wear behavior of thermally sprayed ceramic oxide coatings

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Wear 261 (2006) 1298–1315

Wear behaviour of thermally sprayed ceramic oxide coatings

Giovanni Bolelli, Valeria Cannillo, Luca Lusvarghi ∗, Tiziano ManfrediniDipartimento di Ingegneria dei Materiali e dell’Ambiente, Universita di Modena e Reggio Emilia, Via Vignolese 905, 41100 Modena, MO, Italy

Received 6 July 2005; received in revised form 9 February 2006; accepted 10 March 2006Available online 18 April 2006

bstract

The wear resistance of plasma-sprayed ceramic coatings (Al2O3, Al2O3–13%TiO2, Cr2O3) has been investigated through pin-on-disk and dryand-steel wheel tests, has been correlated to microstructural and micromechanical characteristics (microhardness, fracture toughness) and has beenompared to well-known platings (such as Cr electroplating and electroless Ni) and HVOF-sprayed cermets (WC–17%Co, WC–10%Co–4%Cr).lasma-sprayed ceramics are hard but brittle: dry particles abrasion occurs through splats detachment. The toughest coating (Al2O3) displays

he highest wear resistance, which in fact overcomes HVOF-sprayed cermets and Cr electroplating, when a low number of wheel revolutionsre considered. In pin-on-disk tests, no coating undergoes wear loss against the 100Cr6 ball, that possess lower hardness. Against the alumina
all, Al2O3 and Al2O3–TiO2 coatings show high wear rates and friction coefficients (due to chemical affinity), while Cr2O3 possesses better wearesistance, lower friction coefficient and inflicts less wear on the counterpart. Cr2O3 wear scar consists in plastically deformed splats and debrisorming a quite adherent protective tribofilm.
2006 Elsevier B.V. All rights reserved.

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eywords: Sliding wear; Abrasive wear; Plasma spraying; Ceramic coatings; F

. Introduction

Thermal spraying is often considered as a potential alternativeo traditional coating manufacturing techniques (such as hardhrome electroplating) for the production of wear-resistant coat-ngs [1–4]. In fact, a large variety of hard materials (includingeramics and cermets) can be deposited on a cold (or mod-rately pre-heated) substrate [5,6], thus obtaining very hardoatings while preventing thermal alteration of the substratetself (which invariably occurs in other hardfacing processes,uch as in welding processes), which is a key requirement whenesign tolerances must be satisfied, thin-walled components areeing considered, or heat-sensitive materials (like Al and Mglloys) are being processed.

Among the various thermal spraying techniques, plasma-praying and HVOF-spraying are the most suitable for theroduction of high-quality wear resistant coatings and they are
oth technologically mature processes. HVOF was developed tovercome plasma-spraying limits: it is well known that HVOF-praying gives far superior results than plasma-spraying when
∗ Corresponding author. Fax: +39 0592056243.E-mail address: [email protected] (L. Lusvarghi).

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043-1648/$ – see front matter © 2006 Elsevier B.V. All rights reserved.oi:10.1016/j.wear.2006.03.023

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anufacturing cermet coatings, because of the much higheras jet velocity and lower flame temperature resulting in coat-ngs with extremely low porosity, low splats oxidation and lowarbide decomposition and/or dissolution [7]; in fact, manyxperimental work concerning HVOF-sprayed cermet coatingsested under various conditions exist [8–13]. However, HVOFas its own limits: it produces high quality metallic alloy andermet coatings, but these powders are difficult to prepare andery expensive. HVOF sprayed ceramic oxide coatings are stillnder investigation, but they are not widely used yet, becausehey must be tested and can be deposited just with few ofhe commercial torches [14,15]. Thus, plasma spraying is stillhe most widespread production technique for ceramic coat-ngs, like Al2O3, Cr2O3. Such coatings are more porous andrittle than HVOF-sprayed cermets, because of intrinsic poros-ty of plasma-sprayed coatings due to lower particles in-flightelocity and quenching-induced microcracking in the ceramicplats [6,16]. Anyway, they possess very high hardness, dueo their purely ceramic nature, they are almost insensitive to

any corrosive environments and can stand high temperatures
17]. Thus, if the involved application does not require a liquidight coating, but just wear resistance, plasma-sprayed ceramicxide coatings can be a good solution. Besides, manufactur-ng of atmospheric plasma-sprayed ceramic coatings can be less
mailto:[email protected]

dx.doi.org/10.1016/j.wear.2006.03.023

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xpensive than HVOF-sprayed cermets in many applications,ince powder and processing costs are often lower [18,19], evenhough exceptions may exist and deposition efficiency is also

concern. Therefore, many industrial processes make use oflasma-sprayed ceramic coatings, whose reproducibility is goodnce the optimal set of parameters has been found. For example,he food and medicine packaging industry does not only needear resistance, but also the absence of heavy metals contami-ation: Al2O3 and Al2O3–TiO2 are often used for this reason inhat field. Therefore, plasma-sprayed hard ceramic coatings aretill studied nowadays [20–22].

Therefore, a thorough study of the wear resistance of ther-ally sprayed coatings must involve plasma-sprayed ceramics,hich could represent an economical alternative to HVOF-

prayed cermets in some industrial applications. Much researchelated to the basic wear mechanisms of plasma-sprayed oxidesxists, since such coatings have been studied for a long time23–27]; however, there exist a few works comparing them tohe characteristics of other thermally sprayed coatings as wells to other industrially widespread wear resistant coatings, suchs hard chrome electroplating and nickel electroless plating28,29]. Furthermore, to fully assess the industrial applicabil-ty of thermally sprayed coatings, and of plasma sprayed oxidesn particular, wear maps should be experimentally obtained, as its currently being done for massive sintered ceramics [30–32].herefore, the aim of this study is both to provide an experi-ental assessment of the wear rates, wear mechanisms, friction

oefficients of various plasma-sprayed ceramic coatings underifferent sliding wear conditions through pin-on-disk testing,s a first step towards wear mapping of such materials, and toompare the performance of plasma-sprayed ceramic coatings,VOF-sprayed cermets and metal platings under various wear

onditions, in particular under dry sliding conditions and underry particles abrasion conditions.

. Materials and characterizations

Three plasma-sprayed ceramic coatings, namely Al2O3powder: Sulzer-Metco 105SFP, −31 + 3.5 �m), Al2O3–3%TiO2 (powder: Sulzer-Metco 130, −53 + 15 �m) and Cr2O3powder: Saint-Gobain #3033, −15 + 5 �m), with a NiC-CrAlY (powder: Sulzer-Metco 461NS, −150 + 22 �m) bond
oat to improve adhesion, and two HVOF-sprayed cermetnes, namely WC–17%Co (powder: Tafa 1343, −45 + 15 �m,gglomerated and sintered) and WC–10%Co–4%Cr (pow-er: Sulzer Metco 5847, −53 + 11 �m, agglomerated and
iwha

able 1lasma spraying operating parameters

arameters Al2O3 Al2O3–

pray distance (mm) 105 105ooling gas type and pressure Ar, 8 bar Ar, 8 balasma gases flow rates Ar: 50 slpm H2: 15 slpm Ar: 50 surrent (A) × voltage (V) = power (kW) 580 × 68 = 39.44 690 × 6eeding disk revolution speed (rpm) 12 12arrier gas type and flow Ar, 3.5 slpm Ar, 3.5 s

