Surface & Coatings Technology xxx (2014) xxx–xxx

SCT-19728; No of Pages 7

Contents lists available at ScienceDirect

Surface & Coatings Technology

j ourna l homepage: www.e lsev ie r .com/ locate /sur fcoat

Interface delamination study of diamond-coated carbide toolsconsidering coating fractures

P. Lu a,⁎, X. Xiao b, Y.K. Chou a

a Mechanical Engineering Department, The University of Alabama, Tuscaloosa, AL, USAb Research & Development Center, General Motors Corporation, Warren, MI, USA

⁎ Corresponding author.E-mail address: plu2@crimson.ua.edu (P. Lu).

http://dx.doi.org/10.1016/j.surfcoat.2014.08.0800257-8972/© 2014 Elsevier B.V. All rights reserved.

Please cite this article as: P. Lu, et al., Surf. Co

a b s t r a c t

a r t i c l e i n f o

Available online xxxx

Keywords:Coating fractureDiamond coatingFinite element modelingInterface delaminationScratch testing

Interface delimitation is one of the major failure modes of diamond-coated carbide tools in machining. On theother hand, diamond coatings are prone of cracking easily due to its brittleness, which may affect interfacedelaminations. To study any influence between the two failure modes, micro-scratch testing on diamond-coated carbide tools was conducted and finite element (FE) modeling was developed to simulate the scratchingprocess. In scratch testing, normal and tangential forces as well as acoustic emission signals were recorded todetect coating delaminations and crack initiations. Scratched samples were also observed by optical microscopyto determine the corresponding critical load of delaminations and cracking initiations. In the FE scratch simula-tion, a cohesive-zone interface and the extended finite element method (XFEM) were applied to investigatedelamination and coating fracture behaviors, respectively. The cohesive elements were based on a bilineartraction\separation model and XFEM was implemented to model cracking behavior in a diamond coatingwith a damage criterion of the maximum principal stress.The major findings are summarized as follows. The coating fracture energy has a negligible effect on the criticalload for interface delaminations, and similarly, the interface fracture energy has no effect on the critical load forcoating cracking, indicating that the two failure modes are mostly uncoupled for the testing range in this study.From the experiments and simulations, it is estimated that the coating fracture energy of the samples tested inthis research is in the range of 120 to 140 J/m2, and the diamond-carbide interface fracture energy is from 77to 192 J/m2. Moreover, increasing the coating Young's modulus will increase the critical load for coating delam-inations, but decrease the critical load of coating cracking.

© 2014 Elsevier B.V. All rights reserved.

1. Introduction

Hard coatings such as chemical vapor deposited (CVD) diamondfilms on tungsten carbide cutting tools (typically, cobalt cemented,noted asWC–Co) has attracted significant interest for advanced toolingin machining due to their high hardness and strength, low friction coef-ficient, chemical stability, etc. [1]. Much research has been devoted tostudying the wear mechanism of CVD diamond coated tools. Oles et al.[2] reported that abrasive wear on the tool flank face is a commonwear mechanism of CVD diamond coated tools. However, the life ofCVD diamond-coated tools in machining is dominantly limited by coat-ing interface delaminations [3] or coating fractures [4]. Themajor obsta-cle for cost-effective applications of CVD diamond-coated tools is theinsufficient adhesion between the diamond coating and the carbidesubstrate, which results in coating delaminations during cutting. ForCVDdiamond-coatedWC–Co tools, Polini [5] concluded that the presenceof Co in carbide tools leads to a non-diamond carbon layer formation atthe substrate surface, resulting in a weak interface adhesion. On the

at. Technol. (2014), http://dx

other hand, diamond is very brittle and coating cracking due to localfractures may induce catastrophic tool failures. Thus, it is desired toquantitatively characterize the interface adhesion of diamond-coatedtools and further to know how coating fractures may affect interfacedelaminations.

