutting. The level of tool Von Mises stress usually has close relation with the tool failures such as tool breakage, micro chipping, and coating peeling. Therefore, investigation on cutting tool temperature and tool Von Mises stress plays a significant role in wear mechanism analysis. Figure 5 presents the results about the temperature distribution of uncoated and multi-layer coated tools when each tool machines to the steady cutting stage. In order to quantitatively compare and analyze the cutting temperature of

the used cutters, two rays were made from the tool nose to the cutter body of each cutting tool, respectively. Figure 6 shows the comparison of simulated and measured tool temperature distribution curves of the tool bodies. It is notable that the simulated values are in good agreement with the measured ones, which indicates that the simulation results have sufficient reliability and credibility. It is observed that the peak temperature zones are located on the tool–chip interface and the third shear zone for the uncoated tool as seen in Fig. 5a. It is well known that tool wear is largely dependent on the cutting temperature, which, in turn, depends on how much heat is conducted into the tool during a machining operation [23]. This high temperature in the cutting tool may be caused by sustaining friction, thermal conduction, and elastic–plastic deformation. When the uncoated tool machines to the steady cutting stage, a large amount of cutting heat will dissipate into the rake face, which results in high cutting temperature concentrated at a narrow region adjacent to the cutting edge, where the temperature rises up to about 770 °C. This phenomenon demonstrates that tool rake face also plays the role of dissipating cutting heat. Consequently, as the cutting tool and workpiece keep continuous contact; more and more heat will transfer to the cutting tool, which results in a higher temperature increase. Ascribed to the presence of localized peak

(a)

(b)

Small wear land

Minor crater wear
Fig. 11 Tool morphology of multi-layer coated tool, l01,962 m (×30). a Rake face. b Flank face Fig. 13 Schematic diagram of the depth of cold working of the transient surface

Int J Adv Manuf Technol
Fe O

B
Fe

Fe O

A

Fe C C W W Fe W W W W W Si

A

Fe

B

C
C

W

D

D

O Fe C

Fe

C

Co W Co W W Co W W W Fe

Fig. 14 SEM and EDS analyses of the tool rake face for the uncoated tool at the cutting length of 1,314 m

temperature zone on the tool rake face, the high contact temperatures on the interface can cause significant crater wear in the form of highly localized plastic deformation. In addition, the high temperature will also dramatically undermine and weaken the mechanical properties of the tool material, making the cutter more vulnerable to adhesive wear and oxidation wear. In contrast, the cutting condition of the multi-layer coated tool is better because most of the generated cutting heat is largely dissipated by the chip material instead of the tool rake face. This is mainly because the thermal conductivity coefficients are low on coating layer Al2O3 resulting in the heat flow to the tool to be low [24, 25]. Consequently, the temperature of the coated tool is averagely 80 °C lower than that of the uncoated tool (as depicted in Fig. 6). The reason for this can be attributed to the protective role of the multi-layer coating material in reducing the friction coefficient of the tool–chip interface and preventing the tool from absorbing heat [26]. Summarily, the multi-layer coated tool can have the dominant effect in reducing the cutting temperature in contrast with uncoated tool. Figure 7 shows the simulated Von Mises stress distribution of the used inserts when machining to the steady cutting stage. Two remarkable characteristics of Von Mises stress distributions can be observed among the used cutting tools.
Table 3 EDS results of the regions shown in Fig. 14 (in weight percent) Fe Region Region Region Region A B C D 46.74 55.20 51.02 – O 30.13 27.54 22.72 – C 20.71 11.89 26.26 9.87 Si 2.42 – – – W – 5.37 – 82.62 Co – – – 7.51

One is the distribution law of the stress; the other is the location of the peak Von Mises stress. It is observed that the Von Mises stress distribution of the coated tool is uniform whereas that of the uncoated tool is uneven. In contrast, the peak Von Mises stress of the uncoated tool is localized on both tool rake face and tool nose. This phenomenon demonstrates the cutter has suffered severe friction and pressure on the tool–chip interface and tool nose during the highspeed cutting process. Furthermore, the peak Von Mises stress of the coated tool is mainly concentrated on the tool nose, and its value reaches as high as 6,920 MPa which is about 126 % larger than that of the uncoated tool. This indicates that micro chipping and coat peeling have a great probability of occurring during turning operation of the tool nose. It should be emphasized that the localized peak Von Mises stress around the tool nose may favor the formation of

