Quantitative analysis of frictional behavior of cupronickel B10 at the tool-chip interface during dry cutting

Accepted Manuscript Quantitative analysis of frictional behavior of cupronickel B10 at the tool-chip interface during dry cutting Mohan Hao, Daochun Xu, Fuqiang Wei, Qingqing Li PII:

S0301-679X(17)30448-6

DOI:

10.1016/j.triboint.2017.09.033

Reference:

JTRI 4897

To appear in:

Tribology International

Received Date: 8 July 2017 Revised Date:

7 September 2017

Accepted Date: 27 September 2017

Please cite this article as: Hao M, Xu D, Wei F, Li Q, Quantitative analysis of frictional behavior of cupronickel B10 at the tool-chip interface during dry cutting, Tribology International (2017), doi: 10.1016/ j.triboint.2017.09.033. This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.

AC C

EP

TE D

M AN U

SC

RI PT

ACCEPTED MANUSCRIPT

ACCEPTED MANUSCRIPT

Quantitative analysis of Frictional Behavior of Cupronickel B10 at the Tool-chip interface During Dry Cutting

RI PT

Mohan Haoa, Daochun Xua,*, Fuqiang Weia, Qingqing Lia a School of Technology, Beijing Forestry University, Beijing 100083, China *Corresponding author. Tel: +86 10 62338153; E-mail: [email protected]. Abstract: There are gaps and misconceptions in our understanding of tool–chip friction. The present paper presents a novel method of studying the tribological properties of cupronickel B10 during dry cutting. Three types of experiments—the split Hopkinson pressure bar test, SRV test and orthogonal cutting test—were conducted. Following the tests, topographies of the worn surfaces were analyzed employing scanning electron microscopy. The temperature,

SC

normal load and cutting speed effect coefficient were obtained by taking full advantage of each test. Finally, a new empirical friction model for the dry cutting of cupronickel B10 at the tool–chip interface was obtained. The model clarifies the mechanisms of tool–chip friction.

ac a0 aw Lf At

σ τf µ

µ0

1

Introduction

TE D

β γ0

Nomenclature Shear angle (deg) Fn Normal load (N) Friction angle (deg) FC Cutting force (N) Rake angle (deg) FT Thrust force (N) Nch Compression force at the tool–chip interface Cutting depth (mm) (N) Chip thickness (mm) Fch Shear force at the tool–chip interface (N) Chip width (mm) V Cutting speed (m/min) Tool–chip contact length (mm) KT Temperature effect coefficient Actual apparent contact area (mm ) KF Normal stress effect coefficient Normal stress (MPa) KV Cutting speed effect coefficient T Temperature of the tool–chip interface (°C) Frictional stress (MPa) Tool–chip friction coefficient Tr Room temperature (°C) Static friction coefficient Tm Melting temperature (°C)

EP

∅

M AN U

Key words: Tool–chip friction, Cupronickel B10, Dry sliding/machining, Quantitative analysis

Frictional phenomena that occur during the cutting process are complex because they involve three objects of study,

AC C

namely the tool, chip and workpiece, and various affecting factors. More and more researchers are becoming interested in the phenomena owing to their effects on tool failure, tool life [1] and machining accuracy [2]. After years of research, a system with which to study tool–chip friction has been developed. Studies have presented two kinds of research methods, namely methods involving the cutting process itself and laboratory tests [3, 4, 5]. Two-dimensional orthogonal cutting is a commonly used cutting process approach and its mechanics is introduced by [1, 6]. In the experimental method, a pin-on-disc apparatus, such as an SRV testing machine [7–11], is frequently used to study frictional behavior during cutting. Additionally, the pin-on-ring system [3, 4, 5, 12, 13] is becoming increasingly popular. During an experiment, a pin having cylindrical geometry rotates with helical movement to refresh the contact interface. This produces experimental conditions that are much closer to the real cutting process. To reach realistic values of pressures, temperatures and sliding velocities, Zemzemi F et al presented a new tribometer based on a modified pin-on-ring system [5]. Moreover, the simulation method [14–17], especially the finite element method, has become more and more popular. Although simulation can save time and money, preparatory work is required and simulation results must be verified by experimental results. A predictive machining theory was proposed by comparing finite element results with experimental

ACCEPTED MANUSCRIPT data [17]. Cutting forces are found to decrease with increasing cutting speed mainly because of the evolution of the tool–chip contact length. It is well known that the temperature, normal load and cutting speed are the three major factors affecting frictional behavior in metal cutting. However, to the best of our knowledge, few previous works had investigated the effects of these three factors at the same time. Furthermore, the effects have generally been only qualitatively investigated. The effects of temperature and load on the friction and wear performance of brake pads have been investigated qualitatively [7, 8]. The dual function of tribofilms was put forward in the cited papers. In addition, the effects of tool geometry [18],

RI PT

such as the rake angle, and tool materials on friction behavior have been research hotspots. From the view of material science, numerous common alloys, such as aluminum alloy [3–5, 14, 19], and composite materials [20] have been investigated. Meanwhile, cupronickel B10, an attractive alloy with good mechanical properties and corrosion resistance, has hardly appeared in research on metal cutting.