(2006) 1298–1315 1299

intered), have been manufactured onto C40 steel plates100 mm × 100 mm × 5 mm). The crystalline phases of all thehree ceramic oxide powders have been studied with X-raysiffraction (XRD, Philips PW3710, Cu K� radiation). The sub-trates were grit-blasted with 500 �m alumina grits (Sulzer-

etco Metcolite-C) in a vacuum-operated blasting machineNorblast) prior to coating deposition. Plasma-sprayed coatingsere manufactured with a Sulzer-Metco F4-MB plasma torch,perated in Air Plasma Spraying (APS) mode in a C.A.P.S.lant (Centro Sviluppo Materiali S.p.A., Roma, Italy, co-sharedith Universita la Sapienza, Roma, Italy). The spraying param-

ters for plasma-sprayed oxides are listed in Table 1, whilepraying runs of cermet coatings have been described else-here [29]. Hard chrome electroplating and Ni–14%P elec-

roless plating on C40 steel have also been studied for ref-rence. Both Cr electroplatings (on ground substrates) andi–P electroless platings (on micro-grit blasted substrates)ere industrially manufactured by proprietary processes. The

atter was supplied in the “as-plated” state, without thermalreatment: a heat treatment (5 ◦C/min heating, 400 ◦C treat-

ent temperature for 1h, slow cooling inside the kiln) waserformed on some of the samples (hereafter referred to asi–P tt). The treatment conditions were chosen following

iterature indications in order to achieve the highest possi-le hardness increase through precipitation of nanostructuredi–P crystals [33]. The plasma sprayed coatings microstruc-

ure was determined by XRD and by scanning electronicicroscopy (SEM, Philips XL30) on polished cross-sections

mounted in resin, ground with 400, 800, 1000, 2000 meshiC papers and polished with 3 and 0.5 �m polycrystallineiamond paste). Image analysis was also performed on SEMmages to determine coating porosity (UTHSCSA Image Tool. 3.0 software). Roughness measurement was performed byptical profilometry (ConScan profilometer, CSM Instruments,witzerland), determining the Ra and Rz parameters (UNISO 4287-1). A depth-sensing Vickers microindenter (CSMnstruments, Switzerland) was employed to measure Vickersicrohardness, fracture toughness and elastic modulus on coat-

ngs polished cross-sections. Vickers microhardness was cal-ulated on all coatings by measuring the indentations diag-nals (1 N load, 15 s loading time) from SEM micrographs.racture toughness was determined on thermally sprayed coat-

ngs by high-load, cracked Vickers microindentations producedith 10 N load (5 N for Al2O3–13%TiO2 due to its lowerardness and toughness), measuring the indentation diagonalsnd crack lengths from SEM micrographs and employing the

13%TiO2 Cr2O3 Bond coat

105 105r Ar, 8 bar Ar, 8 barlpm H2: 13 slpm Ar: 45 slpm H2: 15 slpm Ar: 50 slpm H2: 15 slpm5 = 44.85 650 × 67 = 43.55 560 × 69 = 38.64

15 10lpm Ar, 2.5 slpm Ar, 2.7 slpm

1300 G. Bolelli et al. / Wear 261 (2006) 1298–1315

Table 2pin on disk testing conditions

Test 1 Test 2 Test 3 Test 4 Test 5 Test 6 Test 7 Test 8

Sliding speed (m/s) 0.1 0.1 0.1 0.1 0.2 0.2 0.2 0.2Sliding distance (m) 250 250 250 250 250 250 250 250Normal load (N) 1 1 5 5 1 1 5 5P AE 38ν 0.

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in material 100Cr6 Alumina 100Cr6(GPa) 210 380 210

0.3 0.23 0.3

vans–Wilshaw formula [34]:

IC = 0.079(P/a3/2) log(4.5a/c) (2.1)

here KIC = MPa m0.5, a the half diagonal of the indentation�m), c the crack length (�m), and P is the load (mN). Aseported in literature [35], the limitation for the use of thisormula is that the ratio between the crack and the half diagonalength must be between 0.6 and 4.5. This condition has beenerified in all the tests. Even if the formula has been developedor “halfpenny-shaped” cracks, it has been demonstrated thatt is valid also for Palmqvist cracks [36]. The employmentf this formula can already be found in literature for plasmapray ceramic coatings [37] and HVOF cermets [38]. Elasticodulus was determined on all coatings from the unloading

art of instrumented indentation loading-unloading curves, byhe Oliver–Pharr formula [39]. The Poisson’s ratio (requiredor elastic modulus calculation) was assumed to be 0.23or all ceramics and 0.3 for metals (Cr, Ni–P) and cermets.ollowing preliminary tests (which highlighted some effect of

he indentation load on the elastic modulus value, especially forlasma-sprayed ceramics), a 5 N indentation load was chosen,ith 5 N/min loading rate, 4 N/min unloading rate, 15 s loading

ime. A minimum of 20 indentations was performed for eachardness, toughness and elastic modulus measurement. Dryarticles abrasion resistance was tested on all thermally sprayed
oatings in their “as-sprayed” conditions and on “as-plated”ard chrome for reference (the measurement was not performedn Ni plating due to its low thickness causing the substrate to beuickly exposed) by a simple dry sand-steel wheel test, using
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able 3in on disk testing plan (EHC = electrolytic hard chrome)

Test 1 Test 2 Test 3

s-sprayed Al2O3 × × ×s-sprayed Al2O3–TiO2 × × ×s-sprayed Cr2O3 × × ×s-sprayed WC–Co ×s-sprayed WC–Co–Cr ×s-plated EHC ×s-plated Ni–P ×s-plated Ni–P tt ×olished Al2O3

olished Al2O3–TiO2

olished Cr2O3

olishedC–Co

olished EHC

lumina 100Cr6 Alumina 100Cr6 Alumina0 210 380 210 380

23 0.3 0.23 0.3 0.23

200.1 mm diameter Fe360A steel wheel rotating at 75 rpm,EPA 80 alumina grains (180 �m average particle diameter)s abrasive medium with a 1 g/lap mass flux, and applying a0.2 N normal load (Ceramic Instruments AP/87 abrasimeter).his test is a modified version of ASTM G65, with the mainifferences consisting in the use of steel wheel instead of aubber wheel and of corundum in place of Ottawa sand.

Ball-on-disk dry sliding tests were performed with a pin-on-isk tribometer (CSM Instruments, Switzerland) using 100Cr6alls (manufacturer’s nominal hardness: 7 GPa; radius = 3mm)nd sintered alumina balls (manufacturer’s nominal hardness:9 GPa; radius = 3 mm) on (22 mm × 22 mm × 5 mm) coatedlates obtained by cutting the (100 mm × 100 mm × 5 mm)lates. Plates are fixed onto a rotating disk: the ball is fixednto a steady ball holder pressed against the sample surface withnormal load (rotating unidirectional sliding). Eight different

arameter sets have employed, varying the sliding speed, theormal load and the counterpart material. The parameters of thearious tests (labelled tests 1–8) are listed in Table 2. For eachest, the friction coefficient was measured on-line by the instru-

ent; the wear rate of the ball was determined by measuringhe worn cap diameter with an optical microscope and the wearate of the sample was determined by measuring the area ofhe wear track cross-section by optical profilometry (ConScanrofilometer, CSM Instruments, Switzerland: each area value ishe average of four measurements) and calculating the resulting
ear volume. All tests (from tests 1 to 8) were performed onlasma-sprayed ceramic coatings (Table 3): these data will pro-ide important information on wear mechanisms and will serves a basis for wear mapping of such coatings. A reduced set of
Test 4 Test 5 Test 6 Test 7 Test 8

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ests was performed on cermets and on other platings (namelyest 3, 4, 7, 8; Table 3): they provide a good amount of data tovaluate the cermets sliding behaviour and are of use as refer-nce values to compare the different performances. To assess theffect of surface finishing on the friction and wear behaviour,ests 7 and 8 have also been performed on polished plasma-prayed ceramics, HVOF-sprayed WC–17%Co and electrolyticard chrome (Table 3). These tribological samples were, in fact,aboratory polished using diamond abrasive papers with pro-ressively smaller diamond particles, down to 2 �m.

. Results

.1. Microstructure and micromechanical properties

SEM micrographs in Fig. 1A–F describe the microstructuresf plasma-sprayed ceramic coatings, while Cr electroplating andVOF-sprayed cermets chemical and microstructural charac-

eristics have already been described elsewhere [29]. Obvious
icrostructural differences appear between the various plasma-
prayed ceramic coatings. Cr2O3 markedly exhibits a lamel-ar microstructure with a prevalence of elongated interlamellarores; while almost no unmolten particles are found (Fig. 1A).

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ig. 1. SEM micrographs of coatings cross-sections. (A and B) Plasma-sprayed Cr2

ond coat; (E and F) plasma-sprayed Al2O3–13%TiO2 + NiCoCrAlY bond coat.

