failure mode maps in the thin film scratch adhesion

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Tribology Infrrnafional Vol. 30, No. 7, pp. 491498. 1997 0 1997 Elsevier Science Ltd

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re mode maps in the thin scratch adhesion test

The scratch test has been used to assess thin coating adhesion for some time now. In this test a diamond indenter is drawn across the coated surface under an increasing load (either stepwise or continuous) until at some load, termed the critical load, L,, a well-defined failure event occurs; if this failure event represents coating detachment then the critical load can be used as a qualitative measure of coating-substrate adhesion. However, it is well known that a range of possible failure modes can occur and only some of these are dependent on adhesion. Other failure modes which depend on plastic deformation and fracture within the coating, rather than any adhesive failure at the coating substrate interface, may be just as useful in the assessment of coating quality particularly for tribological applications. In this paper the load regimes in which the main adhesion-related failure modes (spallation and buckling) occur as a function of coating thickness will be presented for thermally grown oxide and sputtered nitride coatings. The origin of the observed failure modes and the use of the scratch test to assess coating/substrate adhesion in a more quantitative fashion is discussed in the light of these observations. 0 1997 Elsevier Science Ltd.

Introduction

The scratch test has been used to give a measure of the adhesion of a range of coatings for some time now ‘-6. In the most common version of the test a diamond stylus is drawn across the coated surface under an increasing load until some well-defined failure occurs at a load which is called the critical load, L,. If this test is to be used to assess adhesion then this failure must occur as a result of coating detachment which is not always easy to identify.

The types of failure which are often observed in the scratch test depend critically on the properties of both substrate and coating. If the coating is very soft com- pared to the substrate, considerable plastic deformation will occur within it and the scratch test critical load may be defined as the load at which the coating is scraped off exposing the substrate 7. However, it is not always easy to determine when this has occurred and quantification of the failure mode is difficult. For a hard coating on a softer substrate spallation and

Materiuls Division, Department oj. Mechanical, Materials and Manu- facturivg Erzgineering, Herschel Building Universify of Newcastle, Newcn~tle-llpon-T?irle NE1 7RlJ. UK

buckling failure modes result from interfacial detachment *x9 but a range of other cracks and deformed regions can be observed. Both the spallation and buckling failure modes are amenable to quantification and are discussed in some detail in this paper. For hard coatings on hard substrates the chipping observed in the scratch test is almost identical to the lateral fracture observed in the scratch testing of bulk ceramics. This failure is occasionally observed to coincide with the coating-substrate interface but this is not always the case making the results of the test difficult to interpret. Thus if scratch testing is to be used for adhesion assessment only the spallation and buckling failure modes are really useful. This generally limits the test to the assessment of hard coatings on softer substrates.

In this paper the scratch test failure modes observed for hard coatings (sputtered TIN coatings and thermally grown alumina oxide scales) are briefly reviewed. The theoretical basis for analysing buckle and spa11 failures is introduced and some observations about failure initiation and how the failure mode affects the locus of failure will be made.

Tribology International Volume 30 Number 7 1997 491

Thin film scratch adhesion test: S. J. Bull

Scratch test failure modes

The failure modes in the scratch testing of hard coatings can broadly be split into three categories:

1. Through-thickness cracking (Fig l)-including tensile cracking behind the indenter ‘.lO, conformal cracking as the coating is bent into the scratch track ‘,I’, and Hertzian cracking 8.

2. Spallation (Fig ‘)-including compressive spallation ahead of the indenter ‘,l”, buckling spallation ahead of the indenter ’ or elastic recovery induced spallation behind the indenter ‘2’ ‘.

3. Chipping in the coating (akin to lateral cracking in bulk ceramics).

The type of failure which is observed for a given coating-substrate system depends on the test load, the coating thickness, the residual stress in the coating and the properties of the substrate (e.g. hardness) as well as on test parameters such as indenter radius and sliding speed. Generally the critical load at which a given failure mode first occurs, or occurs regularly along the scratch track is used as a method of coating adhesion assessment. Comparisons between different samples are only valid if the mechanism of failure is

the same, which needs careful post facto inspection to confirm in many cases.