Fig. 15 Schematic diagram of the diffusion model without the MnS lubricant zone

Int J Adv Manuf Technol
Fe O

G
Fe

Fe O

H

C

W

G
Fe W W W W W

H
C

Fe

W

Crack

F E

I

Si

Fe

Fe O

F
Fe

Fe O C W

E

O Fe Fe

I

C

W W Fe W W W W

Fe C W Fe W W W W W Fe

W

Fig. 16 SEM and EDS analyses of the tool rake face for the uncoated tool at the cutting length of 2,692 m

adhering inclusions around the cutting edge because their formation usually requires the condition of high pressure. Overall, the multi-layer coated tool suffers higher Von Mises stress in contrast with the uncoated cutting tool. This abnormal phenomenon shows that the addition of a coating may increase the tool Von Mises stress. In order to reveal this phenomenon, the cutting edge radius should be taken into consideration. In fact, the cutting edge radius of the multi-layer coated tool is 20 μm before the coating process, and the total thickness of the multi-layer coating is nearly 12 μm. As a result, the cutting edge radius reaches 32 μm, which is larger than that of the uncoated tool. The enlarged cutting edge radius will increase the resistance force of separating the chip from the workpiece and thus generates high tool Von Mises stress. 3.2 Wear morphology of uncoated and multi-layer coated tools In these experiments, tool wear was measured as the average flank wear land width, VB, as a function of cutting length (l) as shown in Fig. 8. Figures 9, 10, and 11 show the wear
Table 4 EDS results of the regions shown in Fig. 16 (in weight percent) Fe Region Region Region Region Region E F G H I 38.98 43.89 44.53 46.74 51.02 O 29.37 31.62 26.30 30.13 22.72 C 21.17 15.51 18.74 20.71 26.26 Si – – – 2.42 – W 10.47 8.98 10.43 – –

surface morphology of uncoated and multi-layer coated tools, respectively. As shown in Fig. 8, the used cutting tools have undergone the initial and normal wear stages in the entire turning process. In order to quantitatively analyze the functional relation between VB and l, polynomial function is employed to fit the wear curves of the used inserts. The fitting functions for uncoated and multi-layer coated tools are given by the following equations: Uncoated tool : VB?in micrometers? ? 93:97 ? 10?2 ? 94:43l ? 85:5l 2 ? 3:312 ? 10?1 l 3 ? 4:53 ? 10?3 l 4 Multi ? layer coated tool : VB?in micrometers? ? 3:16 ? 48:24l ? 4:8l 2 ? 0:23l 3 ? 4:55 ? 10?3 l 4 ?7? ? 3:44 ? 10?5 l 5 The above equations can be used to quantitatively calculate the VB value under any cutting length (l) in the range of 0–2,692 and 0–4,758 m, respectively. It should be noted that both Eqs. (6) and (7) have at least 98 % confidence through analysis of variance which is not presented here. During the initial wear stage, tool wear rates of the used cutters are both rapid. This phenomenon can be explained by the fact that the surface roughness value of antithesis surfaces is high, and the actual contact area suffers powerful pressure and severe wear in their early running, which thereby accelerates the wear rate correspondingly. As seen in Fig. 9, the uncoated tool is observed to suffer severe crater wear and flank wear ?6?

Int J Adv Manuf Technol

(a)

machined surface and can be represented as follows:
hs ?
Point 1

 ? dhr ?hr ? hr?i ?100 ? ? in micrometers 103 sm2 dS ?lt ? li ?f

?8?

where hr, hr?i, and li are the current radial wear, initial radial

15000

14000 12000

(b)
12000

(a)
Fe2p3/2 in Fe2O3

Intensity (c/s)

10000 8000 6000 4000

Intensity (c/s)

Fe2p3/2 in Fe3O4

9000

6000

730

725

720

715

710

705

540

535

530

525
15000

Binding energy (eV)

Binding energy (eV)
Fig. 17 XPS spectrum of point 1 on tool rake face at the cutting length of 2,692 m. a Position of point 1. b XPS spectrum of oxygen element