(2006) 1298–1315 1301

he overall coating porosity from image analysis is quite lowabout 6%): the employment of powder particles with smallverage grain diameter reduces the average size of pores dueo splats stacking faults, gas entrapment and unmolten parti-les. In Al2O3 (Fig. 1B), the interlamellar cohesion appearsigher. The porosity is still about 6%, due to more roundedores produced by unmolten particles, splats stacking faultsnd gas entrapment. In Al2O3–TiO2, TiO2 was well melted andartly mixed with alumina, since SEM micrographs highlightamellar areas containing different amounts of titania. There-ore, intersplat adhesion is better than the former coatings, butany vertical microcracks are present. Unmolten alumina par-

icles can also be found due to the coarser average particle sizef this powder; they cause splat-stacking faults. Thus, due tohese defects, the porosity of this coating is the highest amonghe presently tested ceramic coatings (about 9%), notwithstand-ng the better intersplat adhesion. The XRD qualitative analysisf the starting powders showed that the chromia powder fullyonsisted in eskolaite phase (Cr2O3), alumina powder in �-
l2O3 and alumina–titania powder in �-Al2O3 and anatase, i.e.
itanium dioxide low temperature phase. No glass phase wasresent in any powder. From XRD, the chromia coating con-ists in eskolaite; the alumina one mainly consists in �-Al2O3

O3 + NiCoCrAlY bond coat; (C and D) plasma-sprayed Al2O3 + NiCoCrAlY


Table 4Vickers microhardness, indentation fracture toughness, elastic modulus (Oliver–Pharr formula), Ra, Rz for all tested coatings

HV1N (GPa) KIC (MPa m1/2) E (GPa) Ra (�m) Ra (�m) ground Rz (�m) Rz (�m) ground

Al2O3 11.70 ± 1.70 2.57 ± 0.65 184 ± 6 5.00 ± 0.82 0.102 ± 0.050 18.33 ± 2.70 2.435 ± 0.547Al2O3–TiO2 8.18 ± 1.29 1.66 ± 0.45 141 ± 7 8.26 ± 1.29 0.105 ± 0.055 30.46 ± 4.71 2.967 ± 0.961Cr2O3 12.52 ± 1.26 1.67 ± 0.67 185 ± 13 3.07 ± 0.47 0.099 ± 0.047 11.60 ± 1.03 2.121 ± 0.348WC–Co 11.16 ± 0.41 4.02 ± 0.51 255 ± 9 3.25 ± 0.28 0.087 ± 0.036 12.75 ± 0.66 0.917 ± 0.389WC–Co–Cr 10.58 ± 0.46 3.76 ± 0.46 257 ± 16 3.41 ± 0.34 – 12.60 ± 1.50 –Hard Cr 8.42 ± 0.20 Not measurable

(ill-shapedmicrocracks) [29]

188 ± 8 0.518 ± 0.113 0.080 ± 0.033 3.30 ± 0.87 0.775 ± 0.286

Ni–P 5.52 ± 0.62 Not measurable 224 ± 21 2.77 ± 0.24 – 15.36 ± 1.31 –

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(coating too thin)i–P tt 9.17 ± 0.09 Not measurable

(coating too thin)184 ± 6

due to splats quenching), with minor quantities of �-Al2O3 andlassy phase; the alumina–titania consists mainly in �-Al2O3ith some �-Al2O3, glassy phase (Fig. 2, see black arrow)

nd a minor amount of rutile. The very low amount of crys-alline TiO2 indicates that it mostly dissolves in molten Al2O3,s SEM micrographs suggested, contributing to the formationf a low melting point glassy phase, which is probably respon-ible for the better intersplat cohesion. The hardness, fractureoughness, elastic moduli and roughness parameters of the testedoatings are listed in Table 4. It should be noticed that plasma-prayed ceramic coatings have different roughness: roughnesseems to be increasing with increasing powder average particleize. The roughness of plasma-sprayed Cr2O3, HVOF-sprayedermets and electroless Ni–P plating is quite similar, while elec-rolytic hard chrome possesses much lower roughness [29].r2O3 is the hardest of all tested coatings; Al2O3 has similarardness to WC–Co, while WC–Co-Cr is slightly less hard.he microhardness of Al2O3 is lower than that of bulk alu-ina (HV = 20.45 GPa [40]), both because of the intrinsically

ower hardness of �-Al2O3 than �-Al2O3 and because the inden-ation response of a plasma-sprayed material is governed notnly by the intrinsic hardness of the material, but also by the
amellar microstructure, with splat boundaries giving off underoad to facilitate the indenter accommodation [41]. Electroplatedhrome is less hard than all thermally sprayed coatings except forl2O3–TiO2. Ni–P has poor hardness in the as-deposited con-
ig. 2. X-ray pattern of alumina–titania plasma spray coating. The black arrowndicate the broad band caused by the presence of a glassy phase. The labels arehe following: � = �-Al2O3, � = �-Al2O3, � = �-Al2O3, R = rutile (TiO2).

ifppiotstrcpoaianvch

e as Ni–P – Same as Ni–P –

ition, but is definitely improved after the thermal treatment,ue to the two-phase nanocrystalline structure. Al2O3 possesseshe best fracture toughness among plasma-sprayed ceramics,hile Al2O3–TiO2 and Cr2O3 are less tough. Cermets are much

ougher than plasma-sprayed ceramics. The effect of the ther-al treatment on electroless Ni is immediately obvious from

he huge microhardness increase. Cracked Vickers microinden-ations are shown in Fig. 3A–C for plasma-sprayed ceramics:t can be seen that, for alumina and chromia thermally sprayedoatings, crack preferentially propagate parallel to the substratenterface, along splat boundaries. This is particularly evident forr2O3, where SEM micrographs had already highlighted a low

ntersplat cohesion. The abovementioned phenomenon is typi-al also of HVOF coatings. The alumina–titania coating seemso be the most isotropic among all tested coatings, with simi-ar microcracks propagating both parallel and transverse to theubstrate (Fig. 3C).

.2. Dry particles abrasion resistance

The dry sand-steel wheel test results are listed in Table 5.oth wear volume (mm3) and wear rate (mm3/Nm) for increas-

ng number of steel wheel revolutions have been included. Whileor ceramic coatings the wear rate remains quite constant inde-endently of the number of disk revolutions, for hard chromelating and HVOF-sprayed cermets there is a clearly decreas-ng trend, which is better highlighted in Fig. 4. Comparing thebserved wear rates for the various coatings, it can be noticedhat electrolytic hard chrome and HVOF-sprayed WC–Co rankimilarly in this test, with WC–Co–Cr (the less tough amongested cermets) performing worse. At a high number of diskevolutions, such coatings behave better than plasma-sprayederamics, but at a low number of revolutions, oxide ceramicserform definitely better, with the wear rate of the toughestne, i.e. Al2O3, being much lower than that of Al2O3–TiO2nd Cr2O3. The wear scar on plasma-sprayed alumina is shownn Fig. 5, while Fig. 4 compares the wear rate of plasma-sprayedlumina and of HVOF-sprayed WC–Co cermet as a function of
umber of disk revolutions, as a confirmation to the above obser-ations. Wear scar analysis also indicates that plasma-sprayederamics (in particular Al2O3 and Cr2O3), due to their highardness, do not undergo significant plastic deformation-related

G. Bolelli et al. / Wear 261 (2006) 1298–1315 1303

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ig. 3. SEM micrographs of high-load, cracked Vickers microindentations uselasma-sprayed Al2O3 + bond coat; (C) plasma-sprayed Al2O3–13%TiO2 + bon

henomena (such as microcutting and microploughing): theirnly relevant wear mechanism is brittle fracture, and in partic-lar, cracks mostly propagate through splat boundaries, whichre the weakest link in the material, as former indentation testslearly showed.

.3. Pin-on-disk wear test

Hertzian maximum and average contact pressures, maximumub-superficial shear stress and maximum shear stress depth arell reported in Table 6; they have been computed from analyti-
al Hertz’s formulae, using the experimentally evaluated elasticoduli in Table 4 for the samples and assuming E = 210 GPa,= 0.3 for 100Cr6 steel, E = 380 GPa, ν = 0.23 for sintered alu-ina. Clearly, such stresses are valid in a static contact condition,
snin

able 5ear volumes (mm3) and wear rates (×10−3 mm3/Nm) recorded in dry sand-steel w

est/sample Al2O3 Al2O3–TiO2 Cr2O3

revolutionsWear volume 2.03 ± 0.24 Not measured

because the wearrate stays constant

Not measurbecause therate stays co

Wear rate 11.47 ± 1.38

0 revolutionsWear volume 10.46 ± 1.65 14.97 ± 0.88 16.88 ± 1.8Wear rate 10.37 ± 1.64 14.85 ± 0.87 16.74 ± 1.8



verage wear rate 10.14 ± 1.29 13.73 ± 1.47 14.93 ± 2.1

fracture toughness measurement. (A) Plasma-sprayed Cr2O3 + bond coat; (B)t;.

ut do not take into account additional stresses due to friction,hich make the contact more severe. It must be considered thatertzian stresses and pressures are only valid at the beginning ofhe test; in fact, as the test progresses, the pin gets progressivelyorn, the actual contact area increases and the contact pres-

ure consequently decreases. It can be noticed that differencesetween contact pressures and contact stresses in the variousoatings arise due to the different elastic moduli; however, inhe same test, they always remain the same order of magnitude,he highest shear stress differences being about 25% between

C–Co–Cr and Al2O3–TiO2. Most importantly, the maximum
hear stresses always occur within the coating, without any sig-ificant involvement of the substrate and the coating-substratenterface. Thus, even though tested coatings have different thick-esses, pin-on-disk test results can be compared. The pin-on-
heel abrasion test