The present work quantifies the effects of the three factors in analyzing the frictional behavior of cupronickel B10 at

SC

the tool–chip interface during dry cutting. For this purpose, a novel research method for the study of fictional phenomena in metal cutting is presented. A split Hopkinson pressure bar (SHPB) test, SRV test and orthogonal cutting test were conducted for the same experiment material (cupronickel B10). Following the tests, the topographies of the worn surfaces were analyzed by scanning electron microscopy (SEM). This paper considers the friction coefficient an

2

M AN U

important evaluation index of frictional behavior. A new friction model for cupronickel B10 is obtained.

Theoretical friction model

At the tool–chip interface, the frictional stress τf is perpendicular and proportional to the normal stress σ [19, 21]. The friction coefficient is defined as

µ=τf /σ=τf At /Fn .

(1)

In general, friction at the tool–chip interface can be broadly divided into three stages. In the first stage, friction just

TE D

appears between two different metal surfaces. By the time the frictional surface is rough and actual contact only appears on some peaks. The first stage is generally short and stick–slip friction theory is applicable in this stage. In the second stage, a small amount of debris appears at the tool–chip interface. Once formed, wear particles are trapped within the tool–chip interface [22] and carried away by the flow of the work material along the contact to fill in grooves. This increases the area of contact and friction coefficient (Eq. (1)). In the third stage, more wear debris is generated and a

EP

transfer layer forms, which features three-body abrasions. Part of the trapped wear debris thickens the transfer layer, while the remainder is pushed out. It is obvious that the formed transfer layer makes no difference to the normal load. However, owing to the relative motion of the transfer layer, the frictional stress and thus the friction coefficient tend

AC C

to decrease. A balance is eventually reached and the condition at the tool–chip interface remains relatively stable. Note that wear debris mainly comes from the workpiece with lower yield strength. The effect of the transfer layer on friction and wear has been noted by many researchers [7, 8, 12, 23, 24]. The three stages are shown in Fig. 1.

Fig. 1 Three stages of friction

ACCEPTED MANUSCRIPT The temperature, normal load and cutting speed are the three major parameters of friction. Taking account of their effects on tribological phenomena at the tool–chip interface, a new abstract friction law is expressed as µ=µ0 KT KF KV .

3

(2)

Experimental study

3.1 SHPB test

RI PT

An SHPB test was carried out at strain rates ranging from 1000 to 7000 s-1 to get the relationship between temperature and friction. Specimens were made of cupronickel B10 and were in the shapes of cylindrical rods (∅ 6 mm × 4 mm). Applied heat loads were 100, 80, 70, 50, and 40 °C while the room temperature was 20 °C. A K-type thermocouple was used to measure the temperature of specimens. However, the deformation time in the SHPB test was very short and beyond the measurement capability of the thermocouple. The study thus adopted an inverse method to obtain the deformation temperature.

SC

3.2 SRV test

Sliding wear tests were conducted using a cylinder-on-disc machine with a cylinder of cemented carbide YG6 sliding against the stationary disc specimen of cupronickel B10. A line contact was designed to simulate tool–chip

M AN U

friction. The size of the upper specimen was ∅ 15 mm × 22 mm while the size of the lower specimen was ∅ 24 mm × 7.88 mm. The experimental setup and structure of the tribo-pair are displayed in Fig. 2. The investigated material was a commercial BFe10-1-2 cupronickel alloy. WC–Co cemented carbides are commonly used cutting tools having excellent wear resistance, good hardness and good chemical inertness [1, 23]. Their chemical composition is listed in Table 1. The thermal and physical properties of cupronickel B10 and cemented carbide

EP

TE D

YG6 are given in Table 2. Table 3 gives the mechanical properties of cupronickel B10 and cemented carbide YG6.