(b)
14000

Co2p3/2 in Co2O3 Co 2p3/2 in Co3O4

even in the initial wear stage, i.e., l0465 m. Several adhesive chips are found on the tool rake face. These bonding chips often adhere to the main cutting edge by the plastic deformation under the condition of high pressure and temperature. Because the formation of bonding chips usually undergoes an adhesive–peeling–adhesive dynamic process, it will take away parts of the tool material periodically and thereby exacerbate the tool wear. When cutting length (l) reaches 2,056 m, the flank wear land of the uncoated tool has been further expanded, and the tool fails in the normal wear stage as shown in Figs. 8 and 10. On the contrary, the multi-layer coated tool suffers minor tool wear in the entire turning operation ascribed to the effective protection from the coating material. The tool almost has twice the tool life of the uncoated tool. In addition, no obvious crater wear or adhesive chips are found in its normal wear stage, i.e., l01,962 m. From the above analysis, it is concluded that the multi-layer coated tool can obtain better cutting performance and longer tool life than the uncoated tool in high-speed turning of AISI 1215 steel. 3.3 Tool wear rate analysis To evaluate tool wear objectively, a new tool wear evaluation indicator, i.e., the surface wear rate (hs) proposed by Astakhov [27, 28], is introduced in this investigation. The relative surface wear is the radial wear per 103 sm2 of the

Intensity (c/s)

Co2p3/2 in Co

13000 12000 11000 10000 800

795

790

785

780

775

770

Binding energy (eV)
3000

(c)
2500

W4f7/2 in WO3 W4f7/2 in WC W4f7/2 in WO2

Intensity (c/s)

2000 1500 1000 500 45.0

42.5

40.0

37.5

35.0

32.5

30.0

27.5

Binding energy (eV)
Fig. 18 XPS spectra of Fe2p3/2, Co2p3/2, and W4f7/2 of point 1 on the tool rake face. a XPS spectrum of the iron element. b XPS spectrum of the cobalt element. c XPS spectrum of the wolfram element

Int J Adv Manuf Technol
O Fe Fe

2

W

3

2

Zone C

VC
C Fe C Co W Co Co W W W W W

3

Fig. 19 SEM and EDS analyses of the tool flank face for the uncoated tool at the cutting length of 903 m

wear, and the initial length of the tool path, respectively, and lt is the total length of the tool path. As follows from Eq. (8), the relative surface wear is inversely proportional to the overall machined area and, in contrast to it, does not depend on the selected wear criterion. Figure 12 shows the influence of feed rate (f) and cutting speed (vc) on the tool wear rate of uncoated and multi-layer coated tools. It is notable that the influence of feed rate (f) on the tool wear rate varies with the cutting speed (vc). When cutting speed is low, i.e., vc 0100 m/min, increase of feed rate (f) may favor the reduction of surface wear rates of both uncoated and multi-layer coated tools. According to Astakhov [27], there are five factors (feed rate, cutting speed (cutting temperature), transient surface, etc.) affecting the tool wear rate in the machining process. The mutual effect of the above factors may affect the tool wear rate in considerably different ways, depending upon the parameters and characteristics of a particular cutting system. When cutting speed is low, the predominant factor affecting tool wear rate is feed rate (f). On one hand, the increase of feed rate may decrease the length of the tool path (especially for a given length of the workpiece). As a result, the cutting (contact) time decreases, as well as the corresponding tool wear, which leads to the reduction of surface wear rate. On the other hand, increase of feed rate (f) also makes the uncut chip thickness (tun) greater than the depth of cold working (dcw) (as shown in Fig. 13), so the main cutting edge will not cut the cold-worked work material characterized by a greater strength and higher hardness. Therefore, the higher feed rate f, the higher tun/dcw, and then the surface wear rate decreases. However, when the turning