WC–17%Co WC–10%Co–4%Cr Hard Cr

edwearnstant

4.27 ± 0.01 6.79 ± 0.07 3.81 ± 0.1524.19 ± 0.01 38.50 ± 0.38 21.63 ± 0.82

7 8.44 ± 0.02 13.48 ± 0.04 7.52 ± 0.225 8.37 ± 0.02 13.34 ± 0.04 7.46 ± 0.22

0 9.61 ± 0.01 15.30 ± 0.50 8.72 ± 0.175 6.36 ± 0.01 10.12 ± 0.33 5.76 ± 0.11

8 10.82 ± 0.27 16.70 ± 0.35 10.20 ± 0.138 5.37 ± 0.14 8.28 ± 0.17 5.06 ± 0.07

2 Not calculated Not calculated Not calculated


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ig. 4. Al2O3 and WC–Co wear rates in dry sand-steel wheel tests as a functionf number of disks revolutions.

isk results for plasma-sprayed coatings are shown in Fig. 6Asample wear rate), Fig. 6B (pin wear rate) and Fig. 6C (averageriction coefficients). Material build-up on the coatings surfaces recorded on all samples when tested against 100Cr6 due toteel debris transfer from pin to sample, while the pin undergoesignificant material loss. The occurrence of material build-up is
ndicated in Fig. 6A as a negative value of the wear rate. Ther2O3 coating causes much less wear on the steel pin (subse-uently undergoing much less material transfer from the pin) andisplays lower friction coefficients. The wear rate of the steel pin
tps(

able 6ertzian maximum and average contact pressures, maximum sub-superficial shear s

onfigurations

Maximum hertziancontact pressure (MPa)

Average hertzian cpressure (MPa)

l2O3 Test 2–6: 716.9 Test 2–6: 477.9Test 4–8: 1225.8 Test 4–8: 817.2Test 1–5: 620.8 Test 1–5: 413.4Test 3–7: 1061.4 Test 3–7: 707.7

l2O3–TiO2 Test 2–6: 632.9 Test 2–6: 421.9Test 4–8: 1082.2 Test 4–8: 721.5Test 1–5: 560.5 Test 1–5: 373.7Test 3–7: 958.4 Test 3–7: 639.0

r2O3 Test 2–6: 718.6 Test 2–6: 479.1Test 4–8: 1228.8 Test 4–8: 819.2Test 1–5: 622.0 Test 1–5: 414.7Test 3–7: 1063.6 Test 3–7: 709.1

C–Co Test 4–8: 1430.4 Test 4–8: 953.6Test 3–7: 1198.5 Test 3–7: 799.0

C–Co–Cr Test 4–8: 1434.8 Test 4–8: 956.5Test 3–7: 1201.4 Test 3–7: 800.9

r Test 4–8: 1259.8 Test 4–8: 839.9Test 3–7: 1085.0 Test 3–7: 723.4

i–P Test 4–8: 1357.6 Test 4–8: 905.1Test 3–7: 1151.1 Test 3–7: 767.4

i–P tt Test 4–8: 1247.9 Test 4–8: 831.9Test 3–7: 1076.8 Test 3–7: 717.9

ig. 5. SEM micrograph of dry sand-steel wheel wear scar on plasma-sprayedl2O3. Black arrows indicate some particularly evident splats, proof of the brittleetachment of nearby splats.

gainst alumina and alumina–titania is 2–10 times higher thanhat against chromia in all cases (except for Al2O3–TiO2 in test); the wear rate against Cr2O3 never exceeds 10−5 mm3/Nm.he alumina pin undergoes lower wear rates (≈10−6 mm3/Nm)hen sliding against Cr2O3 as well. Overall, Cr2O3 coatingutperforms other plasma-sprayed ceramics in the pin-on-disk
est, concerning both wear rates (undergone and inflicted onins) and friction coefficients. The SEM micrographs of wearcars of plasma-sprayed coatings tested (test 7) against 100Cr6Fig. 7A for Al2O, Fig. 7B for Cr2O3) confirm that, during
tress, and depth of the point of maximum shear stress, for all pin-on-disk test

ontact Maximum hertzianshear stress (MPa)

Maximum shear stressdepth (�m)

Test 2–6: 222.2 Test 2–6: 12.4Test 4–8: 380.0 Test 4–8: 21.2Test 1–5: 192.4 Test 1–5: 13.3Test 3–7: 329.1 Test 3–7: 22.8



Test 4–8: 443.4 Test 4–8: 19.6Test 3–7: 371.5 Test 3–7: 21.4






F rates.d

tcffcipst(saO

l(itiBson

F

ig. 6. Pin-on-disk test results for plasma-sprayed coatings. (A) Coatings wearebris sticking; (B) pin wear rates; (C) friction coefficients.

he test, some steel debris is transferred to the coatings. Wheneramic coatings are tested against alumina pins, a tribofilm isormed (Fig. 8A–C), consisting in plastically deformed debrisrom both sample and pin, and also in plastically deformederamic splats (Fig. 9A–D). On Al2O3 and Al2O3–TiO2 coat-ngs (Fig. 8A and B), besides plastically deformed splats, lot oflastically deformed debris appears in the surface film, whicheems to possess low compactness (Fig. 9A and B). On Cr2O3,he surface film is much more compact and smooth in all cases
Fig. 8C): it mostly consists in plastically deformed debris andplats (Fig. 9C and D), with some alumina debris, which werelso plastically deformed and strongly embedded in this film.ptical micrographs of pin wear scars are in Fig. 10. They high-
taft

ig. 7. SEM micrographs of wear scars on plasma-sprayed Al2O3 (A) and Cr2O3 (B

Negative wear rates indicate material build-up on coating surface due to wear

ight that the steel sample mostly suffered two-body groovingploughing and cutting) but also some adhesive wear. Alumina,nstead, mostly undergoes brittle fracture; besides, some plas-ically deformed wear debris adheres to the pin, forming anrregular surface film, which undergoes grooving. Fig. 11A and

represent three-dimensional axonometric projections of wearcars on Al2O3 and Cr2O3 after test no. 8, reproduced from theptical profilometry results. While Al2O3 has undergone sig-ificant material loss, with pin penetration inside the coating,
he tribofilm on Cr2O3 is smoother than the original surfacend is placed slightly above the mean line of the unworn sur-ace. Table 7 sums up the wear rates of all tested samples inhe pin on disk wear test. HVOF-sprayed coatings exhibit an
) after test no. 7, indicating transferring of steel debris on the coating surface.


F 13%T

eaotaCtt

Nmae

FAs

ig. 8. SEM micrographs of wear scars on plasma-sprayed Al2O3 (A), Al2O3–

xcellent sliding wear resistance: they never undergo a measur-ble volume loss. Sometimes, material build-up is recorded fromptical profilometry. Their performance is therefore comparableo Cr2O3. Electrolytic hard chrome displays unfavourable wear
nd friction characteristics when compared to plasma-sprayedr2O3 and HVOF-sprayed cermets: when tested against sin-
ered alumina, it undergoes a significant wear rate (but lowerhan Al2O3 and Al2O3–TiO2). Remarkably, thermally treated

f

aH

ig. 9. Details of wear scars on plasma-sprayed ceramic coatings after testing agail2O3–13%TiO2, test no. 6; white arrows indicate plastically deformed splats; (C a

plats under the tribofilm.

iO2 (B) and Cr2O3 (C) after test no. 8, showing the formation of tribofilms.

i–P undergoes a negligible wear rate in all tests. Without ther-al treatment, it undergoes very high wear rates both against

lumina and against 100Cr6, being less hard than steel. How-ver, its wear rate becomes negligible in test no. 8. The reasons
or this behaviour will be discussed in the following section.
Fig. 12A and B respectively compare pin wear rates and aver-ge friction coefficients for plasma-sprayed ceramic coatings,VOF-sprayed cermets, Cr electroplating and Ni–P electroless

nst alumina pins. (A) Plasma-sprayed Al2O3, test no. 8; (B) plasma-sprayednd D) plasma-sprayed Cr2O3, test no. 8, with the appearance of undeformed


Fig. 10. Optical micrographs of pin wear scars. (A) Steel pin after test no. 3 against Cr2O3; (B) alumina pin after test no. 4 against Cr2O3.

