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9 Ž. Surface and Coatings Technology Effects of tungsten carbide thermal spray coating by HP HVOF and hard chromium electroplating on AISI 4340 high strength steel Marcelino P. Nascimento a,, Renato C. Souza b, Ivancy M. Miguel a, Walter L. Pigatin c, Herman J.C. Voorwald a a State Uni ersity of Sao Paulo- DMT-UNESP FEG, A. Ariberto Pereira da Cunha, 333-Guaratingueta SP BR-CEP: , Brazil b FAENQUIL DEMAR, Lorena SP BR-CEP: , Brazil c EMBRAER-LIEBHERR EDE, Sao Jose dos Campos SP BR-CEP: , Brazil Received 26 August 1999; received in revised form 5 October 2000; accepted 7 November 2000 Abstract In cases of decorative and functional applications, chromium results in protection against wear and corrosion combined with chemical resistance and good lubricity. However, pressure to identify alternatives or to improve conventional chromium electroplating mechanical characteristics has increased in recent years, related to the reduction in the fatigue strength of the base material and to environmental requirements. The high efficiency and fluoride-free hard chromium electroplating is an improvement to the conventional process, considering chemical and physical final properties. One of the most interesting, environmentally safer and cleaner alternatives for the replacement of hard chrome plating is tungsten carbide thermal spray coating, applied by the high velocity oxy-fuel Ž HVOF. process. The aim of this study was to analyse the effects of the tungsten carbide thermal spray coating applied by the HP HVOF process and of the high efficiency and fluoride-free hard chromium electroplating Ž in the present paper called accelerated., in comparison to the conventional hard chromium electroplating on the AISI 4340 high strength steel behaviour in fatigue, corrosion, and abrasive wear tests. The results showed that the coatings were damaging to the AISI 4340 steel behaviour when submitted to fatigue testing, with the tungsten carbide thermal spray coatings showing the better performance. Experimental data from abrasive wear tests were conclusive, indicating better results from the WC coating. Regarding corrosion by salt spray test, both coatings were completely corroded after 72 h exposure. Scanning electron microscopy technique Ž SEM. and optical microscopy were used to observe crack origin sites, thickness and adhesion in all the coatings and microcrack density in hard chromium electroplatings, to aid in the results analysis Elsevier Science B.V. All rights reserved. Keywords: Tungsten carbide thermal spray coating; Hard chromium electroplating; Abrasive wear; Corrosion; Fatigue; HP HVOF Corresponding author. Tel.: ; fax: Ž. address: pereira@feg.unesp.br M.P. Nascimento $ - see front matter 2001 Elsevier Science B.V. All rights reserved. Ž. PII: S

10 114 ( ) M.P. Nascimento et al. Surface and Coatings Technology Introduction Chromium plating is the most used electrodeposited coating to obtain high levels of hardness, resistance to wear and corrosion and a low coefficient of friction for applications in the aerospace, automotive and petrochemical fields 1,2. Chromium plating properties, such as hardness and microcrack density, change with the bath composition, current density, bath agitation, temperature, etc. 3,4. Among other things, a significant characteristic of chromium electroplating is the high tensile residual internal stresses originating from the decomposition of chromium hydrides during the electrodeposition process 3 5. These high tensile stresses in electroplated chromium coatings increase as thickness increases and are relieved by local microcracking during electroplating. Therefore, basically, microcrack density is related to the high tensile residual internal stresses, hardness, and corrosion resistance 2,3,6. It was observed that the residual stresses through-thickness decrease with the depth of the coating and increase again at the coating substrate interface 7. Bending fatigue tests on samples with different coatings and coating conditions indicate that the fatigue strength is dependent on the fracture behaviour of the substrates and on the hardness and residual stresses at the substrate surface. It was also observed that the hard chromium electroplating reduces the fatigue strength of a component 8. Due to this fact, the design of hard chromium plated components, which are subjected to dynamic loads, may consider this negative influence to guarantee safety during operation. Therefore, the use of effective methods to improve the fatigue strength shall be considered. Shot peening is a well-known process to increase fatigue life of structures subjected to constant and variable amplitude loading. The compressive residual stress obtained by surface plastic deformation is responsible for the increase in fatigue strength in shot peened mechanical components 9. Compressive residual stresses induced by machining processes are also responsible for the improvement in fatigue resistance of AISI 4340 steel 10. Increase in the fatigue crack propagation resistance in AISI 4340 steel with electron beam surface hardening was associated to residual stress distribution and microstructural characteristics 11. However, problems concerning chrome plating, such as health and environmental hazards, increasing costs and a performance not in accordance with the specifications, have resulted in a search to identify possible alternatives 12. Aircraft landing gear manufacturers are considering tungsten carbide Ž WC. thermal spray coatings applied by the high velocity oxy-fuel Ž HVOF. process as an alternative to hard chrome plating. The question to be answered is if the performance of the alternative candidate is at least comparable to results obtained for hard chrome plating. Comparisons of experimental data showed better corrosion resistance for several HVOF coatings with respect to chrome plating. In the case of fatigue and friction tests, the results were acceptable, indicating interesting perspectives on the use of tungsten carbide coating to replace chrome-plating 1. Analysis of the wear performance of tungsten carbide coated samples in the presence of air, aqueous and aqueous abrasive media indicated better results in terms of volume loss and change in surface roughness than for the mild steel substrate 13. The objective of this research is to compare the influence of the tungsten carbide thermal spray coating applied by HP HVOF and hard-chromium plating on the fatigue strength, abrasive wear and corrosion resistance of AISI 4340 steel. S N curves were obtained in rotating bending and axial fatigue tests for the base material, chromium plated, and tungsten carbide coated specimens. 2. Experimental procedures AISI 4340 steel is widely used in aircraft components where strength and toughness are fundamental design requirements. The chemical analysis of the material used in this research indicates accordance with specifications. The fatigue experimental program was performed on rotating bending and axial fatigue test specimens machined from hot rolled, quenched and tempered bars according to Figs. 1 and 2, respectively. The specimens were polished in the reduced section with 600 grit papers, inspected dimensionally and by magnetic particle inspection. Fatigue tests specimens were quenched from C in oil Ž 20 C. and tempered in the range of C for 2 h. The mechanical properties of the material after the heat treatment are: hardness of 39 HRC; yield tensile strength of 1118 MPa, and ultimate tensile strength of 1210 MPa. After final preparation, samples were subjected to a stress relieve heat treatment at 190 C for 4 h to reduce residual stresses induced by machining. Average superficial roughness in the reduced section of the samples was R 2.75 m and a standard deviation Ž S.D.. a of 0.89 m. Rotating bending fatigue tests were conducted using a sinusoidal load of frequency 50 Hz, and load ratio R 1, at room temperature. For axial fatigue tests, a sinusoidal load of frequency 50 Hz and load ratio R 0.1 was applied throughout this study. Both tests consider as fatigue strength the complete fracture of the specimens or 10 7 load cycles. Four groups of fatigue specimens were prepared to obtain S N curves for rotating bending fatigue and axial fatigue tests.