Fig. 2 Experimental setup and structure of the tribo-pair

Table 1 Chemical composition of cupronickel B10 and cemented carbide YG6 (%) Ni

Fe

Mn

Zn

Pb

Si

P

S

C

Sn

Cu

B10

9~11

1~1.5

0.5~1

0.3

0.02

0.15

0.006

0.01

0.05

0.03

-

AC C

cupronickel

cemented carbide YG6

WC

Co

94

6

Table 2 Thermal and physical properties of cupronickel B10 and cemented carbide YG6 cupronickel B10

cemented carbide YG6

Density /

∙

8900 Density /

∙

14.6~15.0

Melting

Thermal conductivity /

Specific heat / J

point / ℃

W/(m·k)

/(kg·K)

1050~1200

398

394

Thermal conductivity/

/

Thermal coefficient of

(m ∙ ℃)

expansion 1E-6 /℃

0.19

4.5

ACCEPTED MANUSCRIPT Table 3 Mechanical properties of cupronickel B10 and cemented carbide YG6

/ GPa

B10

Tensile strength / MPa

Yield strength / MPa

373

315

136

cemented

Hardness

bending strength/

carbide

HRA

/

YG6

89.5

145

compressive stren gth/

/ 460

Endurance limit / MPa 80~90

Modulus of elasticity / GPa 630~640

impact toughness/ / 0.3

RI PT

Modulus of elasticity cupronickel

The tribological behavior of cupronickel B10 dry sliding against cemented carbide YG6 was obtained with an Optimol SRV tribotester. Experimental preset parameters included the test frequency, stroke, load, temperature and duration. The test frequency and stroke were respectively set as 50 Hz and 3 mm. The test temperature varied from 200 to 500 °C while loads varied from 30 to 60 N (50–70 MPa). The friction coefficient for various regimes was obtained

SC

directly from the experimental equipment. All friction tests began with the same run in and all lower specimens were used twice. Each test lasted 1 min.

Following testing, worn surfaces were analyzed by SEM to examine the frictional phenomena (Fig. 9).

M AN U

3.3 Orthogonal cutting test

An orthogonal cutting test without lubrication was carried out on a DMG CTX310 eco CNC lathe (Fig. 3). The workpieces—round bars with an outer diameter of 125 mm, inner diameter of 95 mm and thickness of 15 mm—were made of cupronickel B10. The uncoated cutting tool made of cemented carbide YG6 was an insert type and was mounted on a special-purpose tool holder to have a rake angle γ0 of 0°. To minimize experimental error, the cutting tool was renewed once worn and 20 tool pieces were used overall.

The tests were divided into three groups according to the feed rate (0.05, 0.08 and 0.11 mm/rev). In each test, the

TE D

spindle speed was up to 4000 rpm and the cutting speed was changed from 0 to 1400 m/min. The cutting force FC and feed force FT were measured throughout the orthogonal cutting tests by a K-type dynamometer mounted on the tool turret (Fig. 3). The temperature variation during the machining process was recorded by thermocouples embedded in the cutting tool.

AC C

EP

Following each test, the thickness and width of the chip were measured for further study.

Fig. 3 Experimental setup for orthogonal cutting

4

Results and Discussion

4.1 Effect of temperature on friction The effect of temperature on friction was investigated by conducting the SHPB test and SRV test. Due to the differences between the actual temperature T and the measured temperature T10 in the SHPB test, this paper adopted

ACCEPTED MANUSCRIPT the thermal conduction model T10 =0.9404T-5.229 which was obtained in the paper [24]. The technological details can refer to the literature [24]. The SHPB test provided the thermal softening effect curve (Fig. 4). The thermal effect was obtained as =1-[(T-Tr )/(Tm -Tr )] 1.1026 . The temperature effect coefficient is clearly a decreasing function of temperature.

(3)

Figure 5 obtained from the SRV test shows variations in the friction coefficient with temperature at various loads. The friction coefficient decreases as temperature increases from 200 to 500 °C and is consistent with Eq. (3). At higher

RI PT

temperatures, the yield strength of cupronickel B10 is clearly lower and more wear debris is generated. In this case, the transfer layer more readily forms and thickens, decreasing the friction coefficient. Note that the effect of temperature on the friction coefficient of cupronickel B10 dry sliding against cemented carbide YG6 is limited and the friction coefficient finally becomes steady (Fig. 6).

In order to simplify matters, the present paper focuses on the mean temperature at the tool–chip interface. Actually

SC

the temperature distribution of tool–workpiece–chip interface is also a valuable research topic [25]. The value of temperature effect coefficient changes along the tool–chip interface on account of temperature gradient induced by heat

TE D

M AN U

conduction. The effect of temperature distribution on frictional behavior will be investigated further in a future work.