Table 7Sample wear rates in mm3/Nm after pin on disk test on as-deposited (no grinding) coatings

Test 1 Test 2 Test 3 Test 4 Test 5 Test 6 Test 7 Test 8

Cr2O3 −2.07 × 10−4 1.35 × 10−4 −6.46 × 10−5 −4.91 × 10−5 −5.64 × 10−4 −2.48 × 10−4 −5.79 × 10−5 −5.12 × 10−5

Al2O3 −1.33 × 10−3 −4.98 × 10−4 −3.25 × 10−5 1.95 × 10−4 −1.42 × 10−4 1.12 × 10−3 −3.12 × 10−4 5.80 × 10−4

Al2O3–TiO2 −6.52 × 10−3 −1.02 × 10−3 −3.29 × 10−4 3.06 × 10−4 −1.33 × 10−3 −6.02 × 10−4 −2.36 × 10−4 4.72 × 10−4

Cr 0 0 4.11 × 10−5 0 0 1.90 × 10−4

WC–Co −6.56 × 10−4 −4.59 × 10−5 −1.45 × 10−4 −4.58 × 10−5

WC–Co–Cr −9.86 × 10−5 −2.96 × 10−5 −1.49 × 10−4 −5.18 × 10−5

N −4 .77 × −4 −4

N .17 ×N

poacClthH1wsncfbrAca

alhodcsccpHpwpta

i 7.14 × 10 8i-tt 0 2

ote: Negative values indicate material build-up on the sample.

latings. Alumina pins wear rates are higher by almost one orderf magnitude when tested against cermets than when testedgainst plasma-sprayed Cr2O3. Electrolytic hard chrome alsoauses higher friction coefficients and higher pin wear rates thanr2O3 when tested against alumina pins (especially with the

ower sliding speed), but lower than that caused by cermets. Inhe tests against 100Cr6 pins, electrolytic hard chrome displaysigher friction coefficient than plasma-sprayed chromia andVOF-sprayed cermets. Reversing the results for alumina pins,00Cr6 pins undergo one order of magnitude lower wear rateshen slid against cermets than when slid against Cr2O3. When

liding against alumina, the cermet coatings never undergo sig-ificant material removal, thanks to the formation of a veryompact tribofilm, located on the higher asperities of the sur-ace. EHC is not hard enough to prevent ploughing and cuttingy the alumina pin, though it is tough enough to prevent material
emoval by brittle cracks propagation from pre-existing cracks.s far as thermally treated Ni–P electroless plating are con-
erned, after a short running-in period in which the domes capsre slightly worn (due to very high localized contact pressures),

ssfis

Fig. 11. Three-dimensional axonometric representations of wear scar

10 6.46 × 10 010−5 0 0

film is formed where an oxidized Ni–P compound, possessingow chemical affinity to the counterbody and (probably) highardness, fills the valleys between adjacent domes. Wear scarsf Ni–P electroless platings after tests 4 and 8, respectively, showefinite morphological differences, which will be hereafter dis-ussed and explained. It must be noticed that the sample wearcar after the seemingly anomalous test no. 8 is almost identi-al to that of thermally treated Ni–P, with slightly worn domesaps and Ni–P–O compounds filling roughness valleys. Aluminains undergo both two-body grooving and brittle fracture againstVOF-sprayed WC–Co and thermally treated electroless Nilating respectively, with the latter forming a much smootherear scar on the pin. The effects of polishing on tribologicalerformance of coatings has been the object of a specific inves-igation, comparing the results of tests 7 and 8 performed both ons-deposited and on ground plasma-sprayed ceramics, HVOF-
prayed WC–Co and electrolytic hard chrome. The recordedamples and pins wear rates are listed in Fig. 13A, friction coef-cients are indicated in Fig. 13B. No wear rate is indicated foramples tested against 100Cr6, because only material build-up
s on plasma-sprayed Al2O3 (A) and Cr2O3 (B) after test no. 8.


Ff

igiahtsrfTbaworfwatttlA

Fig. 13. Comparison between pin-on-disk tests no. 7 and no. 8 for as-depositedart

ea1odvam(ctWFig. 14A–D show optical micrographs of some worn alumina

ig. 12. Pin-on-disk test results for all tested coatings. (A) Pin wear rates; (B)riction coefficients.

s found in this case. It must also be noticed that wear rates forround Cr2O3 and WC–Co tested against alumina have beenndicated, while a limited material build-up was recorded ons-sprayed coatings. The wear rates for WC–Co and Cr2O3,owever, are more than two orders of magnitude lower thanhose for the other coatings; thus, their wear performance istill far superior to other tested materials. Such very low wearates could not have been measured on rough samples, but sur-ace grinding has made it possible to determine them accurately.hus, it can be assumed that the overall wear amount undergoney these samples has not actually increased after grinding: likes-sprayed coatings, ground ones do not show a real wear scarith significant material removal from the sample surface, butnly minor surface morphological changes. Thus, samples wearates are not significantly changed in any case, while some dif-erences emerge in pin wear rates and friction coefficients. Pinear rates decrease in all cases (with the exception of 100Cr6 pin

gainst Al2O3–TiO2). Considering alumina pins, Cr2O3 remainshe ceramic coating inflicting less wear to the counterpart, elec-roplated chromium still causes higher wear (comparable to
he other ceramic coatings), but WC–Co now causes muchess wear than hard chrome and plasma-sprayed Al2O3 andl2O3–TiO2, approaching the performance of Cr2O3. Consid-
a(i

nd ground coatings. (A) Sample wear rates in test no. 8 (only positive wearates indicating effective material removal are indicated) and pin wear rates inest no. 7 (steel pin) and test no. 8 (alumina pin); (B) friction coefficients.

ring 100Cr6 steel pins, the wear rates that all ceramic coatingsnd electrolytic hard chrome inflict become very similar (about.3 × 10−5 mm3/Nm), with WC–Co still causing the least wearn the steel pin. The friction coefficients against 100Cr6 pinsecrease for plasma-sprayed ceramics (which display similaralues after polishing) but increase for electrolytic hard chromend HVOF-sprayed WC–Co. Friction coefficients against alu-ina pins decrease for all coatings but Al2O3 and Al2O3–TiO2

which display slightly increased friction) and rankings betweenoatings remain similar, with Cr2O3 causing the lowest fric-ion coefficient against alumina, followed by HVOF-sprayed

C–Co, by hard chrome, and finally by Al2O3 and Al2O3–TiO2.

nd steel pins. The steel pin undergoes two-body abrasive wearcutting and ploughing) when tested against chromia (Fig. 14A);nstead, a film of transferred material is formed on the pin tested


F ings.a lytic

apltisWvasisfmt

4

4

l(eptgfiIr

sshl(uabppcdctbfiftcpmhbidp

ig. 14. Optical micrographs of pin wear scars after testing against ground coatgainst ground WC–17%Co; (C) steel pin after test no. 7 against ground electro

gainst WC–Co (Fig. 14B): it is likely that this film consists inlastically deformed (and probably also partly oxidized) metal-ic debris from the pin itself and from the soft Co matrix ofhe cermet. The pin suffers both abrasive and adhesive wearn tests against electrolytic hard chrome (Fig. 14C). The wearcar on the alumina pin tested against Cr2O3 (Fig. 14D) and

C–Co is very smooth, with some grain boundaries becomingery evident. Profilometry on the wear scar on polished Cr2O3nd WC–Co samples after test no. 8 allowed to notice that theurface roughness of Cr2O3 has decreased because some asper-ties have been worn out and pores (which have been opened beurface polishing) have been filled with wear debris. The sur-ace roughness of WC–Co, instead, has increased, because someetal matrix has been abraded by alumina pin asperities, leaving

he carbide grains protruding out of the surface.