Several experimental methods have been applied to examine thecoating–substrate adhesion [6]. One of widely adopted techniques isscratch testing, which is a useful technique for obtaining comparativevalues of the adhesion strength for hard coatings on a compliant sub-strate [7]. Considering that the coating–substrate interface failure isadhesive, the adhesion strength is a measure when a critical load isreached at which the coating–substrate interface delaminates [8]. Inaddition, scratch testing may also be used to identify coating cracking.von Stebut et al. [9] observed that there is a clear correlation betweenhigh-energy acoustic emission (AE) pulses and coating cracking failures.Similarly, the same argument of the critical load for interface delamina-tions can be applied for coating cracking; i.e., the corresponding criticalload when the cracking failure occurs could be a measurement of thecoating fracture strength.

In scratch testing, a spherical indenter tip slides over a coating sur-face to cause a groove under an incremental normal load. The detailed

.doi.org/10.1016/j.surfcoat.2014.08.080

Fig. 1. (a) Assembly of coating, substrate, and indenter in the FE model, and (b) meshes of the FE model for scratch simulations of CVD diamond-coated cutting inserts.

2 P. Lu et al. / Surface & Coatings Technology xxx (2014) xxx–xxx

description of scratch testing can be found from an earlier study [10].During a scratch test, tangential forces, penetration depths and acousticemission signals can be monitored and the morphology of the scratchtrack can be observed simultaneously or afterwards. For coating adhe-sion evaluations, the normal load that causes the coating to detachfrom the substrate can be considered as the critical load for coating de-laminations [11]. On the other hand, while a hard coating, e.g., diamondin this study, can withstand compressive stresses induced by theindenter to a certain extent, a brittle diamond film may fracture if ahigh tensile or shear stress field is developed [12]. Coating crackingfailure occurs when the mechanical work of the brittle failure in thecoating is equal to the energy release rate from coating cracks. Whenthe mean compressive stress over an area in the coating exceeds acritical value, the coating cracking failure thatfirst occursmaybedetect-able by high-energy AE pulses [7].

Though scratch testingmay offer critical load information for coatingcracking and interface delaminations, it may not shed light of theinteractions between the twomodes by using only limited sets of exper-iments. On the other hand, numerical studies of scratch process model-ing and simulations may be useful to investigate coating cracking andinterface delaminations together. Simulations of a scratch processhave been investigated before. Several challenges associated withscratch simulations include the interface behavior modeling, coatingbrittle failures, and possible interactions between the two, etc. In a

Fig. 2. Traction–separation response examples f

Please cite this article as: P. Lu, et al., Surf. Coat. Technol. (2014), http://dx

previous study about diamond-coated tools [10], a three-dimensional(3D)finite element (FE) simulationwas developed to investigate coatingdelaminations alone when a rigid indenter slides on a diamond-coatedcarbide substrate. In that study, the cohesive zone concept, a bi-linearconstitutive law, was applied to model the interface behavior [13].However, the coating fracture phenomenon which may influence dia-mond coating delaminations was not considered. To better understandthe adhesion of diamond-coated carbide tools, it is essential to investi-gate interface delamination by simultaneously considering the coatingfracture phenomenon. Cracking behavior analysis of hard coatingssuch as CVD diamond is a challenging task too. Recently, the extendedfinite element (XFEM)method in ABAQUS software using enriched ele-ments has been applied for cracking analysis. XFEMwasfirst introducedbyBelytschko and Black [14]. It is an extension of the conventionalfiniteelement method based on the concept of partition of unity by Melenkand Babuska [15], which allows local enrichment functions to be incor-porated into a finite element approximation. XFEM can be used for theestimation of multiple crack propagation in indentation simulations,with or without pre-cracks defined, and it is independently definedfrom the existence of any predefined crack or its propagation pathwith-out alternating the finite element mesh. XFEM has been used to modelsome applied mechanics problems after its development. For example,quasi static crack propagations in 2D and 3D by using XFEMwere intro-duced by Daux et al. [16]. In addition, Combescure's group [17]

or XFEM: (a) linear and (b) nonlinear [22].