11 ( ) M.P. Nascimento et al. Surface and Coatings Technology Fig. 1. Rotating bending fatigue testing specimen For rotating bending fatigue tests rotating bending fatigue specimens were tested without blasting. 12 smooth samples of base material; 13 samples with 160 m thickness of conventional hard chromium electroplated; 13 samples with 100 m thickness of accelerated hard chromium electroplated; and 13 samples with 100 m thickness of tungsten carbide thermal spray coated by HP HVOF process Salt spray test The performance of the coatings was evaluated with respect to chemical corrosion in specific environment. The samples were prepared from normalised AISI 4340 steel with 1 mm thickness, and 76 mm width and 254 mm length, surface roughness Ra 0.2 m, and in the following conditions: 2.2. For axial fatigue tests 10 smooth samples of base material; 15 samples with 160 m thickness of conventional hard chromium electroplated; 7 samples with 100 m thickness of accelerated hard chromium electroplated; and 13 samples with 100 m thickness of tungsten carbide thermal spray coated by HP HVOF process. The tungsten carbide thermal spray coated specimens were blasted with aluminium oxide mesh 90 to enhance adhesion. To compare experimental data, six Fig. 2. Axial fatigue testing specimen. accelerated hard chromium electroplated 16, 36 and 49 m thickness; conventional hard chromium electroplated 16, 36 and 49 m thickness; and tungsten carbide thermal spray coated 100 m thickness. Experimental tests were conducted in accordance with ASTM B 117, in 5 wt.% NaCl, ph of , at 35 C. The samples were supported at 20 from the vertical. The results were analysed by Image Pro Plus software Abrasi e wear test The performance of the coating was also evaluated with respect to abrasive wear. For abrasive wear tests, samples were prepared from annealed AISI 4340 steel with 4 mm thickness and 100 mm square, according to FED-STD-141C. The samples were divided in three groups; two coated with 100- m thickness of accelerated and conventional hard chromium electroplating, respectively, and one group coated with 100- m thickness of tungsten carbide coating. The wear tests were conducted with a Taber abraser, at room temperature, using a 10-N load and CS-17 abrading wheel for hard chromium electroplating and diamond wheel for tungsten carbide coating. The results were analysed by wear index Ž mg 1000 cycles. and total wear Žmg cycles. data.

12 116 ( ) M.P. Nascimento et al. Surface and Coatings Technology Tungsten carbide coating The tungsten carbide thermal spray coating applied by HP HVOF system, used WC powder with 12% Co, resulting in thickness equal to 100 m. The average superficial roughness in the reduced section of the samples was Ra 4 m and a S.D. of 0.39 m, in the as-deposited condition Hard chromium electroplating The conventional hard chromium electroplating was carried out from a chromic acid solution with 250 g l of CrO3 and 2.5 g l of H 2SO 4, at C, with a current density from 31 to 46 A dm 2, and a speed of deposition equal to 25 m h. A bath with a single catalyst based on sulfate was used. The accelerated hard chromium electroplating was carried out from a chromic acid solution with 250 g l of CrO3 and 2.7 g l of H 2SO 4, at C, with a current density from 55 to 65 A dm 2, and a speed of deposition equal to 80 m h. A bath with a double catalyst, one based on sulfate and the other without fluoride, was used. After the coating deposition, the samples were subjected to a hydrogen embrittlement relief treatment at 190 C for 8 h. The average surface roughness of the hard chromium electroplating was Ra 3.13 m in the reduced section and a S.D. of 0.79 m in the as-electroplated condition. For the microcrack determination in both hard chromium electroplating, samples were prepared from normalised AISI 4340 steel Ž R 0.2 m. a, 1 mm thick- ness, 25 mm width and length, and with accelerated and conventional hard chromium electroplating both with 100 m thickness, which resulted in a surface roughness of Ra 0.74 m for the former and Ra 1.6 m for the later, in the as-electroplated condition. The surface microcracks were enhanced through anodic etching for 30 s with a current density equal to 25 A dm 2 in the same chromium bath and later analysed using an optical microscope model Nikon Apophot. All surface roughness data measured in this research was obtained by Mitutoyo 301 equipment using a cut-off of 0.8 mm. The analysis of fracture surface was carried out on rotating bending fatigue specimens by scanning electron microscope, model LEO 435 vpi and Zeiss DSM 950. The metallographic analysis was carried out on optical microscope model Neophot Results and discussion 3.1. Fatigue test The S N curves for the rotating bending and axial fatigue tests for the base metal and coated specimens are presented in Figs. 3 and 4, respectively. Fig. 3 shows that the effect of coating in the rotating bending fatigue test is to decrease the fatigue strength of AISI 4340 steel. The tendency is observed for low number of cycles Ž10 4., high number of cycles Ž10 5. and for the fatigue limit, 10 7 cycles, and is represented in Table 1. One sees that the specimens coated with tungsten carbide applied by the HVOF process show a lower decrease in fatigue strength. This may be attributed to the process itself. It is well known that HVOF thermal spray process produces compressive Fig. 3. S N curves for rotating bending fatigue tests.

13 ( ) M.P. Nascimento et al. Surface and Coatings Technology Fig. 4. S N curves for axial fatigue tests. residual internal stresses within the substrate, which are formed from mechanical deformation on the surface during particle impact. This is confirmed by the through-thickness residual stress behaviour shown in Fig. 5. These surface deformations counteract the tensile shrinkage stresses of the coating caused by fast cooling and solidification as particles strike the surface. These tensile stresses in the coating also generate compressive stresses within the surface of the substrate. However, there was a reduction in the fatigue Table 1 Rotating bending fatigue strength. Rotating bending fatigue strength Group 4 Low cycles Ž High cycles Ž limit Ž 10. Base material 950 MPa Ž 85%. 700 MPa Ž 63%. 615 MPa Ž 59%. ys ys ys Treat. T. carb Ž 100 m. 900 MPa Ž 80%. 610 MPa Ž 54%. 531 MPa Ž 47.5%. ys ys ys Tungsten carb. Ž 100 m. 900 MPa Ž 80%. 570 MPa Ž 51%. 531 MPa Ž 47.5%. ys ys ys Conv. chrome Ž 160 m. 840 MPa Ž 75%. 500 MPa Ž 45%. 321 MPa Ž 29%. ys ys ys Accel. chrome Ž 100 m. 730 MPa Ž 65%. 340 MPa Ž 30%. 280 MPa Ž 25%. ys ys ys Fig. 5. Through-thickness residual stress distribution for WC HP HVOF thermal spray coating.

14 118 ( ) M.P. Nascimento et al. Surface and Coatings Technology Fig. 6. Microcracks surface network in conventional Ž. a and accelerated Ž b. hard chromium electroplating. Anodic etching from 25 2 A dm for 30 s Ž strength of AISI 4340, despite the compressive residual stresses induced by the process. This can be due to the high density of pores and oxide inclusions in the coating that commonly form during the process. Thermal spray is generally conducted in air so chemical interactions occur, notably oxidation, which can be evident in the coating microstructure as oxide inclusions, mainly in grain boundaries 14. These inclusions in coating subsurfaces are possible cracks or nucleation initiation sites. From Fig. 3, it is possible to observe that the aluminium oxide blasting is responsible for a small improvement in the fatigue strength. Considering both hard chromium electroplated rotating bending fatigue results, one sees the negative influence of coating on the fatigue strength of the steel. From the analysis of these two coatings, it is possible to observe the better performance of the conventional hard chromium in relation to the accelerated hard chromium electroplating, despite the higher thickness of the former. This may be attributed to the lower microcrack density of the conventional hard chromium electroplating in comparison with the accelerated hard chromium electroplating as showed in Fig. 6. The microcracks density quantitative analysis indicated median values of 1512 microcracks cm and a S.D. of microcracks cm for the accelerated hard chromium electroplating, and 223 microcracks cm with standard deviation of 57.5 microcracks cm for the conventional hard chromium electroplating. Microcracks form when the high tensile residual internal stresses exceed the cohesive strength of the chromium deposits and affect the fatigue behaviour of a plated part. Therefore, microcrack density arises as a relief of the tensile residual internal stresses, which increase when the chromium thickness increases. Pina et al. 7 showed that the microcrack density changes along the thickness, being higher at the core and lower at the surface of the coating and in the substrate coating interface due to the balance between the residual stresses. On the surface of the coating, the microcracks arise in a network shape, without preferential direction and characterising an equi-biaxial residual stress state. With respect to residual stresses, an inverse behaviour from that observed for the microcrack occurred. Therefore, in general, the higher the microcrack density, the higher the tensile residual internal stresses and or their relief. This means that the accelerated hard chromium electroplating is responsible for higher tensile residual internal stresses and or present the highest crack initiation propagation front amount. However, the different microcrack densities between both hard chromium electroplatings practically produced the same effect in low cycle fatigue, since crack growth occurs after few cycles of fatigue testing. In general, the fatigue strength for all conditions studied was reduced due to the high tensile residual internal stresses, microcrack density, and high adhesion at the coating substrate interface, which allows the crack to grow from the coating through the interface into the base metal. Fig. 5 shows the residual internal stress profile for tungsten carbide thermal spray coating. From experimental points, one sees that the residual internal stresses change throughout coating thickness, from 100 MPa tensile stress near to the surface, reaching 350 MPa maximum tensile stress at a mm depth and decreasing to the maximum compressive stress of 680 MPa at a 0.07-mm depth, increasing again into the base metal, becoming tensile stress at a 0.20-mm depth. The curve of through-thickness residual stress was plotted based on three specimens and obtained by the modified layer-removal method for thermal spray coating and substrates 18,19. The through-thickness residual stresses change from approximately 300 MPa tensile at mm depth to approximately 680 MPa, compressive at 0.06 mm from the surface. This means that the crack initiation may occur easily on the coating surface, but its propagation throughout the thickness may be delayed when the compressive residual stress site is reached.