. Discussion

.1. Microstructure and micromechanical properties

Among plasma-sprayed coatings, Cr2O3 seems to possess theowest intersplat cohesion, as emerging from SEM micrographsFig. 1A and B) and from fracture toughness measurement. How-ver, from the same SEM micrographs, it seems that unmoltenarticles are not the reason for the low intersplat cohesion ofhis coating; in fact, all splats appear well spread, indicating
ood in-flight melting. The as-sprayed coating roughness con-rms a very good molten droplets flattening for this coating.t can be argued that the well-spread splats may have too lowoughness to allow proper mechanical adhesion of new-coming
bpl1

(A) Steel pin after test no. 7 against ground Cr2O3; (B) steel pin after test no. 7hard chrome; (D) alumina pin against test no. 8 against ground Cr2O3.

plats. This phenomenon has already been observed with HVOF-prayed alumina splats [42], which spread very well due to theirigh in-flight velocity and thus possess low mechanical inter-ocking. Furthermore, the quite high melting point of Cr2O32330 ◦C [40]) causes molten splats to solidify very quicklypon impact and does not allow underlying splats to reachsufficiently high temperature to activate chemical intersplat

onding. In fact, it is noticeable, from the vast literature onlasma spraying of ceramics (for instance, papers dealing withlasma-sprayed PSZ [43–47] and Al2O3–TiO2 [25,48] can beompared) that the more high-melting-point the ceramic pow-er is, the more difficult it is to achieve a dense, low porosityoating. In the alumina coating, where intersplat cohesion is bet-er, overall fracture toughness is higher (Table 4 and Fig. 3B)ecause crack propagation along splat boundaries is more dif-cult. The low melting point of titania in Al2O3–TiO2 coatingavours better particle melting, causing opposing effects. Onhe one hand, it improves intersplat adhesion, as Fig. 1C indi-ates, confirming many literature work [25,48]. Probably it alsoromotes chemical intersplat bonding as well. Thus, the above-entioned phenomena lessen coating anisotropy. On the other

and, however, the addition of titania allows the formation arittle glassy phase inside the coating. Furthermore, the betterntersplat adhesion also causes more transverse microcrackingue to tensile quenching stresses. Besides, the coarser averageowder particle size causes the appearance of a significant num-
er of unmolten particles. This is why significant cracks botharallel and transverse to the substrate are formed even at aower load than other plasma-sprayed coatings (5 N instead of0 N). Overall, the result is a less tough and less hard coating than

1 ar 261

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4

dci[ciwefcnattigmbh(iuibpusHwpoatttP

amiaVtoctwbwtapCatetAotoTTTt[

w

w(ht

tcecc

w

The coefficient a1 and the a2, a3, a4 exponents have been eval-uated by non-linear least squares fitting of experimental data:a1 = 0.0121, a2 = 0.1290, a3 = −0.7355, a4 = 1.0290.

Table 8KIC-L values and KIC-L/KIC ratios for the plasma-sprayed ceramic coatings

310 G. Bolelli et al. / We

lasma sprayed alumina alone. It is very important to notice thatetter as-sprayed coating surface finishing can be achieved usingner spray powders; in fact, even though chromia is the mostigh melting point among tested ceramics (thus, in theory, theost difficult to melt), it is the one for which the best surface fin-

shing has been obtained in the as-sprayed state. Having a lowers-sprayed surface roughness is very important for technologicalpplications, because it reduces the number of post-depositionechanical treatments necessary. HVOF-sprayed cermets are

ougher than plasma-sprayed ceramics thanks to their compos-te nature with hard particles in a tough metal matrix and thankso a better cohesion, allowed by much higher particles speed inhe supersonic HVOF flame. Nonetheless, the coatings are stillefinitely anisotropic, with cracks mostly propagating parallelo the substrate. These micromechanical features shall obviouslylay an important role in explaining the results of tribologicalests.

.2. Dry particles abrasion resistance

Former research has shown that, in such kinds of tests, theecreasing wear rate of coatings containing metallic phases isaused by abrasive particles which are progressively embeddedn metals or cermets due to the ductility of the metallic phase29,49]. This phenomenon has not limited to this specific testondition, but it has been also documented under different exper-mental configuration (microscale abrasion) [50]. In this way, theear rate is decreased with increasing sliding distance because

mbedded particles alter the characteristics of the sample sur-ace, protecting it. Thus, wear rates for ceramic coatings can bealculated as the average of measurements performed at differentumbers of disk revolutions, but this is not possible for metalsnd cermets. The wear mechanism for hard chrome plating underhe dry sand-steel wheel test is mixed three-body grooving andwo-body rolling (following literature definitions discussed fornstance in [51]), with indenting abrasive particles getting pro-ressively embedded in the coating. The coating thus undergoesicrocutting, microploughing and also microindentation. It can

e seen that the chrome plating is not hard enough to preventard alumina grains to wear the surface by grooving phenomenaploughing, cutting), but withstands indentation without crack-ng thanks to its sufficient toughness. HVOF-sprayed cermetsndergo two and three-body wear as well, as already discussedn [29]. Ceramic coatings do not undergo ploughing and cutting,ecause of the very high intrinsic hardness of the material, com-arable to the one of the abrasive corundum grains; instead, theyndergo wear due to brittle splats removal. Therefore, their abra-ive wear resistance overcomes that of hard chrome plating andVOF-sprayed cermets, for a low number of disk revolutions,hile at high number of disk revolutions the embedded particlesrotect the metallic and cermet coatings from further groovingr indenting phenomena. So, in the latter conditions, cermetsnd hard chrome plating wear rate is better than ceramics. Since
he wear rate for ceramic coatings remain quite constant andhe wear mechanism is not altered increasing the test dura-ion, the steel wheel test can be considered significant for them.lasma-sprayed ceramic coatings, therefore, seem a technically
C

KK

(2006) 1298–1315

ppropriate choice when dry particles abrasion is the main wearechanism involved. The best performance for ceramic coatings

s shown by the toughest one. It would therefore be appropri-te to elucidate the relationship between fracture toughness,ickers microhardness and abrasive wear resistance, to quan-

itatively ascertain the effect of microstructural characteristicsn this property, in view of an optimization for practical appli-ation. An anomaly is immediately apparent when observinghat Al2O3–13%TiO2 is more abrasion resistant than Cr2O3,hich is harder and possesses similar toughness. Two facts muste considered to explain this seeming inconsistency: the mainear mechanism is splats detachment along splats boundaries;

hus, a quantity directly related to intersplat cohesion should bedopted; in fracture toughness tests, long cracks in the directionarallel to the substrate (i.e. along splat boundaries) appeared forr2O3, while similar cracks were developed in al directions forlumina–titania. Thus, fracture toughness by itself might not behe best parameter to explain abrasive wear behaviour; a param-ter taking into account the toughness in the direction parallelo the substrate and the coating anisotropy must be introduced.n empirical but simple parameter can be computed by usingnly the half-diagonal and average crack length in the direc-ion parallel to the substrate in the Evans–Charles formula, thusbtaining an estimation of the crack resistance in this direction.his parameter will be hereafter referred to as KIC-L (MPa m1/2).he KIC-L values and the KIC-L/KIC ratio are listed in Table 8.he KIC-L/KIC ratio will be employed in a modified version of

he Evans–Marshall formula for abrasive wear rate prediction52]:

= a(E/H)P1/8

K1/2IC H5/8

(4.1)

here a is the material independent constant, w = wear ratemm3/Nm), E the elastic modulus (GPa), H the Vickers micro-ardness (GPa), P the normal load (N), and KIC is the fractureoughness (MPa m1/2).

The formula will be modified by multiplying the denomina-or by the KIC-L/KIC ratio. Since the normal load has not beenhanged, its effect cannot be accounted for and will be consid-red as a part of the coefficient a. Since the formula has beenhanged, it is likely that the exponents value shall have to behanged as well. The resulting expression is as follows:

= a1(E/H)

Ka2IC(KIC-L/KIC)a3Ha4

(4.2)

oating Al2O3 Al2O3–TiO2 Cr2O3

IC-L (MPa m1/2) 1.76 ± 0.88 3.30 ± 1.23 0.85 ± 0.40

IC-L/KIC 0.68 ± 0.34 1.28 ± 0.48 0.51 ± 0.33

G. Bolelli et al. / Wear 261

Fig. 15. Dry sand-steel wheel test results for plasma-sprayed ceramicca(

ptbesfigwrci[opeb

4

tat

wcficbbfiiTsfittogaacfcWpcpbtpocapfcistoltpoiNlpcafmhma

oatings fitted with Eqs. (4.1) and (4.2). (A) Wear rates plottedgainst (E/H)P1/8/(K1/2

IC H5/8); (B) wear rates plotted against (E/H)Ka2

IC(KIC-L/KIC)a3 Ha4 ).

Fig. 15A shows that the Evans–Marshall equation fails inredicting the abrasive wear rates of the various coatings, whilehe new equation (Fig. 15B) fits the experimental data. It shoulde noticed that, in the current work, only three points have beenmployed to fit Eq. (4.2), while, for a proper statistical analy-is, a greater number of points should have been employed. Thet results, therefore, should not be considered as being valid ineneral. The above considerations only seek to indicate a neway for the microstructural interpretation of abrasive wear test

esults of plasma-sprayed coatings. Although there exist doubtsoncerning the effective usefulness of such predictive theoret-cal or semiempirical formulae for thermally sprayed coatings53,54], they can be a useful reference for a simple estimationf the performance of a coating, and a basis for a future a-priorirocess design, once the correlations between operating param-ters and micromechanical properties shall be fully elucidatedy further fundamental research.