.doi.org/10.1016/j.surfcoat.2014.08.080

0

4

8

12

16

20

0 10 20 30 40 50

Pene

trat

ion

dept

h (µ

m)

Normal Force(N)

(b)

(a)

Fig. 3. (a) Microscopic image of a scratch groove, and (b) penetration depth vs. loading force from scratch testing of diamond-coated cutting tool (Sample 1).

3P. Lu et al. / Surface & Coatings Technology xxx (2014) xxx–xxx

investigated dynamic fracture of brittle materials by XFEM, and Baoet al. [18] simulated the initiation and propagation of a cone crack invacuum glazing units by using an XFEM model.

This study aims at better understanding of the delamination behav-ior of diamond-coated carbide tools by considering also coating frac-tures. Micro-scratch testing experiments were first conducted and thecritical loads for coating crack and interface delaminations were identi-fied and analyzed. In addition, a 3D FE model of a scratch process usingABAQUSwas developed with XFEM implemented in a diamond coatingto analyze crack initiations and propagations. Moreover, a cohesive-zone layer between the coating and the substrate was included tostudy the diamond–carbide interface behaviors. The FE model wasthen applied to investigate the effects of the coating fracture energy,the interface fracture energy, and the coating Young's modulus on coat-ing cracking and interface delaminations.

2. Experimental details

The test specimens were CVD diamond-coated carbide tools. Thecarbide substrates were square-shaped inserts (SPG422) of fine WCgrains and 6 wt.% cobalt. For the coating process, diamond films weredeposited using a high-powermicrowave plasma-assisted CVD process.A gas mixture of methane in hydrogen, 750 to 1000 sccm with 4.4 to7.3% ofmethane/hydrogen ratio, was used as the feedstock gas. Nitrogengas, 2.75 to 5.5 sccm, was inserted to obtain nanostructures by pre-venting columnar growth. The pressure was about 30 to 55 Torr andthe substrate temperature was about 685 to 830 °C. A forward powerof 4.5 to 5.0 kW with a low deposition rate obtained a thin coating,5 μm (Sample 1); a greater forward power of 8.0 to 8.5 kW with ahigh deposition rate obtained a thick coating, 25 μm (Sample 2). Thediamond coatings were characterized and the results were reported in

0

10

20

30

40

50

60

70

80

90

0 10 20

Aco

ustic

Em

issi

on(%

)

Normal Fo

AE

Spot 1

Fig. 4. AE response and tangential force vs. normal load from a

Please cite this article as: P. Lu, et al., Surf. Coat. Technol. (2014), http://dx

an earlier study [19]. In summary, the surface roughness, in Ra, wasabout 0.33 μm and the grain size was on the order of 100 nm estimatedby scanning electron microscopy. From nanoindentation testing, theYoung's modulus and hardness were about 700 GPa and 80 GPa,respectively.

A micro-scratch tester from CSM Instruments, Revetest ScratchTester was used to conduct experiments at the room temperature. Themaximum load limit of this system is 200 N. A diamond indenter witha tip radius of 100 μm was used. The scratch speed was 6 mm/minwith a scratch length of 3 mm. A linear loading method was used inthe scratch testing with a loading rate of 100 to 300 N/min, and thetotal loading time for each test was 0.5 min. The unloading rate was10 N/s. During the scratch test, tangential forces, AE signals, and thedepth of the scratch were recorded. A microscopic camera was used toobserve and record the scratch tracks and delamination areas on thesamples after testing. To determine the critical loads of coating crackingand delamination for the tested diamond-coated tools, the scratch testswere carried out using a progressive load method. In addition, theindenter tip was checked before and after each scratch test to confirmits quality by measuring the tip radius using an optical microscope. Ifthere was an abrupt increase of the tip radius observed, the indenterwould be regarded as failed and a new indenter would be replaced forfurther scratch testing.