If the scratch test is to be used for the assessment of coating-substrate adhesion then it is the adhesion- related failure modes which are most important. There is a well-defined range of coating and substrate properties where adhesion-related failures can be observed in the scratch test (Fig 3) and this restricts the range of coating-substrate systems in which the test can be used for adhesion assessment. In general it is most useful for hard coatings on soft or hard substrates where plastic deformation of the coating does not occur to any great extent. However, there is a tendency for the diamond stylus to wear during the test in cases where both substrate and coating are hard so the test has been most widely applied to hard coating-soft substrate systems. In such cases three main types of adhesion- related failures are observed in the scratch test:

1. Buckling-this failure mode is most common for thin coatings ( < 9 pm for an alumina oxide scale on an iron-based alloy substrate, < 12 pm for TIN on high speed steel). Failure occurs in response to the compressive stresses generated ahead of the moving indenter [Fig 4(a)-(d)]. Localised regions containing interfacial defects allow the coating to buckle in response to the stresses. Individual

(THROUGH-THICKNESS CRACKING

BRITTLE TENSILE CRACKING

HERTZ CRACKING

Fig. I Through-thickness cracking failure modes in the scratch test

492 Tribology International Volume 30 Number 7 1997

DUCTILE TENSILE CRACKING

CONFORMAL CRACKING


/ INTERFACIAL FAILURE 1

BUCKLE SPALLATION

WEDGING/SPALLATION

Fig. 2 Intelfacial fuilure modes in the scratch test

Major Scratch Test Failure Regimes

1 lti0”

/ Plastic Deformation / I \I

1

Substrate Hardness

,

Fig. 3 Map of the main scratch test failure modes in terms of substrate and coating hardness

buckles may then spread laterally by the propagation of an interfacial crack. Spallation results when through-thickness cracks form in regions of high tensile stress within the coating [Fig 4(c)]. The presence of plastically piled-up material ahead Or the indenter can enhance this failure mode. Cnce the buckle has occurred the scratch stylus passes over the failed region crushing the coating

RECOVERY SPALLATION

into the surface of the scratch track formed in the substrate. At this time there may be enhanced coating removal or the failure can disappear completely depending on its size and the toughness of the coating. Buckling failures typically appear as curved cracks extending to the edges of or beyond the scratch track. Regions of spallation associated with buckle failures have edges perpendicular to the coating-substrate interface.

2. Wedge spallation. Once a critical coating thickness is achieved the coating becomes too stiff to buckle and reduce the stresses ahead of the indenter. Compressive shear cracks now form through the thickness of the coating before interfacial failure is observed [Fig 4(e)-(h)]. These cracks have slop- ing sides [Fig 4(f)] and the continued forward motion of the stylus drives wedges of adjacent coating under the segment bounded by the shear crack causing decohesion of the interface. Large enough displacements will cause a region ahead of the indenter to spall. In extreme cases the scratch diamond can drop into the hole left by removal of the coating and there is a dramatic increase in scratch width and scratch depth.

3. Recovery spallation-this failure mode is associated with the elastic recovery which occurs behind the stylus as it travels over the coated surface and depends on plastic deformation in the substrate and



Coating in Compression

Increasing applied stress

areas of decohesion

zone of athermal

directions of crack propagation e-

sites of tensile

damage interface

Fig. 4 Schematic of the main stages in adhesion-related coating failure mechanisms, (n)-(d) buckling and (e)- (h) wedging

through-thickness cracking in the coating. After the stylus passes and the scratched region is unloaded the elastic deformation jn the coating-substrate system is relaxed. However, due to the plastic deformation in the substrate, which results in the formation of the scratch track, it is not possible to completely relax the substrate elastic deformation and a residual stress remains. If through thickness cracking has occurred in the coating any residual strain on it can be more completely relaxed-tensile recovery stresses in the coating are converted into shear stresses at the coating- substrate interface near to these cracks. The propagation of interfacial shear cracks due to these stresses can lead to spallation either side of the scratch track. This failure mode is not generally observed for hard coatings on soft substrates where the adhesion is good and will not be discussed further in this paper.

The buckling and spallation failures which are observed for TiN coating and alumina oxide scales are discussed in some detail in the following sections.

Experimental

Samples of stainless steel (304) and the oxide dispersion strengthened alloy MA956 were cut into 20 x 10 x 2 mm sections, polished to a 1 mm diamond finish and cleaned and degreased prior to use. The 304 stainless steel coupons were coated with TIN by sputter ion plating I2 at a temperature of 500°C and a bias voltage of - 35 V. Thicknesses in the range 1-15 pm were deposited and measured by ball cratering. Prior to TIN deposition and samples were sputter cleaned

and a 120 nm titanium interlayer was deposited to promote adhesion. The MA956 samples were weighed and placed in alumina crucibles prior to isothermal oxidation in flowing laboratory air at 1150°C and 1250°C for times up to 1400 h. This produces oxide scale thicknesses up to 20 pm. Both samples and crucibles were weighed before and after exposure to pro- vide an estimate of scale thickness and the amount of spallation. It was confirmed that a weight gain of 1 mg/cm’ equates to a scale thickness of around 5 pm l3 by ball cratering.