15 ( ) M.P. Nascimento et al. Surface and Coatings Technology Table 2 Axial fatigue strength Axial fatigue strength Group 4 a Low cycles Ž High cycles Ž Limit Ž 10. Base material 1330 MPa Ž 119% MPa Ž 101%. 850 MPa Ž 76%. ys ys ys Tungsten carb. Ž 100 m MPa Ž 121%. 850 MPa Ž 76%. 750 MPa Ž 67%. ys ys ys Conv. chrome Ž 160 m MPa Ž 107%. 650 MPa Ž 58%. 400 MPa Ž 36%. ys ys ys Accel. chrome Ž 100 m MPa Ž 103%. 650 MPa Ž 58%. 400 MPa Ž 36%. a Ž 4. Projection of the curves until the respective number of cycles 10. ys ys ys For hard chromium electroplating residual stresses, Pina et al. 7 showed that, despite the fact that microcracks result in residual stress relief in the coating, the stresses still remain high at the surface Žapprox. 800 MPa., decreasing in direction to the core Žapprox. 200 to 300 MPa., and increasing again at the interface to values which depend on the substrate material. Fig. 4 shows the axial fatigue testing results, indicating the decrease in fatigue strength for all specimens coated with tungsten carbide thermal spray and hard chromium electroplating, in comparison to the base material. Comparing the curves, one sees the negative influence of coatings on the fatigue strength of the steel, with the same tendency observed previously in the rotating bending fatigue tests. This behaviour can also be explained by high tensile residual internal stresses, oxide inclusions, pores and microcracks inherent from each process. Table 2 indicates the axial fatigue strength tendency of all specimens groups, based on Fig. 4. The better performance of tungsten carbide coated specimens in comparison to the hard chromium plating may also be attributed to the lower tensile residual internal stresses in the coating of the former, compressive residual stresses on the substrate surface and in the subsurface due to the particles impact effect, as well as by interactions between surface substrate residual stresses. Note in the axial fatigue test that the different microcrack density between both hard chromium electroplatings did not play an important role in the specimens performance, as occurred in rotating bending fatigue tests. A comparison of Tables 1 and 2, as shown in Table 3, indicates the higher fatigue strength shown in axial fatigue test in relation to the rotating bending fatigue tests. In addition to the lower specimen dimensions, this is in accordance with the fact that the rotating bending fatigue tests are more severe as a result of the effect of the bending moment which increases the tensile stresses on surface from where, in general, the fatigue cracks grow. However, Table 3 shows a higher decrease in the fatigue strength as a function of the number of cycles in comparison to the rotating bending fatigue test results, for each level of stress. This may also be due to radial throughout thickness crack propagation to the base metal, in a direction normal to the maximum tensile stress and resulting in lower fatigue life of the specimen 15. Fig. 7 shows a typical fracture surface from the base metal, indicating that the fatigue crack nucleation started at the surface. In Fig. 8, several crack fronts that may be associated to the microcracks density originated from the plating process are represented. From Fig. 9, which represents fracture surface from a rotating bending fatigue specimen electroplated with accelerated hard chromium and tested at 871 MPa, one sees the coating homogeneity, strong adhesion substrate coating, and microcracks distributed along thickness in a radial shape. Fig. 10 shows a micrograph of a tungsten carbide thermal spray coated specimen blasted with aluminium Table 3 Fatigue strength in number of cycles in the rotating bending and axial fatigue tests Fatigue testing a Stress Rotating. bending fatigue data Ž %. Axial fatigue data Ž %. Base mat. W.C. C.H.C. A.H.C. Base mat. W.C. C.H.C. A.H.C. Ž MPa. Cycles Cycles Cycles Cycles Cycles Cycles Cycles Cycles Ž Ž Ž Ž Ž Ž Ž Ž Ž Ž Ž Ž Ž Ž Ž Ž Ž Ž a Values contained in parenthesis are the rate between the number of cycles of the coated material and the number of cycles of base metal, in percentage. a

16 120 M.P. Nascimento et al. r Surface and Coatings Technology 138 (2001) Fig. 7. Typical fracture surface from base metal. Rotating bending fatigue test. oxide. To compare, Fig. 11 indicates a coating profile without the blasting treatment. It is possible to observe that the aluminium oxide blasting is responsible for the increase in roughness at the substratercoating interface, with a consequent improvement in adhesion. The important characteristics of the hard chromium electroplating are the homogeneity of the coatings and Fig. 8. Fracture surface from 100 m of accelerated hard chromium electroplating. the excellent adhesion with the base metal, represented in Fig. 12. From Figs , it is also possible to observe that, in both cases, the base metal microstructure was not affected by the deposition process. Fig. 9. Fracture surface of samples accelerated hard chromium electroplated and tested at 871 MPa Ž500 =..

17 ( ) M.P. Nascimento et al. Surface and Coatings Technology Ž. Fig. 11. Tungsten carbide coating specimen 100. Fig. 10. Tungsten carbide coating aluminium blasted specimen Ž Abrasi e wears tests The abrasive wear resistance of HP HVOF WC coating and hard chromium plating was evaluated, and the results in terms of wear weight loss are represented in Table 4 and Fig. 13. Comparing the abrasive wear resistance, one sees the better performance of samples coated with WC, with lower wear weight loss than the hard chromium electroplated specimens. This may be attributed to the higher hardness and oxide content into the tungsten carbide coating. Coatings of high oxide content are usually harder and more wear resistant 14. However, initially, in the first 1000 cycles, an almost equivalent value of wear weight loss was observed for the tungsten carbide coating and conventional hard chromium electroplating, and a higher wear weight loss for the accelerated hard chromium. This higher hardness of tungsten carbide coating, which should result in lower wear weight loss in comparison with the conventional hard chromium electroplating, was damaged due to its higher surface roughness. With respect to both hard chromium electroplating, in the subsequent cycles, the wear weight loss of the accelerated hard chromium electroplating decreases with the increase in number of cycles in a parabolic way, resulting after cycles in lower wear weight loss than the conventional hard chromium plating. This may be explained by the through-thickness hardness variation, in accordance with Table 5, in which the microhardness data of tungsten carbide coating and conventional hard chromium electroplating is also indicated. The lower hardness on the accelerated hard chromium electroplating surface and its increase through-thickness may explain the decrease in the wear weight loss after a number of cycles. Theoretical calculations of the respective wear depth caused by abrasive wheels after cycles were 38.0 and 40.8 m for the accelerated and conventional hard chromium electroplating, respectively, and 9.50 m for the tungsten carbide thermal spray coating. Therefore, for both hard chromium electroplatings the wear depth values are associated to the hardness change site through-thickness. This may also be associated to the higher microcrack density and higher hardness in the accelerated hard chromium electroplating, Ž. Fig. 12. Hard chromium electroplated specimen 100.