3. Scratch process simulations

A 3D FE model was developed to simulate the scratch process usingABAQUS software. Onehalf of the specimenwasmodeled because of thesymmetry condition. The finite element model for the specimen isshown in Fig. 1(a): 1.25 mm long, 1.0 mm wide and 1.02 mm high.Different scratch configurations for simulations were conducted. The

0

5

10

15

20

25

30

35

40

30 40 50

Tan

gent

ial F

orce

(N)

rce(N)

FtSpot 2

scratch test on diamond-coated cutting tool (Sample 1).

.doi.org/10.1016/j.surfcoat.2014.08.080

(a) (b)25 µm 25 µm

Fig. 5.Magnified images of (a) Spot 1 for a normal load around 35.8 N, and (b) end of the scratch (load of 50 N) from a scratch test on diamond-coated cutting tool.

4 P. Lu et al. / Surface & Coatings Technology xxx (2014) xxx–xxx

diamond indenter had a tip radius of 100 μm and was assumed rigid.During the scratch simulation, the indenter moved over the top surfaceof the diamond coating for a distance of 500 μm, to reduce unnecessarycomputational costs, with progressive loading. The frictional coeffi-cient between the indenter and the coating surface was set as 0.1.Fig. 1(b) shows the details of themeshes of themodel. An 8-node linearbrick (C3D8) was used to mesh the diamond and the substrate with aprogressive meshing method. For the meshes along the scratch direc-tion, the average element size was 8.3 μm with a minimum of 5.7 μm,and the minimum size was 5.6 μm in the transverse direction.

To simulate cracking in a diamond coating by XFEM, enrichedelements were applied to the coating. The enrichment functions forthe elements typically consist of a near-tip asymptotic function thatcaptures the singularity around the crack tip and a discontinuous func-tion that represents the jump in the displacement across the cracksurfaces [20]. The approximation for a displacement vector function uwith the partition of unity enrichment is in the form of [20]:

u ¼ ∑Ni¼1Ni xð Þ ui þ H xð Þαi þ∑4

l¼1∑2j¼1 Fl

j xð Þbilh i

; ð1Þ

whereN is thenumber of nodes in themesh;Ni(x) is a usual nodal shapefunction for the nodal i; ui is the usual degrees of freedom (DOFs) for thenode i; αi and bil are the DOFs associated with the Heaviside “jump”function H(x), with value 1 above and below the crack. The asymptoticcrack tip function for an isotropic elastic material, Fl(x), can be given as:

Fl xð Þ ¼ ffiffiffir

psin

θ2;

ffiffiffir

pcos

θ2;

ffiffiffir

psinθ sin

θ2;

ffiffiffir

psinθ cos

θ2

� �; ð2Þ

where (r, θ) are the local polar coordinates at a crack tip.The formulae and laws that govern the behavior of XFEM-based

cohesive segments for a crack propagation analysis are very similar tothose used for cohesive elementswith the traction–separation constitu-tive behavior. The available traction–separationmodel in software such

Table 1Summary of the critical load results from scratch testing on CVD diamond-coated cuttingtools.

Sample Coatingthickness/μm

Tipradius/μm

Max.load/N

Critical loadfor crackinitiation/N

Critical loadfor coatingdelamination/N

1 5 100 50 17 35.8100 50 20 29.6100 50 29 42.9

Average 22 36.12 25 100 150 27 52.9

100 80 18 42.8Average 22.5 47.9

Please cite this article as: P. Lu, et al., Surf. Coat. Technol. (2014), http://dx

as ABAQUS assumes initially a linear elastic behavior followed by theinitiation and evolution of damages, shown in Fig. 2.

The coating material, diamond, was modeled as an elastic solid, andcracks for a brittle diamond coatingwill be initiatedwhen themaximumprinciple stress in the coating reaches the critical value, obeying thetraction–separation law, with a mixed mode behavior of power law(power = 1). Thus, the combining effects of normal and shear modeswill be considered. For the properties of diamond coatings, typicalvalues were assigned: Young's modulus E = 1200 GPa and Poisson'sratio υ = 0.07. The fracture toughness of CVD diamond coatings hasbeen reported in a wide range, between 5 and 13.5 MPa·m1/2 [21].Using the data of the Young's modulus in the literature, the fractureenergy of CVD diamond coatings could be calculated by the following:

∅n ¼ KIC2=E: ð3Þ

This yields a fracture energy for CVD diamond coatings between24 J/m2 and 172 J/m2, with the mean value of 98 J/m2, which will beused in the numerical analysis for both the normal and shear modes,considering they are equal.