.3. Pin on disk wear test

It is a basic principle that, in a tribological coupling betweenwo different surfaces, the harder one wears the softer one: sincell tested coatings but Ni–P plating without treatment are harderhan 100Cr6, they never undergo a wear loss except for Ni–P

lscF

(2006) 1298–1315 1311

ithout thermal treatment. The only way by which materialould be removed from samples surface would be by crackormation and propagation by cyclic contact fatigue; however,t appears that contact stresses have not been high enough toause such phenomenon. Tribological phenomena in couplingsetween the steel pin and tested coatings seem to be influencedy the coating roughness. In many cases, the steel pin forms alm of transferred material on the coatings, thus, after a running-

n period, a steel–steel contact governs tribological phenomena.he morphology of the transferred film is influenced by coatingurface finishing: considering as-sprayed ceramic coatings, thelm is thinner and smoother on the Cr2O3 one, which possess

he lowest surface roughness, while it is coarser and thicker onhe others. This explains the lower friction coefficient recordedn the as-sprayed chromia coating. When ceramic coatings areround, they all cause a similar friction coefficient and pin weargainst the 100Cr6 pin, because they all form a quite smoothnd thin film of transferred material. On the electrolytic hardhrome plating, no significant transferred film is formed, but theriction coefficient is the highest among tested coatings: in thisase, chemical affinity between the surfaces must be considered.

hile ceramic coatings have no chemical affinity for the 100Cr6in, thus adding no further source of friction and wear, a definitehemical affinity between the two metallic surfaces of electro-lated chromium and 100Cr6 certainly exists. This is confirmedy optical micrographs indicating adhesive wear taking place onhe steel pin in this case. Therefore, when the chrome plating isolished, the friction coefficient slightly increases, because theverall extent of the actual contact area is increased. The sameonsiderations are valid for as-deposited Ni–P platings as well;fter the thermal treatment, instead, the appearance of the Ni3Phase and the formation of oxidized phases on the coating sur-ace definitely lowers the chemical affinity, reducing the frictionoefficient. A very peculiar situation occurs for cermet coatings:n the as-sprayed condition, the contact is localized on coatingurface asperities, so that the pin is initially worn by few pro-ruding carbide particles and the steel debris form a transfer filmn the coating. After polishing, the contact area becomes mucharger and also involves the soft Co matrix. In this situation,he chemical affinity between the soft Co matrix and the steelin becomes relevant, a film of transferred material is formedn the pin itself and the friction coefficient increases, althought remains lower than that for hard chrome and as-depositedi–P, both because of transfer material oxidation and because a

arge part of the cermet surface consists of ceramic WC particlesossessing low chemical affinity for steel. The friction coeffi-ient and pin wear rate decrease with increasing contact load forlmost all coatings and with increasing sliding speed as well,or ceramic coatings. Considering as-sprayed ceramic and cer-et coatings, where a transferred steel film is formed, under a

igher contact load the transferred film on the sample surface isore rapidly built up, reducing wear caused by the hard coating

sperities. Furthermore, under a higher normal load, a higher
ocal temperature is reached in the contact point; thus, the steelurface is more rapidly oxidized, reducing the friction coeffi-ient. The latter event also occurs for increased sliding speed.or electrolytic hard chrome, where no significant transferred

1 ar 261

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teel film appears, it is likely that the increased contact pointemperature favours metallic surfaces oxidation, slightly reduc-ng chemical affinity troubles. Sintered alumina is harder thanll tested ceramic coatings; therefore, it is able to inflict wearn them and, at the same time, it undergoes much less wearhan 100Cr6 (wear rates ≤10−5 mm3/Nm), Fig. 10B. The pinndergoes grooving phenomena (ploughing and cutting) whichre probably due to the formation of a thin, softer surface filmonsisting in wear debris, and also undergoes a limited extentf brittle cracking and grain pull-out, with rougher wear scarsor higher recorded wear rates. Comparing the present wearates and wear mechanisms with literature work on wear of sin-ered ceramics [30–32], the present alumina pin wear seemsn a border-line situation between mild and severe wear. Thein wear rate seems to be increasing both with normal load,hich favours brittle cracking, and with sliding speed, which

ncreases flash temperature in the contact point favouring ther-al shock cracking [30], although a few exceptions exist in the

resent experimental data set. Obviously, since brittle fractures involved in alumina pin wear, reducing the contact pressuren surface asperities by sample grinding lowers wear aluminain wear rates and also lowers friction coefficients. This is par-icularly obvious for the WC–Co cermet: the wear rate it inflictsn the alumina pin decreases by more than two orders of magni-ude after grinding. The measured wear rate (<10−7 mm3/Nm)nd the smooth pin wear scar, with no brittle cracking, clearlyndicates the occurrence of a mild pin wear regime (accordingo [30–32]) in tests against ground WC–Co and Cr2O3. Fric-ion coefficient and pin wear are quite high for electrolytic hardhrome and for Al2O3 and Al2O3–TiO2 coatings: for these lat-er, chemical affinity is likely to be the dominating effect. Forard chrome, some wear debris from coating material is trans-erred to the alumina pin, so that a metal–metal contact arises,ncreasing the friction coefficient.

Tests against alumina pins also allow to evaluate the wearesistance of the various coatings, since alumina is hard enougho inflict wear on most of them. It appears that surface tri-ofilms formation is an important tribological phenomenon inost cases: these tribofilms consist in plastically deformed wear

ebris, plastically deformed coating material or in a surfacehemical alteration of the coating. If a tribofilm is not formed,he coating cannot oppose to continuous material removal byhe harder counterpart asperities, as for electrolytic hard chromend as-deposited Ni–P. The tribofilm formation, however, isnly beneficial if the tribofilm itself possesses adequate cohe-ion. Let us examine the case of as-sprayed ceramic coatings.heir tribofilm consists in plastically deformed splats and wearebris: plastic deformation is caused by high local contact pres-ures and by high local contact temperatures, which are due toeat generation by friction and to poor thermal conductivity oferamic materials. A significant compressive stresses distribu-ion appears just below the surface at contact points becausef high contact pressures: it is known that, under high idro-
tatic stress, ceramics can plastically deform at the microscale.esides, high local temperatures lessen the asperities hard-ess, favouring plastic deformation. Tribological studies onoth plasma-sprayed [23,24,55] and massive [30–34] ceram-
mfita

(2006) 1298–1315

cs confirm the occurrence of microscale plastic deformation ineramics: material sliding at discrete “shear faults” (for instance,rain boundary sliding) is thought to be the cause of suchhenomenon. However, the wear scar appearance in massiveeramics, as the above mentioned literature works report andhe present observations of alumina pin wear scars confirm, is

uch different from the one in plasma-sprayed coatings: in mas-ive ceramics, the plastically deformed film is very thin andompact, with few debris embedded in grain boundaries. Inlasma-sprayed coatings, instead, a thicker film appears, withhigher amount of debris on all the contact surface: this is

robably due to material detachment from surface asperities,ecause of the quite high initial surface roughness. Besides, theplats themselves seem to undergo extensive plastic deforma-ion: it is known that, due to rapid cooling, splats consist in verymall crystals (≤1 �m) [16]; thus, a high overall grain bound-ry surface provides a great amount of shear faults for plasticeformation; moreover, splat boundaries themselves may con-titute favourable areas for plastic slipping [23,24]. With thisssumption, plastic deformation without sub-surface crackings more favoured in the Cr2O3 coating, where splat cohesion isower. Thus, surface splats on the chromia coating are quite freeo deform plastically without sub-surface cracking, while, onther coatings, splats are less free to deform plastically and areore easily cracked due to elastic stresses accumulation. Thus,

n Al2O3 and Al2O3–TiO2, the film has a quite loose structure.uring the test, it is continuously removed and reformed andoes not protect the coating, accounting for the higher wear rate.n plasma-sprayed Cr2O3, the compact film effectively protects

he surface from damage; in fact, SEM micrographs (Fig. 9Cnd D) indicate that, below this film, there are unaltered splats.urthermore, the film smoothness may partly account for the

ower friction coefficient recorded in this case. Another reasonor the lower friction coefficient might be the appearance of CrO2nd CrO3 on the very surface of the film, reducing the frictionoefficient [55]. Furthermore, the higher hardness of the Cr2O3oating implies the compact tribofilm, once formed, cannot beasily abraded by pin surface asperities or loose debris. On cer-ets, the occurrence of a tribofilm has already been documented

n literature [56,29], and, as explained in [29], its formation andts compactness are the effect of the exceptional ductility ofhe Co-based metal matrix. The metal matrix, in fact, is plasti-ally deformed around contact points because of high contacttresses, and part of the soft metal is abraded by asperities onhe counterbody. The plastically deformed metal strongly holdshe small carbide particles, some of which have been crackedut, nonetheless, are still embedded. Such particles slightly pro-rude out of the metal matrix and, being very small and closeo each other, prevent further abrasion of the metal matrix. Theetached metal debris (together with some very small WC frag-ent) mix with the alumina debris and acts as bonding phase,

llowing debris sticking onto the cermet surface. Therefore, filmreas consisting in alumina particles and WC fragments held by
etal matrix are found next to the protruding particles. Thislm prevents any further wear on the coating, in fact, the con-
act is localized on WC particles, which are harder than thelumina counterbody. The hardness of these carbide particles,