Scratch testing was performed using a CSEM manual scratch tester fitted with a Rockwell “C” diamond (120” cone with a 200 pm radius hemispherical tip). This is a dead-loaded machine where a separate scratch is made for each applied load. A scratch length of 3 mm was used and loads were applied from 200 g upwards in 200 g increments. The tester is fitted with acoustic emission monitoring equipment which can detect emission in the vicinity of 100 kHz which was used as an on-line failure monitor. Buckle failures lead to a small increase in acoustic emission whereas wedge spallation leads to a much more dramatic increase. However, although acoustic emission can given an indication of failure mode, careful reflected light microscopy exam- ination was necessary to confirm this and determine the critical load.

In this sort of scratch test, care has to be taken in setting the critical load criterion since it is known that there is a distribution of flaws at the coating/substrate interface 6. In this study the load at which failures are first observed to occur regularly along the scratch track, correlating with an increase in acoustic emission, was used to avoid problems with isolated defects dominat- ing the results. The failure mode is somewhat subjec-


tive, but since the total number of wedge spallation failures was low, a full Weibull statistics analysis, as detailed in 6, was not possible.

Results

Failure modes

For both TIN and alumina scales two types of scratch test failure modes related to adhesion were observed (Fig 5). For thjn TIN coatings buckling failures were observed within the track [Fig 5(a)] but these extended outside the track for the thin oxide scales on MA956 [Fig 5(c)]. For thicker coatings, wedging failures were observed [Fig 5(b) and (d)]. The wedge is most appar- ent at the front of the spalled region where the compressive stresses on the coating are greatest. The sides and rear of the spalls are delineated by through-thickness cracking which is less obviously wedge-like. For the TIN coatings the transition to wedging spallation occur! at a slightly greater thickness than for the oxide scale on MA956 (Figs 6 and 7).

Failure load regimes and stresses

For all coatings investigated the critical load for buckle formation increases as the coating thickness increases (e.g. Figs 6 and 7). Wedging spallation does not occur until higher coating thicknesses and the critical load for wedging failure decreases as the coating thicknesses increases. In some cases (e.g. near the wedge/buckle transiGon) it is possible to see both types of failure mode on the same scratch track, but generally only one type of failure is observed.

The ,;tresses responsible for coating detachment, cr,, are a combination of the residual stress remaining in the coating at room temperature, crR, and the stresses introduced by the scratch stylus, cs. Thus

a, can be measured for both TIN coatings and the oxide scales on MA956 by X-ray diffraction using

FeCrAI and FeCrAlY: 1100°C Oxidation

500

0

0 2 4 6 8

Thickness (pm)

Fig. 5 Scanning electron micrographs of scratch test failure modes in TiN coatings: (a) buckling and (b) wedge spallation; and alumina oxide scales on MA9.56: (c) buckling and (d) wedge spallation

Thin film scratch adhesion test: S. .I. Bull

Sputter Ion Plated TiN

! 0 2 4 6 8 IO

Thickness (pm)

12 14 16

Fig. 6 Variation of the critical load for coating detach ment in the scratch test with coating thickness for sputter ion plated TiN on stainless steel

4Qxl

3500

3ooo

2500

2ooo

1500

IOIYJ

500

MA956

Thickness (urn)

Fig. 7 Vuriation of the critical load for coating detachment in the scratch test with scale thickness for an alumina oxide scale grown on MA956 at 1250°C and 1150°C

the well-known sin’ $ method ‘I. Table 1 tabulates measurements made in this study. The technique is most applicable to the thickest coatings due to the X- ray penetration depth which is greater than the coating thickness in all cases. No significant variation in residual stress as a function of coating thickness was observed in the films tested here.

However, it is much more difficult to determine the stresses introduced by the scratch diamond, us. No adequate theoretical model exists to predict the stresses generated ahead of a moving indenter in a bulk elastic- plastic material, let alone in a coated system 15. It is necessary to estimate the stresses using other methods.