The carbide substrate was modeled as an elastoplastic materialwith isotropic hardening obeying the Ramberg–Osgood law, Young'smodulus E = 619.5 GPa, Poisson's ratio υ = 0.24, yield strengthσy = 3.605 GPa, and hardening exponent n = 0.244 [22].

In addition, a bilinear cohesive zone model (CZM) was included tomodel the coating–substrate interface mechanical behavior, detailedin [22]. The interface cohesive zone properties between the coatingand the substrate such as the strength and characteristic length werederived from a previous study [22] with the characteristic values as:543 MPa and 314 MPa for maximum normal and shear strengths,1.41 μm displacement of failure, and an upper bound interface fractureenergy of 387 J/m2.

4. Results and discussion

Scratch testing conducted on Sample 1 included three repeated testswith a maximum load of 50 N. Fig. 3(a) shows the microscopic image ofone typical scratch groove and Fig. 3(b) displays the corresponding re-sults of the penetration depth along with the normal force on Sample1. It is observed that the penetration depth increases smoothly untilthe normal force reached 17N, followed by a sharp transition, indicatingthat the coating crack was initiated then. Further, the penetration depthbegan to rapidly increase, when the normal force was around 35.8 N,until reaching a maximum value of 18.7 μm. It is concluded that theindenter might be in contact with the substrate at about 35.8 N, whichcan be considered as the critical load for coating delaminations; themicroscopic image of the scratch tracks confirmed the results.

Fig. 4 shows the AE signal and tangential force vs. applied normalload during the scratch test. The tangential force has a noticeable

.doi.org/10.1016/j.surfcoat.2014.08.080

Fig. 6.Maximum principal stress contour (from scratch simulations) in the coating and substrate areas when a crack is initiated.

5P. Lu et al. / Surface & Coatings Technology xxx (2014) xxx–xxx

decrease at the normal load of about 17 N (Spot 1), which is also anindication of the coating cracking. On the other hand, the AE signalonly showedminor fluctuation features at about 17 N, related to coating

87

Unit: 1000 GPa

Unit: 1000 GPa

87

Fig. 7.Normal stress contour (from scratch simulations) at the cohesive-zone interface during (a

Please cite this article as: P. Lu, et al., Surf. Coat. Technol. (2014), http://dx

cracking. It is also observed that the tangential force increases smoothlybefore the normal load reached about 35.8 N, but followedwith consid-erable fluctuations. In addition, an abrupt amplitude decrease of AE

46

552Sliding direction

229.128

(a)

(b)

Sliding direction

46

518Sliding direction

(c)

) coating delamination initiation, (b) end of scratch process, and (c) completely unloading.

.doi.org/10.1016/j.surfcoat.2014.08.080

Table 2Comparison of the critical loads in scratch testing on CVD diamond-coated tools frombetween finite element (FE) simulations and experiments.

Sample tc/μm

Crack criticalload/N

Delaminationcritical load/N

Coatingfractureenergy/J/m2

Interfacefractureenergy/J/m2

Experiment FE Experiment FE

1 5 22 20.3 36.1 36.7 140 1922 25 22.5 20.5 47.9 51.5 120 77

tc: coating thickness.

0

20

40

60

80

50 100 150 200 250 300

Cri

tical

Loa

d(N

)

Coating fracture energy(J/m^2)

Interface delamination Coating crack

Fig. 8. Effect of coating fracture energy on the critical loads of coating cracking and inter-face delamination from scratch simulations of diamond-coated cutting tool.