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n the other hand, may cause some wear on the pin (Fig. 12A).s far as thermally treated Ni–P is concerned, the tribofilm is

ormed by chemical alteration of the coating surface: once theomes caps are smoothed by high contact pressures, the con-act point temperature is high enough to form oxidized Ni–P–Oompounds which definitely lessen the friction coefficient andontribute (together with the high hardness and toughness ofhe nanostructured coating) to arrest any further coating wear.he response from the Ni–P electroless plating in test 8, where

he pin and sample wear rate and friction coefficient resemblehermally treated Ni–P, deserve a special comment. The contactemperature must have been high enough to cause nanocrystal-ization along the contact area, so that in the contact area, theoating behaves almost exactly like the thermally treated one.t is interesting to observe how an improves surface finishinglters the wear behaviour. On Cr2O3, the higher actual contactrea reduces initial contact stresses around surface asperities, sohat less debris appears. Thus, the film only consists in plasticallyeformed splats, with few debris filling open pores. As a result,very smooth wear scar appears, and no material build-up is

ecorded, but only a very low wear rate, as shown in Section 3.n WC–Co, the higher actual contact area prevents local plasticeformation of the metal matrix around surface asperities, as itappened in as-sprayed coating. Besides, on the ground surface,arbide particles are already slightly protruding, because theoft Co matrix is much more easily ground and removed. Thus,he contact is already located on the very hard carbide particles.hus, no tribofilm formation occurs because the favourable con-ition of carbides slightly protruding out of the metal matrixas already been accomplished by the grinding operation: theround surface itself acts as the tribofilm. The only noticeableear phenomenon is the removal of a small amount of Co matrixy the highest counterpart asperities: once the carbide particlesre protruding out enough and the counterpart surface asperi-ies are worn away, no further wear occurs on the coating. Themoothness of the wear scars, little amount of wear debris ando material build-up for ground Cr2O3 and WC–Co coatings isonsistent with the onset of a mild wear regime for the aluminain. This accounts for the absence of material build-up on theurface and for the very low measured wear rate. Some literatureork [57,58] indicates that wear due to brittle crack propaga-

ion could occur: in this case, the contact pressure was not highnough. Besides, crack propagation was reported to be causedy metal matrix embrittlement due to carbides dissolution: inhis case, the optimal choice of spraying parameters preventedn excessive extent of carbides dissolution, so that the coatings tough enough to prevent crack propagation under reasonableontact loads. It is practically unusual to find application whereertzian contact pressures are allowed to exceed 1 GPa [59].

. Conclusions

Three plasma-sprayed oxide ceramic coatings, namely
r2O3, Al2O3, Al2O3–13%TiO2, have been characterized in
erms of microstructure and micromechanical properties. Theirry particles abrasion resistance (dry sand-steel wheel test)nd pin-on-disk wear test response with 100Cr6 steel and

(2006) 1298–1315 1313

intered alumina counterbodies (sample wear rate, pin wearate, friction coefficient) have been compared to the behaviourf two HVOF-sprayed cermet coatings (WC–17%Co andC–10%Co–4%Cr), electrolytic hard chrome, untreated and

hermally treated electroless Ni–P coatings. The experimentalesults lead to the following conclusions:

Cr2O3 coating is the hardest and most anisotropic amongplasma-sprayed ceramics, due to low interlamellar cohesion;Al2O3–13%TiO2 is the most isotropic but also the less hardand less tough one, due to the formation of an alumina–titaniaglassy phase which favours intersplat adhesion but turns outto be quite brittle. WC–17%Co and WC–10%Co–4%Cr arealso anisotropic, but, thanks to a ductile metal matrix and toa composite microstructure, are tougher than plasma-sprayedceramics.Plasma-sprayed ceramics are the better-performing coatingsin dry particles abrasion conditions, because they neverundergo abrasive grooving but only splats detachment. Inparticular, plasma-sprayed Al2O3 shows the best dry parti-cles abrasion resistance, thus, its employments are suggestedunder such conditions. Fracture toughness, by itself, is nota good parameter to predict plasma-sprayed ceramics dryparticles abrasion resistance: the parameter KIC-L, related tothe crack propagation resistance in the direction parallel tothe substrate (i.e. intersplat cracking), must be introduced.A combination between KIC-L and Vickers microhardness issuggested to predict the dry particles abrasion resistance ofplasma-sprayed ceramics. Cermets have similar performanceto electrolytic hard chrome.In the pin-on-disk test, the ability to form a smooth andcompact surface film by local plastic deformation is the keyproperty determining coatings performance. Plasma-sprayedAl2O3 and Al2O3–13%TiO2 do not form an adequately com-pact tribofilm and therefore show unfavourable properties interms of sample wear rate, pin wear rate and friction coef-ficient. Cr2O3, instead, has the ability to form a compacttribofilm by splats plastic deformation; thus, its performancesare comparable to HVOF-sprayed ceramics. Among the thick(≥100 �m) coatings tested in this study, plasma-sprayedCr2O3 and HVOF-sprayed cermets are the best performingcoatings under all aspects and are all candidate coatings forelectrolytic hard chrome replacement. Plasma-sprayed chro-mia, performing similarly to cermet coatings though havingpotentially lower production costs, seems an interesting solu-tion. In particular, Cr2O3 powders, although more expensivethan most ceramic ones (at least thrice more expensive thanalumina), are cheaper than cermets. However, potential dis-advantages, which must be considered in a proper coatingchoice, include the rather low deposition efficiency of Cr2O3(≤45% in conventional APS [60]) and environmental prob-lems due to possible formation of Cr6+ by in-flight thermalalteration of the coating material during spraying. Some strict
rules have been recently issued in USA due to this problem[61]. Further development of this research should also involvemore environmental friendly alternatives to thermally sprayedCr2O3 able to reproduce its excellent wear behaviour.

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Polishing generally improves the tribological performance oftested coatings; however, the friction coefficient of WC–Cocoating against 100Cr6 has been increased because chemicalaffinity troubles between steel and metal matrix have beenenhanced. As-sprayed Cr2O3 is smooth enough to cause lim-ited wear on the counterpart and limited friction coefficients,so that, in certain instances, it could also be employed in theas-sprayed condition. This points out to the need for properdeposition parameters and the advantage in using spray pow-ders with a quite small average particle diameter (it must notbe too small, however, to prevent droplets vaporization in theplasma jet and to confer molten droplets an adequate kineticenergy to flatten upon impact).The present data are also a good basis for constructing wearmaps of thermally sprayed coatings, in particular for plasma-sprayed Cr2O3, which is of high practical interest due toits favourable wear resistance. Such wear maps would bevery useful in view of an increasing industrial applicationsin many fields where traditional plating techniques are stillbeing employed. In particular, it is to be noticed that, in allcases considered here, the chromia coating is in the mild orultra-mild wear regime; however, this depends on the factthat this coating has, in all cases, been able to form a com-pact tribofilm by plastic deformation: since the tribofilm isentirely ceramic (thus, intrinsically brittle), once a criticalcontact pressure has been exceeded, a tribofilm detachmentand, consequently, transition from mild to severe wear arelikely to occur. The identification of this transition is of keypractical importance and shall be a future development of thepresent work.

cknowledgements

The authors thanks Centro Sviluppo Materiali S.p.A. (Roma,taly), in particular Mr. Edoardo Severini, Mr. Francesco Barulli,

r. Valerio Ferretti and the head of the surface engineering unit,ng. Fabrizio Casadei, for the thermally sprayed coatings manu-acturing. We are also grateful to Galvanica Nobili S.r.l. (Maranoul Panaro, MO, Italy) for hard chrome platings and to Argos.p.A. (Monteveglio, BO, Italy) for Ni–P electroless platingsanufacturing. A special thank to Ms. Lisa Calamai for her

aluable contribution to the experimental activity. Partially sup-orted by PRRIITT (Regione Emilia Romagna, Italy), Net-LabSurface & Coatings for Advanced Mechanics and Nanome-hanics” (SUP&RMAN).

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wear behavior of thermally sprayed ceramic oxide coatings

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