Fig. 5 Comparison of the effect of temperature on the friction coefficient

AC C

EP

Fig. 4 Curve of the thermal softening effect

Fig. 6 Effect of temperature on the friction coefficient at 30 N

ACCEPTED MANUSCRIPT 4.2 Effect of the normal load on friction The effect of the normal load on friction was investigated in an SRV test. The evolution of the apparent friction coefficient versus load is plotted in Fig. 7 for different temperatures. The friction coefficient rises as the normal load rises from 30 to 60 N. In the SRV test, the friction coefficient can be expressed as µ=µ0 KT KF ,

(4)

where µ0 takes a value of 0.586 at 10 N and 20 °C, which approaches to the initial value.

RI PT

In theory, when the normal stress is zero, it has no effect on friction such that the stress effect coefficient is 1. The stress coefficient KF curve is obtained (Fig. 8) using Eqs (3) and (4) and the theoretical value mentioned above. The relationship is given by

KF =e5.15478E-5σ +0.00558σ+0.00546 . (5) The friction coefficient increases with the normal stress. More wear debris is generated with increasing normal 2

SC

stress. Friction later goes through the second and third stages mentioned in section 2. Wear debris is carried away by the flow of the work material along the contact to fill in grooves (Fig. 9), which increases the contact area. The transfer layer forms once the wear debris reaches a certain amount. At an excessive load, part of the generated wear debris firmly

EP

TE D

evolution of the apparent friction coefficient.

M AN U

pressed on the lower specimen forms into ridges and the transfer layer becomes harder to generate. This accounts for the

AC C

Fig. 7 Comparison of the effect of the normal load

Fig. 8 Stress effect coefficient #$ versus normal stress

ACCEPTED MANUSCRIPT (a)

(b)

Fig. 9 SEM photographs of lower specimens subjected to the SRV test at (a) 30 N and 400 °C and (b) 40 N and 400 °C

4.3 Effect of the cutting speed on friction The effect of the cutting speed on friction was investigated in an orthogonal cutting test. To clarify the friction condition, a force system applicable to orthogonal cutting was developed by ME Merchant

Nch =FC cos γ0 -FT sin γ0 , Fch =FC sin γ0 +FT cos γ0 . The external friction coefficient is obtained as µ=Fch /Nch , Lf =ac sin (∅+β-γ0 ) / sin ∅ cos β,



∅= tan [ac cos γ0 /(a0 -ac sin γ0 )], -1

-1

SC

β= tan µ. An expression of the normal stress is obtained from Eqs (6) to (11) as

RI PT

[26]:

γ0 =0°

(8) (9) (10) (11)

(12)

M AN U

σ=Nch /(Lf aw ) &'( σ=FC /{ac aw sin ( tan-1 ac / a0 + tan-1 FT / FC ) / [sin( tan-1 ac / a0 ) cos( tan-1 FT / FC )] }.

(6) (7)

The relationship between the cutting speed and friction is obtained from Eqs (2), (3), (5) and (12) (Fig. 10). A new equation for the effect of the cutting speed on friction is proposed by fitting a curve: KV =1.0365 -0.0419ln(V+2.3281). (13) The result obtained is similar to the friction model, revealing the connection between the adhesive friction coefficient and the local sliding velocity presented by Zemzemi F et al. [5].

It is clear that the cutting speed effect coefficient is lower than 1 in the overall process, and the cutting speed thus has a weakening effect on friction. The cutting speed effect coefficient remarkably decreases in the range of 0–20 m/min 1. 2.

TE D

while it is relatively stable in the range of 20–1400 m/min. Two frictional regimes are hereby defined. V ≤ 20 m/min (regime 1), where the friction coefficient decreases with increasing cutting speed.

V > 20 m/min (regime 2), where the friction coefficient remains stable.

It is known that the chip flow speed is basically proportional to the cutting speed. As the cutting speed rises, the relative motion is greater and more wear debris is generated. In this case, friction has a declining trend. Friction then

AC C

EP

remains relatively stable with cutting speed because most of the wear debris formed is pushed out.

Fig. 10 Cutting speed coefficient #- versusc cutting speed

ACCEPTED MANUSCRIPT 4.4 Friction model A friction model for the dry cutting of cupronickel B10 at the tool–chip interface is proposed according to Eqs (2), (3), (5) and (13): µ=0.586× .1-[(T-Tr )/(Tm -Tr )]

1.1026

/ e5.15478E-5σ

2 +0.00558σ+0.00546

×[1.0365 -0.0419ln（V+2.3281）].