40

60

80

100

itica

l loa

d(N

)

Interface delamination Coating crack

6 P. Lu et al. / Surface & Coatings Technology xxx (2014) xxx–xxx

signals (Spot 2) exists at the load around 35.8 N, followed by a series ofcontinuous high-amplitude AE peaks. Both of tangential force and AEsignal evidences imply that the diamond coating delaminated at a normalload of 35.8 N. The scratch groove was observed by the microscope aftertesting. Fig. 5(a) shows the microscopic image around Spot 2, which hasa corresponding normal load of about 35.8 N. It can be noted that aroundthis location, coating delaminations have propagated, with the substratelayer exposed. Fig. 5(b) displays the microscopic image at the end ofthis particular scratch test, and it shows continuous coating delaminationwith a comparable width at the end of the scratch.

The averaged critical loads for coating cracking and interface delam-inations from multiple tests are 22 N and 36.1 N, respectively. Scratchtests with Sample 2 (coating thickness of 25 μm) were conducted in asimilar manner and the methods and approaches discussed above toidentify the critical loads were followed. It was found that the averagedcritical loads for coating cracking and interface delaminations frommultiple tests are 22.5 N and 47.9 N, respectively. Table 1 lists thesummary of the scratch testing results with the critical loads compared.It is noted that the critical loads for coating cracking are similar betweenthe two samples with different coating thicknesses.

The critical loads from scratch testing resultswere used to determinethe fracture energy values, for both coating cracking and interfacedelaminations, in the FE scratch simulations. A progressive loadingmethod with maximum load of 80 N was applied during the scratchprocess, followed by an unloading process. Combinations of variouscoating fracture energy and interface fracture energy levels were testedin the scratch simulations to compare the results from the scratchtesting. For Sample 1, the coating fracture energy and the interfacefracture energy are 140 J/m2 and 192 J/m2, respectively, which resultedin the critical load values close to the experimental results. Fig. 6 showsthe maximum principal stress contour of the coating–substrate modelwhen cracking was initiated in the diamond coating at a normal forceof 20.3 N. It can be noted that the corresponding maximum principalstress has just exceeded the critical strength of the diamond coating,8.14 GPa.

Fig. 7 shows the normal stress contour of the cohesive-zone elementsat the interface, also showing a coating delaminated area, at 3 locations:(a) delamination initiation, (b) end of scratch, and (c) completelyunloading. The coating delamination initiation occurred when theindenter scratched over a distance of 230 μm from the scratch startingpoint (corresponding to 36.7 N normal force as the delamination criticalload). In addition, the coating delamination propagates outwards in thetransverse direction along the scratch and the delamination size in-creased slightly along the scratch direction upon unloading, Fig. 7(c).Using the approach above, the coating fracture energy and interfacefracture energy of Sample 2 were also determined, with the results

Table 3Parameters and corresponding ranges of values used in the parametric study of scratchsimulations.

Parameter Low level Median level High level

Coating fracture energy (J/m2) 100 200 300Interface fracture energy (J/m2) 60 140 220Coating Young's modulus (GPa) 600 900 1200

Please cite this article as: P. Lu, et al., Surf. Coat. Technol. (2014), http://dx

listed in Table 2. The coating fracture energy and the interface fractureenergy of Sample 2 tested are 120 J/m2 vs. 140 J/m2 of Sample 1 and77 J/m2 vs. 192 J/m2 of Sample 1, respectively.

The developed FE scratch simulation was employed to investigatethe effect of the coating fracture energy, the interface fracture energy,and the coating Young's modulus on coating cracking and delamina-tions with different combinations listed in Table 3. The fracture energyof CVD diamond coating was in the range of 100 J/m2 to 300 J/m2, theinterface fracture energy was from 60 J/m2 to 220 J/m2, and the coatingYoung's modulus ranges from 600 to 1200 GPa. Further, the diamondcoating thicknesses was kept constant as 15 μm. The results of criticalloads for coating cracking and interface delaminations were analyzed.