18 122 ( ) M.P. Nascimento et al. Surface and Coatings Technology Table 4 Abrasive wears weight loss Ž. Abrasive wear Taber abraser Cycles Tungsten carbide Accelerated hard chromium Conventional hard chromium Ž N. Total mg mg 1000 Depth Ž m. Total mg mg 1000 Depth Ž m. Total mg mg 1000 Depth Ž m Median 1.89 mg 1000 cycles 2.71 mg 1,000 cycles 2.91 mg 1000 cycles Standard 0.75 mg 1000 cycles 2.34 mg 1000 cycles 0.88 mg 1000 cycles deviation and so in the higher amount of edges, resulting in lower fracture toughness and higher brittleness. In addition, the higher the crack density, the higher the amount of previously detached solid particles, which are suppressed in the microcracks and decrease the wear strength. This may result in micro-cutting, which is considered to be the predominant wear weight loss mechanism 16. Hard chromium electroplates with hardness of approximately Vickers were found to have the best frictional wear resistance, if the hardness was obtained as-deposited or by moderate heat treatment of harder deposit Salt spray tests The results of the corrosion testing, performed in a qualitative way, were obtained by visual inspection and by image analyser software of the specimen surface after exposure to salt spray test. Both coatings completed the tests fully corroded. Table 6, Figs. 14 and 15 show the results of the salt spray test after 24, 48 and 72 h, for all the cases. For the HP HVOF tungsten carbide coating, a better corrosion resistance was observed after sealing the application before testing. However, HP HVOF tungsten carbide coating did not quite protect the substrate against the aggressive action of salt spray environment after 72 h in the test chamber. The HP HVOF thermal spray tungsten carbide process has a high content of pores and oxides, which can be detrimental towards corrosion strength 14. In addition, it is well known that the HP HVOF tungsten carbide coating does not yield consistent thickness, which also made easy the salt spray action on sites of low thickness. On the other Fig. 13. Abrasive wear weight loss vs. number of cycles.

19 ( ) M.P. Nascimento et al. Surface and Coatings Technology Table 5 Through-thickness HV microhardness with 1 N load Microhardness-HV Coatings Surface Core Interface Accel. hard chromium Conv. hard chromium Tungsten carbide protection to the salt spray corrosion. However, here also, there was no protection of the substrate against the aggressive action of salt spray environment. This corrosion is due to the high content of pores and microcracks inherent to the process itself, that act as canals, leading the corrosive process to the substrate coating interface, and getting intensity. hand, HP HVOF tungsten carbide coating shows the better performance in comparison to the hard chromium electroplating. Bodger and co-workers 1 observed no corrosion products in tungsten carbide samples of 200 m thickness over 30 days in salt spray test. Higher nominal thicknesses of tungsten carbide lead to discontinuous micropore distributions along the thickness, which may delay the corrosive process. With respect to both hard chromium plated samples, Fig. 14 clearly shows the higher salt spray resistance of the accelerated hard chromium electroplated specimen during all tests. In relation to the accelerated hard chromium plated specimen with 49 m thickness and subjected to 48 h in salt spray environment, it can be observed that its surface showed approximately 5% corrosion products. On the other hand, in the same condition, the corrosion of the conventional hard chromium electroplated specimen was full, i.e. visually, 100% was corrosion. This experimental behaviour is related to the number of microcracks in the deposit in a way that the greater the microcrack density, the more corrosion along the sample surface and, so, better protection against corrosion 3. In spite of the higher microcrack density of the accelerated hard chromium electroplating, the surface roughness measurements indicates lower values than the conventional hard chromium electroplating, as mentioned before. In general, the corrosion resistance is related to the surface roughness of a part; i.e. the higher surface roughness, the higher the corrosion attack due to higher surface area 17. Therefore, the conventional hard chromium electroplating process yielded lower density and, consequently, deeper microcracks. It is also clear that the increase in the thickness enhanced the hard chromium 4. Conclusions The effect of tungsten carbide thermal spray coating applied by HP HVOF process and hard chromium electroplating for the rotating bending and axial fatigue tests was to decrease the fatigue strength of AISI 4340 steel. The influence is more significant in high cycle fatigue tests than in low cycle fatigue tests. The decrease of the fatigue strength was higher in chromium electroplated specimens than in tungsten carbide coated specimens. The higher rotating bending fatigue strength of the conventional hard chromium in comparison to the accelerated hard chromium electroplating, despite the higher thickness of the former, is associated with the lower microcrack density of the conventional hard chromium electroplating. A small increase in rotating fatigue strength was obtained for tungsten carbide thermal spray coated specimens blasted with aluminium oxide in comparison to samples without superficial treatment. No change in the microstructure of the base metal due to deposition process was observed for tungsten carbide thermal spray coating applied by HP HVOF process and for chromium electroplating. For axial fatigue tests, the negative influence of coatings on the fatigue strength of the steel followed the same tendency observed in rotating bending fatigue tests. Analysis of the hard chromium electroplated results revealed that the different microcracks density did not play an important role in their performance. Table 6 Results of the salt spray test in 24, 48 and 72 h Salt spray test Coating Tungs. carbide Accel. hard chrome Conv. hard chrome Time 100 m 16 m 36 m 49 m 16 m 36 m 49 m 24 h 30% 80% 10% OK 90% 70% 50% 48 h 50% 100% 30% 5% 100% 100% 100% 72 h 100% 100% 100% 100% 100% 100% 100%

20 124 ( ) M.P. Nascimento et al. Surface and Coatings Technology spect to both hard chromium electroplated samples, the results indicate clearly the higher salt spray resistance of the accelerated hard chromium electroplated specimens. Acknowledgements The authors are grateful for the support of this research by CAPES, FAPESP, EMBRAER-LIEB- HERR EDE and CTA AMR. Fig. 14. Salt spray test results for hard chromium electroplated samples after 48 h. References The wear weight loss tests showed better results for the HP HVOF tungsten carbide coating in comparison to the chromium electroplating. An initially higher wear weight loss for the accelerated hard chromium electroplating occurred, decreasing continuously with the increase in test cycles. Both coatings completed the test of 72 h with full corrosion. For the HP HVOF tungsten carbide coating, a better corrosion resistance was observed after sealing application before testing. With re- Fig. 15. Salt spray test results for tungsten carbide coated samples after 72 h. 1 B.E. Bodger, R.T.R. McGrann, D.A. Somerville, Plating and Surface Finishing Ž September J.M. Tyler, Metal Finishing Ž October A.R. Jones, Plating and Surface Finishing Ž April G. Dubpernell, F.A. Lowenheim, Modern Electroplating Ž H.S. Kuo, J.K. Wu, J. Mater. Sci. 31 Ž K.L. Lin, C.-J. Hsu, J.-T. Chang, J. Mater. Eng. Perform. 1 Ž J. Pina, A. Dias, M. François, J.L. Lebrun, Surf. Coat. Technol. 96 Ž S. Hotta, Y. Itou, K. Saruki, T. Arai, Surf. Coat. Technol. 73 Ž S. Wang, Y. Li, M. Yao, R. Wang, J. Mats. Process. Technol. 73 Ž Y. Matsumoto, D. Magda, D.W. Hoeppner, T.Y. Kim, J. Eng. Ind. 113 Ž J.R. Hwang, C.-P. Fung, Surf. Coat. Technol. 80 Ž D.C. Bolles, Welding J., October Ž W. Coulson, E.R. Leheup, M.G. Marsh, Trans. IMF 73 Ž. 1 Ž Available from World Wide Web: http: member.aol. com englandg tsc.htm 15 H.E. Boyer, Am. Soc. Metals. Metals Park, Ohio Ž H. Wang, W. Xia, Y. Jin, Wear 195 Ž E. Budke, J. Krempel-Hesse, H. Maidhof, H. Schussler, Surf. Coat. Technol. 112 Ž D.J. Greving, E.F. Rybicki, J.R. Shadley, J. Thermal Spray Technol. 3 Ž.Ž 4 December L. Pejryd, J. Wigren, D.J. Greving, J.R. Shadley, E.F. Rybicki, J. Thermal Spray Technol. 4 Ž.Ž