operation is conducted under the high-speed cutting condition, i.e., vc 0500 m/min, surface wear rate (hs) is nearly proportional to feed rate (f). This is mainly because in such case, the cutting temperature becomes a dominant factor, and it can be higher than the optimal cutting temperature (referred to as the cutting temperature, at which the combination of minimum tool wear rate, minimum stabilized cutting force, and highest quality of the machined surface is achieved) [27]; then, any increase of feed rate (f) will lead to a direct increase of cutting temperature. The high cutting temperature may aggravate the tool wear (adhesive wear, oxidation wear, etc.) on both tool rake and flank faces, which thereby accelerates the surface wear rate (hs) correspondingly. In addition, the multi-layer coated tool suffers lower tool wear rate than the uncoated tool whether under a low-speed cutting condition or high-speed cutting condition. This phenomenon can be attributed to the beneficial role of the coating material in improving the cutting condition of the tool–chip interface, e.g., reducing the cutting forces, decreasing the cutting temperature, reducing friction, etc., which slows down the wear rate correspondingly. 3.4 Wear mechanism for uncoated tool In order to further reveal the wear mechanism of the uncoated tool in turning of AISI 1215 steel, SEM, EDS, and XPS were conducted on the worn tool surfaces. Figure 14 presents the SEM and EDS analyses (Table 3) of the tool rake face at the cutting length of 1,314 m. It was

Fig. 20 SEM and EDS analyses of tool flank face for uncoated tool at the cutting length of 2,692 m

Int J Adv Manuf Technol Table 5 EDS results of the points shown in Fig. 19 (in weight percent) Fe Point 2 Point 3 66.27 – O 24.77 – C 8.96 9.87 W – 7.51 Co – 82.62

Intensity (c/s)

seen that the rake face was worn out and some materials adhered to the worn area. An EDS detector was used to examine four regions of the crater wear area (regions A, B, C, and D as shown in Fig. 14). Region D was an unworn one, and EDS analyses showed that the uncoated tool substrate consisted of W, Co, and C elements. Regions A, B, and C were all worn ones including a large amount of Fe and O elements as shown by EDS analyses, indicating that there was a strong chemical affinity of the work material with the uncoated carbide tool in the turning operation. In addition, no Mn or S element was detected on the crater area (regions A, B, and C). The above phenomenon demonstrated the occurrence of adhesive wear and the absence of sulfide (MnS) lubricant zone on the tool rake face during the turning process. As illustrated in Fig. 15, the lack of a sulfide lubricant zone on the tool rake face also made the cutter and the chip have direct contact in the machining process. Consequently, the uncoated tool was deprived of the protective effect of adhering inclusions and tended to suffer severe friction and impact between cutting tool and workpiece. In such case, the predominant adhesive wear would cause the bonding layer to be worn away easily and further accelerate the tool wear rate. When machining to the cutting length of 2,692 m, the crater wear area was further expanded as shown in Fig. 16 and Table 4. Comparing Fig. 16 with Fig. 14, regions E, F, and G along the main cutting edge contained a large amount of Fe, O, W, and C elements. The presence of W and C elements also indicated the crack occurred at these areas, which meant that the adhered layers of the worn tool would be detached from the tool if the tool continued to cut the AISI 1215 steel. Besides, the absence of Mn and S elements on regions E, F, G, H, and I also demonstrated that the MnS inclusions did not form a lubricant zone on the tool rake face. Because the bonding layers were usually formed under a high-temperature environment, it would provide the condition for oxidation wear on the tool surface. To investigate the chemical composition of the oxidation resultants, an
Table 6 EDS results of point 4 and region J shown in Fig. 20 (in weight percent) Fe Point 4 Region J 16.62 56.95 O – 34.36 C 8.27 8.69 W 67.99 – Co 7.12 –