(14)

Three major effect coefficients curves are obtained after substituting SHPB test data into Eq (14) (Shown as Fig. 11). The normal stress effect coefficient is stable due to the marginal influence of cutting speed to the normal stress. The

RI PT

temperature effect coefficient decreases as cutting speed increases from 0 to 200 m/min because of a rise in the cutting temperature. It is obvious that cutting speed and temperature have large impacts on the frictional behavior of cupronickel

M AN U

SC

B10 in the cutting process. And to optimize the cutting parameters, the two factors should be in consideration.

TE D

(a) Feed rate f=0.05 mm/rev

(b) Feed rate f=0.11 mm/rev

Fig. 11 Three major effect coefficients against cutting speed

5

Conclusions

EP

An SHPB test, SRV test and orthogonal cutting test were conducted to clarify the frictional behavior of cupronickel B10 at the tool–chip interface. Three stages of friction were proposed in this paper. The main achievements and conclusions of the present work are summarized as follows. The paper proposed and adopted a new method of studying metal cutting. The method involves three kinds of tests,

AC C

1.

namely an SHPB test, SRV test and orthogonal cutting test. By taking full advantage of each test, the stand-alone effects of temperature, normal load and cutting speed are obtained. This novel idea can be applied in other metal studies. 2.

The temperature has a strong effect on the frictional behavior of cupronickel B10. The friction coefficient was found to decrease as temperature rose from 200 to 500 °C owing to the thicker transfer layer.

3.

During the SRV test, the friction coefficient increased with load (from 30 to 60 N) and was more prominent at lower temperature (200 and 300 °C). The experimental results provided a new quantitative equation that clarifies the effect of normal stress on friction.

4.

The effect of cutting speed on friction was quantified and a novel equation was proposed. The friction coefficient decreases as the cutting speed rises (Fig. 10). Two frictional regimes were identified. The cutting speed clearly plays an important role in regime 1.

5.

Taking account of three factors (i.e., temperature, normal load and cutting speed), an empirical friction model for

ACCEPTED MANUSCRIPT the dry cutting of cupronickel B10 was proposed. The model not only provides input conditions for the dynamic simulation of cupronickel B10 but also predicts the friction coefficient. And in the course of cutting of cupronickel B10, cutting speed and temperature need more attention than normal load through comparative analyses.

Acknowledgements This work was supported by the National Natural Science Foundation of China (51405021).

RI PT

References [1]. Cheng K. Machining dynamics: fundamentals, applications and practices, Springer, London (2008).

[2]. Ai X, Zhao J, Liu ZQ, Deng JX, Liu JG. Research and development of high speed cutting tribology. Key Eng Mater 315-316 (2006) 401-405.

[3]. Ben Abdelali H, Courbon C, Rech J, Ben Salem W, Dogui A, Kapsa P. Identification of a friction model at the tool-chip-workpiece interface in dry machining of a AISI 1045 steel with a TiN coated carbide tool. J Tribol 133 (4)

SC

(2011).

[4]. Ben Abdelali H, Claudin C, Rech J, Ben Salem W, Kapsa P, Dogui A. Experimental characterization of friction coefficient at the tool–chip–workpiece interface during dry cutting of AISI 1045. Wear 286 (2012) 108-115.

M AN U

[5]. Zemzemi F, Rech J, Ben Salem W, Dogui A, Kapsa P. Identification of a friction model at tool/chip/workpiece interfaces in dry machining of AISI4142 treated steels. J Mater Process Tech 209 (8) (2009) 3978-3990. [6]. Altintas Y. Manufacturing automation: metal cutting mechanics, machine tool vibrations, and CNC Design. 2nd Edition, Cambridge University Press (2012).

[7]. Zhang SY, Li YY, Qu SG, Chen WP. Friction and wear behaviour of brake pads dry sliding against semi-interpenetrating network ceramics/Al-alloy composites. Tribol Lett 38 (2) (2010) 135-146. [8]. Zhang SY, Feng SS. Friction and wear performances of brake material dry sliding against a composite with a semi-interpenetrating network structure of ceramics and Al-alloy. Tribol Int 44 (3) (2011) 248-257. Lett 42 (2) (2011) 141-148.