21 Journal of Materials Processing Technology 152 (2004) Alternative to chromium: characteristics and wear behavior of HVOF coatings for gas turbine shafts repair (heavy-duty) Tahar Sahraoui, Nour-Eddine Fenineche, Ghislain Montavon, Christian Coddet Universite de Technologie de Belfort-Montbeliard, LERMPS, Site de Sevenans, Belfort 90010, France Received 12 February 2003; received in revised form 27 January 2004; accepted 2 February 2004 Abstract Hard chromium plating is usually used to restore to original dimensions the worn surfaces of gas turbine shafts. However, such a technology presents harmful effects on the environment and the public health and it exhibits, moreover some intrinsic technical limitations. HVOF (high-velocity oxy-fuel) thermal spraying process appears as more environmentally friendly than chromium plating process but exhibits also lower potential production costs when compared to hard chromium deposits. In such a way, HVOF process appears as an alternative to hard chromium plating for shafts repair, by reducing the frequency of maintenance operation and repair and by deferring the need to fabricate or to buy replacement parts for used engines of previous generation. The purpose of this study was to investigate and to compare microstructural properties, wear resistance, and potentials of HVOF sprayed Tribaloy -400 (T-400), Cr 3 C 2 25%NiCr and WC 12%Co coatings for a possible replacement of hard chromium plating in gas turbine shafts repair. It was shown that thermal spray coatings exhibit the adequate properties compared to electrodeposited hard chromium coatings Elsevier B.V. All rights reserved. Keywords: Coating; Carbides; Hard chromium; Coefficient of friction; Wear; Repair 1. Introduction The most severe degradation modes that gas turbine shafts have to face are friction and wear. Surface damages generated by the sliding contact with bearings limit the life of the shafts and therefore reduce their durability and reliability [1]. Hard chromium plating is usually used to restore the original dimensions of worn surfaces of the shafts, pumps and compressors [2 4]. However, during the last decade, there has been an increasing concern surrounding the processing of chromium coatings using electroplating. This is due to environmental, health and safety considerations associated with the handling, storage, and disposal of hexavalent chromium (Cr 6+ ) compounds normally used during the plating process [5 9]. Chromium ions reduced in the electrolytic process are in hexavalent state (Cr 6+ ), known to show an exacerbated carcinogenic character, among other developed pathologies. These very harmful effects, like nasal perforation have been reported since 1827 [10]. Thus, the daily average exposure Corresponding author. Tel.: ; fax: address: tahar.sahraoui@utbm.fr (T. Sahraoui). level (PEL) 1 of the operators to hexavalent chromium is today relatively low, and the new standards to come, in particular in the United States, but also in Europe, will be even more drastic. From a current PEL reactulized in 1996 of 0.1 mg m 3, the new standards would recommend a PEL of mg m 3 [11]. The effects of rejections of chromium and of rinsing effluents on the acidification, as well as on the entrophication of the grounds and waters, are particularly disastrous. This is why the environmental standards, already more strictly reinforced than a few years ago, will also become even more severe in a near future. Currently, and since 1999, the maximum limit of chromium discharge in water amounts to 2.77 ml per day and to 1.71 ml per day on average per month. The recommendations of the new standards would amount to 0.3 ml per day and to 0.2 ml per day for the daily maximum rejections and the daily average rejections per month, respectively. Moreover, the low wettability of hard chromium by water based media and oil leads also sometimes to problems related to lubricant adhesion during sliding friction conditions [12]. 1 Permissible exposure limit, daily average concentration level (8 h per day and 40 h per week) that must not be exceeded /$ see front matter 2004 Elsevier B.V. All rights reserved. doi: /j.jmatprotec

22 44 T. Sahraoui et al. / Journal of Materials Processing Technology 152 (2004) Table 1 Sale turnovers comparison for various surface treatment techniques Processes 1990 (M ) 2000 (M ) 2010 a (M ) Thermal spray process Physical vapor deposition process Chemical and electroplating processes ( hard chromium mainly) Data for years 1990 and 2000 and perspectives for year 2010 for the European Economic Community (EEC) [22]. a Estimations. Table 2 Comparison of coating manufacturing costs for thermal spraying and for hard chromium plating in the case of a component simulating a small hydraulic ram of 50 mm diameter and 500 mm length [2,23] Materials Tungsten carbide 10%Co 4%Cr ( ) Raw material cost (powder) 30 7 Energy cost (oxygen and kerosene) 7 7 Labor cost for surface preparation, spraying and component handling (20 min) Nickel chrome boron silicon hardfacing alloy ( ) Total coating cost Hard chromium plating (power + direct labor + chemical + misc + waste treatment) ( ) Various processes offer themselves as alternatives to chromium plating: i.e., trivalent chromium plating, physical vapor deposition (PVD), and thermal spraying, among others. Trivalent chromium process permits to obtain deposits of hardness, abrasion resistance and friction coefficient equivalent to those resulting from hexavalent solutions, but the behavior in regards with corrosion remains very poor [13]. Coatings manufactured implementing physical vapor deposition process exhibit very high hardness (i.e., HV) [14] and permit the improvement of the wear and friction properties [12]. However, the performance of these coatings remains in some cases limited because of their low thickness (1 10 m), their moderate load capacity and their abrupt interface [15]. Coatings deposited by thermal spraying appear also as alternatives to hard chromium plating [16], in spite of the realization of the parts in a unit way, of the limitations in minimum dimensions of the parts to be coated, and of the limitations in the minimum internal diameter of the cylinder that can be coated (i.e., Φ mini = mm) [17]. Faced with this combination of difficulties and limits found in these processes, both environment-friendly and more flexible coatings implementing thermal spraying process have been used for a few years, providing a better productability. It should be noted that the sales turnovers relating to thermal spraying process know a remarkable increase, Table 1 [18,19]. The emergence and the development of the high-velocity oxy-fuel (HVOF) thermal spraying process over the last decade has enabled the production of very dense metallic and cermets coatings, offering an effective substitute for hard chromium plating for certain jet engine components, among other components. [20 24]. With HVOF systems, the feedstock material is heated to near or above its melting point by a high-velocity combustion gas stream (resulting from continuous combustion of fuels including propane, propylene, methane or hydrogen) and deposited [23,24]. The parts actually repaired by hard chromium plating considered in this work are shafts of THM , 1202, and 1203 gas turbine engines. It has been noted that shaft repair using hard chromium plating process requires a long period of standstill [2], and the manufacturing costs are very high (up to five times, in a first approximation) compared to those of thermal spraying process [2,25], Table 2. The present study was undertaken to investigate the potential of WC 12%Co, Cr 3 C 2 25%NiCr, and Tribaloy -400 coatings manufactured by HVOF spraying to replace electrodeposited chromium. The targeted applications involve those in which the coated surfaces could be exposed to friction and wear, like sliding bearings of the shafts of the compressor rotor, Fig Materials and experiments 2.1. Coating manufacturing Experiments involved the manufacturing and characterization of coatings deposited by HVOF spraying or by electrodeposition and the quantification of their friction and wear behaviors. 2 MAN Turbomaschinen AG, GHH BORSIG, Steinbrinkstrasse 1, Oberhausen, Germany.