XPS spectrum of point 1 on the tool rake face was conducted after a total cutting length of 2,692 m. The XPS survey spectra were fitted by using the XPS peak fit software. Figure 17 showed that the spectrum was dominated by the presence of Fe2p3/2, Co2p3/2, W4f 7/2, and O1s peaks, since the bonding energies of O1s in different oxides are too close to find the exact chemical compositions of the oxidation area. It is necessary to investigate further according to the XPS spectra of Fe2p3/2, Co2p3/2, and W4f7/2 of point 1. Figure 18 shows the XPS spectra of Fe2p3/2, Co2p3/2, and W4f7/2 of point 1 on the tool rake face. It was notable that the oxidation resultants of point 1 were a combination of Fe2O3, Fe3O4, Co2O3, Co3O4, WO2, and WO3. This evidence demonstrated that the adhered Fe element and the tool substrate elements Co and W had already reacted with oxygen from the air atmosphere. Since the oxidation resultants were often soft, they could be easily detached from the tool surface and took away parts of the tool material under the severe impact of turning force, further aggravating tool cracking and micro chipping. In addition, the peeled areas would also be replaced periodically by new bonding layers, which made the tool suffer a bonding–peeling–bonding dynamic process. Moreover, since the sulfide lubricant zone could not form on the tool rake face, it led to a high friction coefficient between tool–chip interfaces, making tool rake face suffer high cutting temperature as shown in Fig. 5a. As a result, crater wear was exacerbated. Figures 19 and 20 show the SEM and EDS analyses (Tables 5 and 6) of tool flank face after a total cutting length of 903 and 2,692 m, respectively. At the cutting length of 903 m, a non-uniform wear was dominated as shown in Fig.19. The maximum width of tool wear zone C (VC) (according to the ISO 3685 standard (1993)) approached 0.35 mm. Some bonding layers were also observed along the main cutting edge. With the progression of the turning operation, the bonding layers were soon detached from the tool and tool flank wear land was further expanded as can be
12000
Fe2p3/2 in Fe2O3 Fe2p3/2 in Fe3O4

10000

8000

6000

4000 725 720 715 710 705

Binding energy (eV)
Fig. 21 XPS spectrum of Fe2p3/2 of region J on the tool flank face

Int J Adv Manuf Technol Fig. 22 SEM and EDS analyses of the tool rake face for the multi-layer coated tool before the turning operation
O Al

M

M

seen in Fig. 20. The flank wear land was full of adhered workpiece/chip material (region J), and a point (point 4) now appeared without this adhered layer. This point presented mostly material from the tool substrate, proving that the wear was caused by attrition. The attrition behavior can be described as the cyclic adhesion and removal of workpiece/chip material from the tool, which also causes removal of tool particles [29, 30]. Under these conditions, microscopic particles of the tool were pulled out and dragged away together with the material flow. So when the flank wear land image was recorded, particles from the tool in point 4 had already been removed, exposing the tool substrate (WC/Co grain). In addition, the scratches of this flank wear were also parallel to the cutting direction that appeared in the region of the tool nose radius, indicating that abrasive action also occurred in this wear land. Through the XPS analysis of region J on the tool flank surface (as seen in Fig. 21), the adhered Fe element had reacted with the oxygen element from the air atmosphere. The main chemical resultants were a combination of Fe2O3 and Fe3O4. Therefore, it was confirmed that the tool flank

face suffered adhesive, abrasive, and oxidation wear in turning of AISI 1215 steel. These three wear patterns had expanded the flank wear land and accelerated the tool wear due to their negative effect on the main cutting edge, making the tool fail even in the normal wear stage (l02,056 m). Therefore, it was observed that the cutting condition of the uncoated tool was exceedingly poorly ascribed to the absence of the sulfide lubricant zone and lack of protection from the coating material in turning of AISI 1215 steel. Its rake face mainly suffered crater, oxidation, and adhesive wear while the flank face suffered adhesive, abrasive, and oxidation wear. Since the uncoated tool is likely to suffer severe adhesive–chip phenomenon and flank wear, it is not suitable for high-speed turning of AISI 1215 steel. 3.5 Wear mechanism for multi-layer coated tool Figures 22 and 23 present the SEM and EDS analyses (Tables 7 and 8) of the tool rake face before and after the turning operation (l04,758 m). Based on Fig. 22, the main

N
P O

P
O Mn C Al S Mn Mn Mn

S Mn

N

Mn O C

Al

O

S Mn S Mn Mn

Fig. 23 SEM and EDS analyses of the tool rake face for the multi-layer coated tool at the cutting length of 4,758 m

Int J Adv Manuf Technol Table 7 EDS results of region M shown in Fig. 22 (in weight percent)