In the scratch testing of TIN coatings on stainless steel or other steel substrates it has been shown that the



Table 1 Residual stress at room temperature in the coatings determined by X-ray diffraction (sin2 $ method)

Coating-substrate system Compressive residual stress

(GPa)

5 pm TIN on stainless steel 7 pm TiN on stainless steel

5 pm Al,03 scale on MA956

6.03 i 0.07 6.15 f 0.14

Formed at 1250°C 3.91 f 0.09 Formed at 1150°C 3.69 + 0.10

critical load for coating detachment is reduced as the residual stress in the coating increases 16. In this case the residual stress is changed by using different levels of ion bombardment (bias voltages) during deposition and can be measured by X-ray diffraction. Equating the change in critical load measured in the scratch test on fully dense TIN coatings with the difference in residual stress measured allows a calibration factor to be determined. In the case of the TIN films tested here: 1 g in the scratch test is equivalent to a 0.6 MPa compressive stress ahead of the indenter.

For MA956 oxidation experiments carried out at 1150°C and 1250°C produce alumina scales with a difference in residual stress of 224 MPa (Table 1). Scratch testing these scales (Fig 7) shows a difference in failure load of 590 g for wedge spallation failures (i.e. I g is equivalent to 0.38 MPa). The scales produced by oxidation at 1150°C require a higher scratch test load to produce failure than the scales produced by oxidation at 1250°C as expected from their lower residual thermal stress. There is no significant difference in failure load for the two test temperatures for the buckle failures which implies that the residual stress on the coating is much less important in this case.

Using the two calibration factors the failure stress/thickness behaviour for the two materials is shown in Fig 8. Considerably higher failure stresses are observed for the TiN coatings which are known to be very adherent.

Discussion

Much of the theoretical background for modelling of buckling and spallation failure modes has been developed for oxide scales under conditions of thermally induced stresses. According to Evans I7 the critical buckling stress a, is given by

where E, and u, are the Young’s Modulus and Pois- son’s ratio of the coating, f the coating thickness and R the radius of the circular area of interfacial detachment necessary to cause buckling. This predicts that the critical buckling stress should increase with coating thickness as is observed here. Using Equation (2) and the data in Fig 8 and Table 2, R can be calculated

--e MA956 Buckle -- MA956 Spa11

10

Thickness (Nm)

Fig. 8 Varintiorz of the compressive stress to initiate failure in the scratch test with coating thickness ,for TiN coatings and aluminn scales or1 MA956 grown nt 1250°C

Table2 Properties of TIN and alumina coatings used in the calculations

Material Young’s Poisson’s ratio modulus (GP)”

&OS 380 0.26 TiN 600 0.21

“Measured from phonon dispersion curves on bulk materials- values measured from scales or coatings by other techniques are often lower

and is plotted against the measured width of the scratch track in Fig 9 for both TiN and MA956. The corre- lation between the track width and the size of the defect responsible for buckling is very good. Changing

10 20 30 40 50 60 70 80 90

Calculated buckle radius (pm)

Fig. 9 Plots of calculated interjbciul defect radius against scratch truck width for TN contirzgs on stainless steel and alumina scales on MA956



the Young’s Modulus of the coating, which might be necessary to take into account the fact that the properties of scales and coatings are not the same as those of bulk materials, will alter the size of R but does not affect the relationship with the track width. It is thus likely that the plastic deformation in the substrate, which defines the scratch width at the failure loads observed here, is responsible for the initiation of buckling. in the region below the coating the constraint from the harder coating on the deformation of the softer substrate leads to the generation of high shear stresses and the propagation of an interfacial shear crack. It is this shear crack which provides the initial disbond which causes buckling. Since plastic deformation in the substrate is necessary to form the buckle- initiating defect this explains the well-known obser- vatior: that the critical load for coating detachment increases as the substrate hardness increases and the size of the scratch track at a given load is reduced Is.

The wedge spallation failure mode depends on two distinct processes occurring 17. Firstly a compressive shear crack must form in the coating and then interfacial detachment occurs. According to Evans I7 the biaxial stress necessary to cause the wedge crack, uW, is given by

(3)

where y is the fracture energy and A the width of the wedge spalled region. This is independent of the thickness of the coating. The biaxial stress to produce the spall, o,,, after shear cracking has occurred is given by

(41

where 3/~ is the interfacial fracture energy. The total failure stress is given by the sum of these two.

Clearly by plotting the measured failure stress, aF, against the reciprocal of the square root of coating thickness, the two contributions can be separated. This is shown in Fig 10 for the two types of coating. Table 3 gives values for yF and (T, for TIN and MA956. In fact :he biaxial stress assumption, whilst valid for the thermal and growth stresses, is not strictly valid for the scratch induced stresses which are closer to uniaxial. How-ver, in both cases investigated here the scratch induced stresses are much smaller than the biaxial therr_lal and growth stresses so this is a reasonable first approximation.