23 T. Sahraoui et al. / Journal of Materials Processing Technology 152 (2004) Fig. 1. THM gas turbine shafts: (a) worn surfaces; (b) hard chromium electrodeposition on shafts (bearings). The considered feedstocks were WC 12%Co (Amdry ), Cr 3 C 2 25%NiCr (Amdry 5260) and Tribaloy -400 (Amdry 19155) powders for thermal spray coatings. Table 3 presents the chemical composition and the manufacturing method of the powders. The three feedstocks exhibit a particle size distribution ranging from 11 to 45 m. Fig. 2 shows their spherical morphology. Spray coatings were manufactured implementing thermal spraying by supersonic flame. The spraying equipment used was a commercial Sulzer-Metco CDS system operating with methane and oxygen as combustion gases and using a barrel of three in length. The m thick coatings were deposited onto AFNOR 25CD4 low carbon steel samples, Table 4. The HVOF spray parameters are given in Table 5. Hard chromium coatings were realized in a private company specialized in this field following the state of the art industrial standards. The same substrate nature, i.e., AFNOR 25CD4, was considered and hard chromium coatings were about 350 m thick. 3 Sulzer-Metco AG, Rigackerstrasse 16, 5160 Wohlen, Switzerland. Table 3 Chemical composition of powders: (a) Amdry a (Tribaloy -400), (b) Amdry a 5260 (Cr 3 C 2 25%NiCr) and (c) Amdry a 1301 (WC 12%Co) (production method: agglomerated and sintered) Element wt.% (a) Amdry a Co Balance Cr 8.50 Mo Si 2.46 Ni <0.50 Fe <0.25 P <0.03 S <0.03 (b) Amdry a 5260 C 9.23 Cr Balance Ni 19.6 (c) Amdry a 1301 Co 11.4 WC Balance C (free) <0.20 Fe <2.0 a Sulzer-Metco AG, Rigackerstrasse 16, S160 Wohlen, Switzerland.

24 46 T. Sahraoui et al. / Journal of Materials Processing Technology 152 (2004) Table 4 AFNOR 25CD4 steel composition (wt.%) C 0.25 Si 0.25 Mn 0.70 Cr 1.05 Mo 0.25 Fe Balance Table 5 High-velocity flame spraying conditions Oxygen flow rate (SLPM) 420 Powder carrier gas (SLPM) 20 Fuel gas (methane) (SLPM) 180 Nozzle (in.) 3 Spray distance (mm) 300 Table 6 Mean chemical composition of the brass used in this study as counter samples Cu Balance Pb 1.9 Ni 1 Zn 8.1 After metallographic preparation consisting in sample cutting using an abrasive saw, in sample mounting in epoxy rings and in pre-polishing and polishing using diamond slurries on an automatic polishing system to enhance reproducibility, coatings were observed and image analysis was implemented to determine the coating porosity level. Coating upper surface roughnesses were measured using a mechanical stylus. Average roughness (R a ) and mean square roughness (RMS) [26] were determined 10 times on three different samples of each nature and the results were averaged. Vickers hardness values were determined on coating polished cross-sections under a load of 0.3 kg (i.e., HV 0.3 kgf ). Six measurements were performed on three different samples of each nature and the results were also averaged Coating friction and wear behavior characterization Fig. 2. Initial feedstock powder morphology: (a) T-400; (b) Cr 3 C 2 25%NiCr; (c) WC 12%Co Coating structural characterization For microstructural analyses and worn surfaces observation, scanning electron microscopy (SEM), energy dispersive spectroscopy (EDS) and X-ray diffraction (Cu K radiation; i.e., 1.54 Å wavelength) were used. Friction and wear tests were performed on the one hand using a pin-on-disc (POD) arrangement on a CSEM 4 tribometer, and on the other hand using an Amsler machine. The POD test is a model test for determining friction characteristic and wear behavior of two solid surfaces being in sliding contact, Fig. 3. The pins were coated and slided against a disc made of brass (C31600), Table 6. This material is very often used for sliding bearing purposes, among other materials [27], and was selected to perform friction 4 Centre Suisse d Electronique et de Microtechnique SA, Jaquet Droz 1, 2007 Neuchâtel, Switzerland.

25 T. Sahraoui et al. / Journal of Materials Processing Technology 152 (2004) Fig. 3. Pin-on-disc tribometer implemented in this study. and wear tests in more severe conditions compared to other materials, such as babbitts, for example. The selected normal loads applied to the pin were 5, 10 and 20 N. As a result of rotational speed and radius of the sliding path, the selected relative velocities between pins and surfaces were 0.14, 0.70 and 0.84 m s 1. The Amsler tribometer permits to determine the wear of materials by friction. Fig. 4 shows schematically the experimental configuration. The coated discs were sliding against discs made of the same brass material than the one used for the POD experiments. The forces with which the samples were pressed together during the tests were 245, 490 and 735 N. The radial velocities were 0.52 m s 1 for the coated samples (down disc) and 0.47 m s 1 for the brass discs (upper disc), inducing a sliding between the discs of 10%. In order to study the tribological behavior of coatings in severe conditions, the tests were carried out without any lubrication. The wears experienced by the samples during the tests were determined in terms of mass, by weighting each sample before and after the test, and also in terms of samples dimension evolution, by measuring the diameter of each Amsler disc and the length of each pin before and after the tests. Prior to the friction and wear tests, the coatings were ground, and the counter samples surfaces were polished. 3. Results and discussion 3.1. Coatings structural characterization Fig. 4. Amsler specimen geometry Tribaloy -400 The Tribaloy -400 is primarily made of cobalt and molybdenum. Fig. 5 illustrates the X-ray diffraction patterns of the feedstock powder and the as-sprayed coating, respectively. The coating exhibits crystallized phases which consist primarily in cobalt (cubic) solid solution and Co 3 Mo 2 Si laves phase. Regarding the peak intensities and their broadening, the authors assume also that

26 48 T. Sahraoui et al. / Journal of Materials Processing Technology 152 (2004) Table 7 Coatings characteristics Deposits Average roughness (as-deposited), R a ( m) a Average roughness (grinded), R a ( m) a Porosity level (%) Average micro-hardness (HV 0.3 kgf ) b Tribaloy ± ± ± 72 Cr 3 C 2 25%NiCr 8.38 ± ± ± 93 WC 12%Co 7.98 ± ± ± 176 Hard chromium plating 1.94 ± ± ± 51 a Average value ± standard deviation from 10 data points. b Average value ± standard deviation from six data points. there is a mixture of amorphous and extremely fine crystalline phases (nanocrystallines). These results agree with those obtained by Xiao-Xi and Zhang for HVOF sprayed Tribaloy -800 [28]. The cross-sectional view of the as-sprayed coating, Fig. 5c, shows a fairly homogeneous structure with a high packing density. The typical layered structure results from oxide stringers between the lamellae, indicating an oxidation of the molten particles. Table 7 illustrates the coating porosity content, the Fig. 5. Tribaloy -400: (a) optical micrograph of the as-sprayed coating; (b) X-ray diffraction patterns of powder; (c) X-ray diffraction patterns of the as-sprayed coating. Fig. 6. Cr 3 C 2 25%NiCr: (a) optical micrograph of the as-sprayed coating; (b) X-ray diffraction patterns of powder; (c) X-ray diffraction patterns of the as-sprayed coating.

27 T. Sahraoui et al. / Journal of Materials Processing Technology 152 (2004) upper surface average roughness and its average microhardness Cr 3 C 2 25%NiCr EDS analyses confirmed the attended chemical composition of the powder. The X-ray diffraction patterns of the feedstock powder and the as-sprayed coating are displayed in Fig. 6. A significant peak, characteristic of a substantial amount of elemental chromium (Cr), is observed on the coating and on the powder spectra. The intensities of Cr 3 C 2 peaks are similar to those of the powder, and Cr 3 C and Cr 7 C 3 peaks are not observed. However, minor Cr 3 Ni 2 peaks are observed. These results confirm that dissociation of carbides during the spraying does not significantly occur. A broad maximum in the range of θ seems to indicate an amorphous phase in the deposit. Therefore, it is considered that nickel chromium alloy and some Cr 3 C 2 carbide are melted and some amorphous phase forms, as precised by Otsubo et al. [29,30]. The Cr 3 C 2 25%NiCr coating exhibits a fairly high hardness due to Cr 3 C 2, among other characteristics, Table WC 12%Co The properties and performances of tungsten carbide cobalt coatings are attributed to a complex function of carbide size, shape and distribution, matrix hardness and toughness, and solution of carbon in the cobalt matrix. A coating should hence retain a large volume fraction of finely distributed tungsten monocarbide (WC) to achieve optimum wear properties. This is largely dependent on the Fig. 7. WC 12%Co: (a) optical micrograph of the as-sprayed coating; (b) X-ray diffraction patterns of powder; (c) X-ray diffraction patterns of the as-sprayed coating. Fig. 8. Typical evolution of the friction coefficient vs. the sliding distance.