Al Region M 45.10

O 54.90

(a)

elements of the tool rake face were Al and O belonging to the coating material before the turning test. By comparing Fig. 23 with Fig. 22, it was notable that a bright and smooth land was formed along the main cutting edge after the turning operation. Three regions, N, O, and P of the smooth land, were examined by the EDS detector. It was seen that the examined regions O and P all contained Mn, S, Al, O, etc. Through the XPS analyses of point 5 on the tool rake face after a total cutting length of 4,758 m, the main compounds of the smooth land were MnS and Al2O3, indicating the existence of the manganese sulfide (MnS) lubricant zone as shown in Fig. 24. In order to reveal the effect of the sulfide lubricant zone on the tool wear. Figure 25 presents a schematic diagram of a diffusion model with an adhering MnS lubricant zone. Under the tool–chip contact condition, the inclusions in the AISI 1215 steel in cutting zones may become soft and be in a plastic state [10]. Ascribed to the severe tool–chip friction and pressure, the inclusions will obtain further deformation and be extruded on the tool rake face. When the inclusions reach the rake face, the inclusions can accumulate on the tool surface as a protective layer and form a lubricant zone. As illustrated in Fig. 25, this sulfide (MnS) lubricant zone can cut off the direct contact between the tool rake face and chip material and prevent welding or seizure of the work material onto the tool, which favor the reduction of the abrasive wear. Due to the excellent wetting and adhering ability between the inclusions and the (Ti, N) Al coating, the chip and cutting tool will not adhere easily, and thus, the adhesive wear is reduced. Moreover, the adhered layer can inhibit the diffusion of work material as a diffusion boundary. Besides, the low heat conductivity of adhering inclusions also makes the temperature of the layer–tool interface lower than that of the chip–layer interface. As a result, the diffusion wear is dramatically decreased. From the aspect of physical analysis, though the presence of sulfide cannot reduce the internal friction coefficient, it does have excellent sufficient adhesion to the cutting tool and can cover the rake face in the form of films in proper
Table 8 EDS results of the regions shown in Fig. 23 (in weight percent) Al Region N Region O Region P – 20.78 5.29 O – 37.64 20.06 Mn 59.05 9.69 26.13 S 40.95 5.29 14.90 C – 26.60 33.62

Point 5

14000

(b)
Mn2p3/2 in MnS

13000

Intensity (c/s)

12000

11000

10000 655
6000

650

645

640

635

630

Binding energy (eV)

(c)
Al2p in Al2O3

5000

Intensity (c/s)

4000

3000 85 80 75 70 65

Binding energy (eV)
Fig. 24 XPS spectra of point 5 on the tool rake face at the cutting length of 4,758 m. a Position of point 5. b XPS spectrum of the manganese element. c XPS spectrum of the aluminum element

cutting condition, which prevents the adhesion on the chip– tool interface and reduces the external friction. The reduction of the external friction coefficient can decrease the cutting force, cutting temperature, and tool wear. In addition, it can also inhibit the formation of built-up edge and greatly improve the surface finish of the machined surface. The sulfide lubricant zone can dramatically improve the cutting conditions of the primary and second shear zones. In the primary shear zone, MnS is elongated along the shear plane direction, which decreases the actual anti-shearing area and reduces the sliding contact area between different layers of metal. As a matter of fact, MnS can be regarded as

Int J Adv Manuf Technol Fig. 25 Schematic diagram of the diffusion model with the MnS lubricant zone

the microcrack in the steel which will increase the chip breakability during machining operation [10, 31]. In the second shear zone, the sulfide can be elongated very long, and its thickness is less than 0.1 μm. Furthermore, the density of MnS is very high in the bottom of the chip, making the chip easy to shear off. When the sulfide inclusions extend to the contact interface between the chip bottom and the rake face, they attach to the tool surface and reduce the adhesion between the tool rake face and the work material. As a result, they reduce the tool–chip contact length, increase the shear angle, and decrease the cutting forces. Ascribed to the beneficial effect of sulfide inclusions, tool life is improved significantly. As can be seen in the magnified area R in Fig. 26 and Table 9, abrasive wear was found to take place on the tool flank face where many grooves were found on its surface. The type of wear is considered to be due to microscopic hard particles contained in the work material and/or dislodged abrasive grains from the tool material abrading the multi-layer coating. This is essentially a micro-cutting process that produces chips and leaves grooves, and the abrasive wear rate (wa) can be represented as [32]:

wa ?