Fig. 8 plots the critical loads for coating cracking and interfacedelaminations at different coating fracture energy levels. The coatingYoung's modulus is 900 GPa, the interface fracture energy is 140 J/m2,and the coating thickness is 15 μm. The critical load for coating crackingincreases almost linearly, from 9.3 N to 21.3 N, with the increase of thecoating fracture energy, from 100 J/m2 to 300 J/m2. This is intuitive sincethe coating fracture energy is directly related to the critical load forcoating cracking initiations. However, it can also be noted that thecritical load for interface delaminations decreases slightly from 56.7 Nto 49.8 N, if the coating fracture energy increases from 100 J/m2 to300 J/m2. Generally, with the increase of coating fracture energy, thecoating will become more difficult to crack, less external energy willbe dissipated for crack initiations and propagations, and thus, moreexternal energy will be dissipated for coating delaminations, whichwill reduce the delamination critical load, also increase the coatingdelamination size. However, since coating cracking occurs only in alimited area, the effect on delaminations seems to be limited.

0

20

50 100 150 200 250

Cr

Interface fracture energy(J/m^2)

Fig. 9. Effect of interface fracture energy on the critical loads of coating cracking and inter-face delamination from scratch simulations of diamond-coated cutting tool.

.doi.org/10.1016/j.surfcoat.2014.08.080

0

20

40

60

500 700 900 1100 1300

Cri

tical

load

(N)

Coating Young's modulus(GPa)

Interface delamination Coating crack

Fig. 10. Critical load vs. Young's modulus of diamond coatings for coating cracking andinterface delamination from scratch simulations.

7P. Lu et al. / Surface & Coatings Technology xxx (2014) xxx–xxx

Fig. 9 plots the results of critical loads for coating crack and interfacedelamination under different interface fracture energy levels. The coat-ing Young's modulus is 900 GPa, the coating fracture energy is 200 J/m2,and the coating thickness is 15 μm. For the interface fracture energy inthe range of 60 to 220 J/m2, the critical load for coating crack initiationsis about 15.5 N. Thus, the interface fracture energy has a negligible effecton the critical load for coating crack initiation, since coating fracturegenerally occurs before coating delamination.

Fig. 10 plots the results of the critical loads for coating cracking andinterface delaminations for varied coating Young's moduli. The coatingfracture energy is 200 J/m2, the interface fracture energy is 140 J/m2,and the coating thickness is 15 μm. The critical load for coating crackingdecreases with the increase of the Young's modulus, approximately9.2 N, for the coating Young's modulus increased from 600 GPa to1200 GPa. In contrast, the critical load for interface delaminationsincreases with the increase of the coating Young's modulus, which isconsistent with a previous study [10]. Hence, in diamond coatingdesigns, a smaller Young's modulus may improve the coating crackingresistance, but simultaneously may sacrifice the resistance for coatingdelaminations.

It must be pointed out that the current study has limited measure-ments in the experiments, which may not be sufficient for thoroughmodel verifications; more experimental work is strongly needed inorder to achieve quantitative comparisons to the models.

5. Conclusions

In this study, micro-scratch tests were conducted on diamond-coated carbide substrates to investigate the interface adhesion andfractures of diamond coatings. In addition, a 3D scratch simulation FEmodel with XFEM embedded and cohesive-zone interface incorporatedwas also developed with the fracture energy values, for both coatingand interface, extracted from the results of critical loads from scratchtesting. A systematic study was carried out to investigate the effects

Please cite this article as: P. Lu, et al., Surf. Coat. Technol. (2014), http://dx

of the coating fracture energy, the interface fracture energy and thecoating Young's modulus on the critical loads of coating cracking andinterface delaminations. The following are the summarized results:

(1) The developed 3D FE model may be applied to simulate coatingcracking and interface delamination behaviors during scratchtesting of a diamond-coated carbide substrate.

(2) From the FEmodel and scratch testing experiments, it is estimat-ed that the coating fracture energy of the samples tested in thisstudy ranges from 120 to 140 J/m2, and the interface fractureenergy ranges from 77 to 192 J/m2.