28 50 T. Sahraoui et al. / Journal of Materials Processing Technology 152 (2004) Table 8 Evolution of the average friction coefficient of several coatings vs. the applied load and the sliding velocity for POD experiments Tribaloy -400 Cr3C2 25%NiCr WC 12%Co Hard chromium plating Sliding velocity (m s 1 ) Tests Load (N) Variability, σ/µ (%) Friction coefficient, µ ± σ Variability, σ/µ Friction coefficient, µ ± σ Variability, σ/µ (%) Friction coefficient, µ ± µ Variability, σ/µ (%) Friction coefficient, µ ± σ P ± ± ± ± P ± ± ± ± P ± ± ± ± P ± ± ± ± P ± ± ± ± µ: average value from 428 data points; σ: associated standard deviation. minimizing of decarburization of WC, which potentially occurs at the high temperatures associated with thermal spray processes [31,32]. The manufactured coatings contain a high concentration of tungsten monocarbide crystals, namely WC, as distinguished by the higher proportion of dark gray phase, Fig. 7a. X-ray diffraction analysis confirms also the presence of a larger percentage of WC, Fig. 7b and c. This result derives from the optimization of the spray parameters, which limits the decomposition process. X-ray diffraction pattern also indicates a substantial amount of W 2 C, but did not reveal substantial amounts of brittle cobalt containing-subcarbides in the coating, such as Co 3 W 3 C, Co 6 W 6 C, Co 6 W 4 C, etc. [33]. Instead, it shows a broad maximum between 40 and 48 2θ, which is very likely characteristic of microcrystalline or amorphous materials [34]. The cobalt and excess carbon are probably present in the coatings in an amorphous states. The coating is very hard and presents a low rate of porosity, Table Friction and wear behaviors Pin-on-disc experiments Fig. 8 displays a typical evolution of the friction coefficients versus the sliding distance. Table 8 synthesizes the evolution of the average friction coefficients for each coating. Whatever is the coating nature and the applied load, the average friction coefficients remain very similar, around However, a weight loss is recorded for the counter bodies of Tribaloy -400 on the pins tested under a low load (5 N). Increasing the load (i.e., from 5 to 10 and 20 N) causes significant weight losses for the brass counter bodies of WC 12%Co and Cr 3 C 2 25%NiCr coatings. This can be explained by the high hardnesses of cermet materials, which are mainly due to the shape, the quantity, and the distribution of carbides, as well as their behavior under various levels of loads. However, it should Fig. 9. Weight loss of the counter bodies rubbing against the coated pins.

29 T. Sahraoui et al. / Journal of Materials Processing Technology 152 (2004) be noted that the weight losses of counter bodies are even more significant in case of hard chromium deposits, Fig. 9. In all cases, as the testing load was increased, a tendency towards a transition in wear behavior was revealed. At low loads, the increase of friction coefficient can be explained by an adhesive mechanism which lead to surface polishing. For higher loads, the observed decreasing of friction coefficients after the initial break-in can be attributed to crushing and aggregation of the debris due to fragmentation of mating asperities, producing a third body acting as a solid lubricant [1,35]. The same mechanism is very likely noted when the sliding velocity decreases. Fig. 10 shows the scar surfaces of coatings tested for various levels of load. Pull-outs of matter are also observed on the brass counter bodies, characterized by specific widths and depths of scars for each coated pin and each condition test. The depths and widths of scar surfaces increase when the applied normal load is increased (i.e., 96 m width, maximum value) Amsler experiments The Amsler tests permit to evaluate the wear rate of the coatings by recording the coated sample weight losses, as well as by analyzing the surface damages, under various applied loads. Fig. 11 shows a typical example of Fig. 10. Optical views of the scars formed on the counter samples (brass)/hvof coated pins.

30 52 T. Sahraoui et al. / Journal of Materials Processing Technology 152 (2004) Fig. 11. Typical evolution of the moment vs. the sliding distance. Table 9 Evolution of the dimensionless moment and the wear work vs. the applied load for Amsler experiments Tests Load (N) Tribaloy -400 Cr 3 C 2 25%NiCr WC 12%Co Hard chromium plating Moment (N m) Wear work Moment (N m) Wear work Moment (N m) Wear work Moment (N m) A A A Wear work the evolution of the friction moment according to the sliding distance (i.e., 2500 m in the case of a WC 12%Co coating in this example). Such an evolution explains why, during the tests, the samples cross a grinding period, resulting in a temporary increase of the moment value because of the high friction at the mating surfaces, Table 9. A stability of the moment is reached thereafter, due to the reduction in friction of the new mating polished surfaces. When increasing the load, friction becomes increasingly higher and is correlated to a weight loss of the coatings, Fig. 12. This is due primarily to wear and fatigue of the materials. From a general point of view, the worn surfaces of coatings show a continuous transferred brass film formed on the coated surfaces, Fig. 13. The worn surfaces of WC 12%Co and Cr 3 C 2 25%NiCr coatings reveal an evidence of grain pull-out, or scratching. Hard particles (i.e., carbides) embedded in the surface layer acted as abrasive grains wearing samples and counter samples, Fig. 13c. With increasing the load, more material is removed and more important becomes the effect of the third body abrasion: higher the wear of the coating and the counter sample are. Compared to Cr 3 C 2 25%NiCr worn surfaces, the wear tracks on WC 12%Co coating surfaces show little evidence of scratching and fractures, Fig. 13b. Compared to carbide-based coatings, Tribaloy -400 coatings appear more ductile and exhibit severe wears, occurring by partial delamination of coatings and causing grooves, Fig. 13a. In comparison to the HVOF carbide-based coatings, the chromium plating was severely damaged. Significant fractures and very large and deep grooves were worn into surfaces: the electroplated chromium is more easily abraded than the carbide-based coatings, Fig. 13d. It should be emphasized that, with regard to hardness, wear and abrasion resistance, HVOF sprayed WC 12%Co coatings are superior to Tribaloy -400 and Cr 3 C 2 25%NiCr sprayed coatings, and far superior to electrodeposited hard chromium. Fig. 12. Evolution of the weight loss of the coated samples vs. the applied load.

31 T. Sahraoui et al. / Journal of Materials Processing Technology 152 (2004) Fig. 13. Typical SEM micrographs of the worn surfaces of the coatings (V = 0.52 m s 1 ).