TE D

[9]. Wang ZY, Xia YQ, Liu ZL. Study the sensitivity of solid lubricating additives to attapulgite clay base grease. Tribol [10]. Totolin V, Minami I, Gabler C, Dörr N. Halogen-free borate ionic liquids as novel lubricants for tribological applications. Tribol Int 67 (4) (2013) 191-198.

[11]. Cui XH, Li SH, Jin YS, Song HH. A study of tribological performance and tribo-induced graphitization of

EP

meso-carbon microbeads. Tribol Lett 14 (2) (2003) 53-59. [12]. Shorowordi KM, Haseeb ASM, Celis JP. Velocity effects on the wear, friction and tribochemistry of aluminum MMC sliding against phenolic brake pad. Wear 256 (11–12) (2004) 1176–1181.

AC C

[13]. Rech J, Arrazola PJ, Claudin C, Courbon C, Pusavec F, Kopac J. Characterisation of friction and heat partition coefficients at the tool-work material interface in cutting. CIRP Ann - Manuf Techn 62 (1) (2013) 79–82. [14]. Zhang YC, Mabrouki T, Nelias D, Gong YD. FE-model for Titanium alloy (Ti-6Al-4V) cutting based on the identification of limiting shear stress at tool-chip interface. Int J Mater Form 4 (1) (2011) 11-23. [15]. Bahi S, Nouari M, Moufki A, El Mansori M, Molinari A. A new friction law for sticking and sliding contacts in machining. Tribol Int 44 (7-8) (2011) 764-771. [16]. Samyn P, Schoukens G, Quintelier J, De Baets P. Friction, wear and material transfer of sintered polyimides sliding against various steel and diamond-like carbon coated surfaces. Tribol Int 39 (6) (2006) 575-589. [17]. Atlati S, Moufki A, Nouari M, Haddag B. Interaction between the local tribological conditions at the tool–chip interface and the thermomechanical process in the primary shear zone when dry machining the aluminum alloy AA2024–T351. Tribol Int 105 (2017) 326-333. [18]. Romero PA, Anciaux G, Molinari A, Molinari JF. Friction at the tool-chip interface during orthogonal nanometric machining. Modell Simul Mater Sci Eng 20 (5) (2012) 55007-55022(16).

ACCEPTED MANUSCRIPT [19]. Xu DC, Feng PF, Li WB, Xu C, Ma Y. A friction model having multiple factors for high-speed cutting of aluminum alloy 6061-T6. Int J Adv Manuf Tech 84 (9-12) (2016) 1941-1951. [20]. Chardon G, Klinkova O, Rech J, Drapier S, Bergheau JM. Characterization of friction properties at the work material/cutting tool interface during the machining of randomly structured carbon fibers reinforced polymer with Poly Crystalline Diamond tool under dry conditions. Tribol Int 81 (2015) 300-308. [21]. Zhang CY, Lu JP, Zhang FP, Butt SI. Identification of a new friction model at tool-chip interface in dry orthogonal cutting. Int J Adv Manuf Tech 89 (1-4) (2017) 921-932.

RI PT

[22]. Godet M. The third-body approach: A mechanical view of wear. Wear 100 (1) (1984) 437-452.

[23]. Zhou FJ. A new analytical tool-chip friction model in dry cutting. Int J Adv Manuf Tech 70 (1-4) (2014) 309-319. [24]. Lei N, Xu DC. Deformation temperature and material constitutive model of cupronickel B10. J Mech Sci Technol 31(8) (2017) 3761-3767.

[25]. Sawangsri W, Cheng K. Investigation on partitioned distribution of cutting heat and cutting temperature in micro

SC

cutting, I J M M S 9 (2) (2016) 173-195.

[26]. Qi HS, Mills B. Formation of a transfer layer at the tool-chip interface during machining. Wear 245 (1) (2000) 136-147.

[27]. Kagnaya T, Boher C, Lambert L, Lazard M, Cutard T. Microstructural analysis of wear micromechanisms of

M AN U

WC–6Co cutting tools during high speed dry machining. Int J Refract Met H 42 (2014) 151-162. [28]. Merchant ME. Mechanics of the Metal Cutting Process. I. Orthogonal Cutting and a Type 2 Chip. J App Phys 16 (5)

AC C

EP

TE D

(1945) 267-275.

ACCEPTED MANUSCRIPT Highlights A friction model for the dry cutting of cupronickel B10 is proposed. The stand-alone effects of temperature, normal load and cutting speed are quantified. A new method of studying metal cutting friction is proposed.

AC C

EP

TE D

M AN U

SC

RI PT

Two frictional regimes are defined by the cutting speed.