(3) The coating fracture energy has a negligible effect on the criticalload for interface delaminations, and similarly, the interface frac-ture energy does not affect the critical load for coating cracking,implying that the two failure modes are mostly unaffected byeach other for the testing range in this study.

(4) In addition, a higher coating Young's modulus will increase thecritical load for coating delaminations, but will reduce the criticalload for coating cracking.

Acknowledgments

This material is based upon work supported by the National ScienceFoundation under Grant No. CMMI 0928627.

References

[1] J. Karner, M. Pedrazzini, I. Reineck, M.E. Sjöstrand, E. Bergmann, Mater. Sci. Eng. AStruct. Mater. Prop. Microstruct. Process. 209 (1996) 405.

[2] E.J. Oles, A. Inspektor, C.E. Bauer, Diam. Relat. Mater. 5 (1996) 617.[3] J.M. Sanchez, S. El-Mansy, B. Sun, T. Scherban, N. Fang, D. Pantuso, W. Ford, M.R.

Elizalde, J.M. Martinez-Esnaola, A. Martin Meizoso, J. Gil-Sevillano, M. Fuentes, J.Maiz, Acta Mater. 47 (1999) 4405.

[4] H. Chai, Int. J. Fract. 119 (2003) 263.[5] R. Polini, P. D'Antonio, S. Lo Casto, V.F. Ruisi, E. Traversa, Surf. Coat. Technol. 123

(2000) 78.[6] A.J. Perry, Thin Solid Films 107 (1983) 167.[7] D. Barnes, S. Johnson, R. Snell, S. Best, J. Mech. Behav. Biomed. Mater. 6 (2012) 128.[8] R. Jaworski, L. Pawlowski, F. Roudet, S. Kozerski, F. Petit, Surf. Coat. Technol. 202

(2008) 2644.[9] J. von Stebut, F. Lapostolle, M. Bucsa, H. Vallen, Surf. Coat. Technol. 116–119 (1999)

160–171.[10] P. Lu, X. Xiao,M. Lukitsch, A. Sachdev, Y.K. Chou, Surf. Coat. Technol. 206 (2011) 1860.[11] F.X. Liu, F.Q. Yang, Y.F. Gao, W.H. Jiang, Y.F. Guan, P.D. Rack, O. Sergic, P.K. Liaw, Surf.

Coat. Technol. 203 (2009) 3480.[12] Y. Xie, H.M. Hawthorne, Surf. Coat. Technol. 155 (2002) 121.[13] X.P. Xu, A. Needleman, Solid State Phenom. 35–36 (1994) 287.[14] T. Belytschko, T. Black, Int. J. Numer. Methods Eng. 45 (1999) 601.[15] J. Melenk, I. Babuska, Comput. Methods Appl. Mech. Eng. 39 (1996) 289.[16] C. Daux, N. Moës, J. Dolbow, N. Sukumar, T. Belytschko, Intl. J. for Numer. Meth. in

Eng, 48 (2000) 1741.[17] A. Combescure, A. Gravouil, H. Maigre, J. Réthore, D. Grégoire, IUTAM Bookseries —

Symp. on Discret. Meth. for Evolv. Discontin, 2007, p. 185.[18] M. Bao, J. Yang, X. Liu, Y. Bao, in: Bos, Louter, Nijsse, Veer (Eds.), On Challenging

Glass 3: Conf. on Architec. and Struct. Applica, IOS Press, 2012, p. 387.[19] J. Hu, Y.K. Chou, R.G. Thompson, J. Burgess, S. Street, Surf. Coat. Technol. 202 (2007)

1113.[20] N. Moes, J. Dolbow, T. Belytschko, Intl. J. for Numer. Meth. in Eng., 46(1999), p. 131.[21] R.H. Telling, J.E. Field, Diam. Relat. Mater. 8 (1999) 850.[22] J. Hu, Y.K. Chou, R.G. Thompson, Surf. Coat. Technol. 203 (2008) 730.

.doi.org/10.1016/j.surfcoat.2014.08.080