The interfacial fracture energy for MA956/alumina is reasonably low and represents predominantly brittle failure with little crack tip plasticity. In fact the observed value is similar to that determined for bulk alurrina from fracture toughness tests ( - 20 J/m” 19) but is higher than values calculated for alumina scales frorr tensile test data. For instance, Schutze 2o quotes values of scale fracture toughness between 0.4 and 1.3 MPam”* which give fracture energies between 0.2 and 2.2 J/m’ using yF = KlC2/2E. The discrepancy between the measured values for scales and bulk ceramics probably arises due to the existence of other crack energy absorption mechanisms in the bulk

250 250 260 260 270 270 280 280 290 290 300 300 310 310 320 320 330 330

UP (dR) UP (dR)

8.6

7.8

7.6

200 200 220 220 240 240 260 260 280 280 300 300 320 320 340 340

Fig. 10 Variation of failure stress with Id for (a) TiN coatings on stainless steel and (b) alumina scales on MA956

Table 3 Compressive failure stress and intetfacial fracture energy for TIN and MA956 determined from wedge spallation failures

Substrate- coating

Interfacial Failure stress, fracture energy, u, (GPa)

yF (J/m*)

MA956/AI,O, 28” 3.72 304rTiN 451 2.52

“Average value-the interfacial fracture energy varies with oxidation temperature.

materials (e.g. microcracking ahead of the crack tip) which are not active in the scales where the measured values are close to the surface energy of the material as expected. Careful microscopy shows that the failure crack propagates at or close to the interface for MA956 wedge spallation so some crack-tip plasticity is probably occurring.



The much higher interfacial fracture energy for TIN is representative of more ductile failure where more blunt- ing of the crack tip can occur by plasticity. The failure crack appears to propagate along the interface in this case since the fracture energy is between that expected for TIN and the substrate. This is confirmed by microscopy - it is not possible to detect any TIN at the bottom of wedge spalls, nor any substrate material attached to spalled debris by X-ray analysis in the scanning electron microscope.

The measured fracture stress for alumina on MA956 is close to the residual stress in the coating measured by X-ray diffraction showing that spontaneous shear cracking probably occurs on cooling but at low temperatures. Wedge spallation has been observed to occur on cooling alumina scales by 1070°C on MA956 pre- viously . 2’ This temperature drop generates to a compressive thermal stress of - 3.9 GPa in the absence of creep which is reduced to 3.7 GPa when high temperature creep is allowed to relax some of the substrate stresses. The fact that failure occurs at a low temperature means that the oxide on the bottom of spalled regions has not grown after spallation to disguise the fracture surface.

The calculated compressive fracture stress for TIN is smaller than the tensile fracture stress observed in bending experiments ( - 5GPa 22. after correction with an estimated value residual stress which may be in error) but is greater than the maximum tensile residual stress which is observed when TIN is deposited on low expansion substrates by sputter ion plating ( - 0.5 GPa 23). The value measured here thus looks reasonable for compressive failure. Since it is much lower than the residual stress in the coating it is expected that shear cracks form on cooling the coating- substrate system to room temperature after deposition. However, due to the ductile nature of the interface they do not propagate along it and cause spallation. This confirms the previous observations that hard ductile substrates give the best adhesion for TIN coatings *.

Conclusions

Scratch testing can give useful information about the adhesion of coatings provided that careful identification of the failure modes is carried out. Buckling failure modes predominate for thin flexible hard coatings on a ductile substrate. In these cases plastic deformation of the substrate leads to an interfacial defect which initiates failure. Wedge spallation occurs for thicker, stiffer coatings. To achieve this a compressive shear crack forms through the thickness of the coating and propagates down to the interface. Material can then spa11 as the wedge lifts the coating away from the substrate.

To improve resistance to buckling it is necessary to reduce the size of interfacial defects and the extent of

plastic deformation immediately below the coating. This means that hard, ductile substrates are preferred. To reduce the susceptibility of the coating-substrate system to wedge spallation increasing the toughness of the coating to prevent shear crack propagation and ensuring that interfacial crack propagation actually occurs in a ductile fashion is required.

To derive maximum benefit from the scratch test better theoretical models for the stress fields generated by the moving indenter in coating-substrate systems are required.

Acknowledgements

This work was supported by the Corporate Research Programme of AEA Technology and builds on results obtained under BRITE/EURAM contract BE7972 on ‘How to improve the failure resistance of alumina scales on high temperature materials’.

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