dWa ? cvc dt

?9?

where wa is the abrasive wear rate, Wa is the mass loss due to abrasion of the tool, t is cutting time, c is a constant, and vc is cutting speed. Region Q, at the bottom of the flank wear land, was filled with elements from sulfide inclusions (S and Mn) via the EDS analyses, indicating that the inclusions from the steel matrix had reached the wear land. The flank wear of the tool was affected markedly by the behavior of sulfide inclusions in the third shear zone amongst other things. Ascribed to the high deformability of sulfide inclusions, the microvoids initiated by them were thinner and more elongated, and tended to close easily in the third shear zone [9]. The existence of microvoids in the steel matrix would be also helpful for the reduction of tool flank wear, which slowed down the flank wear rate and prolonged the tool life. Based on the above analysis, it was concluded that the multi-layer coated tool was conducive to the formation of the sulfide (MnS) lubricant zone in the turning operation, which thereby reduced the tool wear and improved the
Abrasive grooves

Q

VB

R
Fe Mn

S

Q

6

Mn Fe Mn C Fe

Fig. 26 SEM and EDS analysis of tool flank face for multi-layer coated tool at the cutting length of 4,758 m

Int J Adv Manuf Technol Table 9 EDS results of the regions shown in Fig. 26 (in weight percent) Fe Region Q 48.58 Mn 21.94 S 15.41 C 14.07

cutting performance of the cutter. The multi-layer coated tool played an effective role in antifriction wear and obtained longer tool life due to the presence of the sulfide inclusions in turning of the steel. Adhesive wear, crater wear, and abrasive wear were observed to be the main wear mechanisms for the tool.

4. The multi-layer coated tool was confirmed to form the sulfide (MnS) lubricant zone along its main cutting edge. The sulfide inclusions can form a diffusion boundary inhibiting diffusion of the work material and play the lubricant effect on the tool–chip interface. In addition, the substances can initiate thin microvoids in the third shear zone which decreases tool flank wear. The cumulative effect of coating material and sulfide inclusions result in the excellent cutting performance of the tool and minor tool wear. It was also observed that the tool rake face suffered crater wear and adhesive wear while the flank face suffered abrasive wear.

4 Conclusions In this study, an FEM and experimental study on high-speed turning of AISI 1215 steel had been performed using uncoated and multi-layer coated carbide tools. FEM simulations of cutting temperature and tool Von Mises stress were both studied in order to facilitate the wear mechanism analysis. Based on the comprehensive analysis, the major conclusions can be drawn from this research: 1. The uncoated tool was observed to play the role of dissipating a large amount of cutting heat during the turning process. This phenomenon caused the rake face to suffer a high concentrated cutting temperature which resulted in severe crater wear. In contrast, cutting heat for the coated tool was mainly dissipated by the chip material which can be attributed to the lower thermal conductivity of the coating material (especially the Al2O3 coating layer) causing a large amount of cutting heat to transfer to the chip. Therefore, the coated tool suffered a relatively lower temperature than that of the uncoated tool. 2. Tool stress simulation demonstrated that Von Mises stress distribution of the uncoated tool was uneven, and its peak stress zones were located on both tool rake face and tool nose. However, the peak Von Mises stress of the coated tool was only focused on the tool nose, and its value reached as high as 6,920.84 MPa which was 126 % larger than that of the uncoated tool ascribed to the enlarged cutting edge radius by the coating material. 3. The uncoated tool was not conducive to form the sulfide (MnS) lubricant zone. Severe adhered chips and high tool wear rate were observed in the turning operation due to the poor cutting condition. Therefore, it was not an alternative method for processing this steel. Its rake face mainly suffered crater, oxidation, and adhesive wear, and the flank face suffered adhesive, abrasive, and oxidation wear.

Acknowledgments The authors acknowledge the financial support of the National Natural Science Foundation of China (No.50935001), Important National Science & Technology Specific Projects (2012ZX04003-051, 2012ZX04012-021 and 2011ZX04015-031), National Key Basic Research Program under grant nos. 2010CB731703 and 2011CB706804 and “Shu Guang” project supported by Shanghai Municipal Education Commission and Shanghai Education Development Foundation (2011-2012).

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