32 54 T. Sahraoui et al. / Journal of Materials Processing Technology 152 (2004) Conclusions The results of the present study on HVOF sprayed coatings and electrodeposited chromium can be summarized as follows: From a general point of view, HVOF carbide-based coatings exhibit higher hardnesses and superior performances in wear resistance than hard chromium coatings. In comparison to chromium plating, with regard to hardness, wear and abrasion resistance, HVOF sprayed WC 12%Co coatings were far superior. In the case of shafts repair, hard chromium plating process proved to be time consuming and to involve high costs. High hardness and good wear resistance performances of WC 12%Co HVOF sprayed coatings support their candidature for the replacement of hard chromium plating in this field. Acknowledgements The authors gratefully thank F. Pendrak and O. Landemarre (LERMPS) for their very valuable help. LERMPS is a member of the Institut des Traitements de Surface de Franche-Comté (ITSFC, Surface Treatments Institute of Franche-Comté), France. ANVAR (Agence Française de l Innovation, French Agency for Innovation) is gratefully acknowledged for its financial support under grant program No. J I References [1] L. Prchlik, S. Sampath, J. Gutleber, G. Bancke, A.W. Ruff, Friction and wear properties of WC Co and Mo Mo 2 C based functionally graded materials, J. Wear 249 (2001) [2] DMN, Sonatrach, Laghouat, Algeria, Hard chromium electroplating of THM shafts, Internal Report, April [3] US Hard Chrome Alternatives Team (HCAT), Joint Group on Pollution Prevention (JG-PP), Canadian Hard Chrome Alternatives Team (C-HCAT), Validation of WC Co, WC Co Cr HVOF or Tribaloy 800 thermal spray coatings as a replacement for hard chrome plating on C-2/E-2/P-3 and C-130 propeller hubs and low pitch stop lever sleeve, joint test protocol, November 17, [4] B. Sartwell, K. Legg, P.E. 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33 T. Sahraoui et al. / Journal of Materials Processing Technology 152 (2004) [31] J. Nerz, B. Kushner, A. Rotolico, Microstructural evaluation of tungsten carbide cobalt coatings, J. Therm. Spray Technol. 1 (2) (1992) [32] J. Stokes, L. Looney, Properties of WC Co components produced using the HVOF thermal spray process, in: Proceedings of the First International Thermal Spray Conference (ITSC 2000), Montreal, May 8 11, 2000, pp [33] H. Nakahira, Y. Harada, K. Tani, Properties of tungsten carbide cermet coatings by high energy flame spraying, in: Proceedings of the International Symposium on ATTAC, Osaka, Japan, 1988, pp [34] C.J. Li, A. Ohmori, Y. Harada, Formation of an amorphous phase in thermally sprayed WC Co, J. Therm. Spray Technol. 5 (1) (1996) [35] D.M. Kennedy, M.S.J. Hashmi, Characteristics of tribological coatings applied to tool steels under dynamic abrasion test conditions, in: K.N. Strafford (Ed.), Surface and Coating Technology, Special Volume, University of South Australia, 1993.

34 Materials and Design 24 (2003) Structure and wear behaviour of HVOF sprayed Cr C NiCr and 3 2 WC Co coatings Tahar Sahraoui*, Nour-Eddine Fenineche, Ghislain Montavon, Christian Coddet Laboratoire d Etudes et de Recherches sur les Materiaux, les Plasmas et les Surfaces (LERMPS), Universite de Technologie de Belfort-Montbeliard (UTBM), Site de Sevenans, Belfort, France Received 4 December 2002; accepted 13 March 2003 Abstract Hard chrome plating is used to restore the original dimensions to worn surfaces of gas turbine shafts.however, its use is about to decrease due to some intrinsic limitations of its deposits and the toxic and carcinogenic characteristics of the hexavalent chromium.during the last decade high velocity oxy-fuel (HVOF) thermal sprayed cermet coatings play an important role in industrial applications where exceptional friction and wear resistance are required.the purpose of this study is to investigate and to compare the microstructure, wear resistance and potentials of HVOF sprayed Cr3C 2 NiCr and WC Co coatings for a possible replacement of hard chromium plating in gas turbine components repair.it has been shown that coatings exhibit high hardness with a high volume fraction of carbides being preserved during the spraying, and have different wear behaviour Elsevier Science Ltd.All rights reserved. Keywords: Cermets; Spraying; Wear 1. Introduction Carbides based cermets coatings are widely used against wear and corrosion in gas and oil industries. Their wear resistance is three to five times that of electroplated chromium, and their manufacturing costs are low w1 3x.They are often implemented using high velocity oxy-fuel (HVOF) thermal spraying process which allows to mitigate the decarburisation phenomenon occurring during the spraying.it was pointed out that the wear behaviour of coatings depends on the microstructure and the volume fraction of carbides being preserved during the deposition process w4,5x.the present study was undertaken to investigate and to compare in various test conditions properties and wear behaviour of Cr3C 2 NiCr, and WC Co coatings deposited using HVOF thermal spraying for possible applications in gas turbine shafts repair (heavy-duty) (Fig.1). *Corresponding author.tel.: q y16; fax: q address: tahar.sahraoui@utbm.fr (T.Sahraoui). 2. Materials and experiments The feedstock were Cr C 25%NiCr (Amdry 5260) 3 2 and WC 12%Co (Amdry 1301) powders.table 1 presents the composition and the production method of the powders used.coatings were carried out by thermal spray with supersonic flame (HVOF).The torch used was a commercial Sulzer-Metco CDS system operating with methane and oxygen as fuel gases and using a nozzle of 3 psi.coatings were deposited to a thickness of mm on a 25CD4 discs substrate (rim).sem and X-ray diffraction were used to characterize various features of the materials.images obtained using the optical microscope were employed in image analysis to determine the level of porosities present in the deposit. Interfacial indentation, hardness and surface roughness of coatings were also evaluated.wear tests were performed using an Amsler machine.the loads used were: 245, 490 and 735 N.The rotation speed of the coated discs was 200 RPM, whereas that of the brass discs (counter-samples) was 180 RPM (Fig.2).Prior to the wear tests the coating discs were ground, and the Brass counter-samples surfaces were polished.applied loads /03/$ - see front matter 2003 Elsevier Science Ltd.All rights reserved. doi: /s (03)

35 310 T. Sahraoui et al. / Materials and Design 24 (2003) Fig.1.Worn shafts of THM gas turbine engines (degradation of hard chromium deposit). and sliding velocities are selected according to the design table shown in Table Results and discussion 3.1. Characterization The Cr C 25NiCr powder is shown at high magni- 3 2 fication in Fig.3 and the X-ray diffraction pattern of the Cr C 25NiCr as-sprayed coating is shown in Fig A great peak characteristic of a substantial amount of elemental chromium (Cr) is both observed on the coating and powder spectrums.the intensity of Cr C 3 2 peaks are similar than that of the spraying powder, and Cr C, Cr C peaks cannot be observed.however, weak Cr Ni peaks can be observed.a broad maxima in the 3 2 range of two theta seems to indicate an amorphous phase in the deposit.hvof sprayed Cr C NiCr coatings, exhibit an appreciable hardness and high porosity (Table 3). The properties and performance of tungsten carbide coatings is attributed to a complex function of carbide size, shape and distribution, matrix hardness and toughness, and solution of carbon in the cobalt matrix.a coating must retain a large volume fraction of finely distributed tungsten monocarbide (WC) to achieve the optimum wear properties.this is largely dependent on the minimizing of decarburisation of WC, which can Fig.2.Amsler sample dimensions. Table 2 Design table and results of the Amsler wear tests Tests Load Cr3C 2 25NiCr WC 12Co (N) Moment Wear Moment Wear (N m) Work a (Nm) Work a A =10 y =10 y A =10 y =10 y A =10 y =10 y a Results are reduced dimensionless. readily occur at high temperatures associated with the thermal spray process w6x.the coatings contain a much higher concentration of tungsten monocarbide crystals WC, as distinguished by the higher proportion of dark gray phase (Fig.5).X-ray diffraction analysis confirmed the presence of a larger percentage of WC (Fig.4). This result was expected due to the higher flame velocity and lower flame temperature of the HVOF process, which would limit the decomposition process.x-ray diffraction also indicated a substantial amount of W C, 2 but did not reveal substantial amounts of cobalt containing-subcarbides in the deposits as Co W C, Co W C, Co W C, etc.the results of chemical analysis indicated 6 4 Table 1 Chemical composition and production method of powders (a) Cr C 25NiCr, (b) WC 12Co 3 2 (a) Cr C 25NiCr 3 2 Powder Production method Element C Cr Ni Amdry 5260 Agglomerated and sintered Wt.% 9.23 Balance 19.6 (b) WC 12Co Powder Production method Element Co WC C (free) Fe Amdry 1301 Agglomerated and sintered Wt.% 11.4 Balance

36 T. Sahraoui et al. / Materials and Design 24 (2003) Fig.3.SEM observations of powders: (a) Cr C 25NiCr powder; (b) WC Co powder. 3 2 porosity.the indentation tests in interfaces revealed a great resistance which results in a good adherence of the deposit and the substrate (Table 3) Friction and wear tests Fig.4.X-ray diffraction pattern of the as-sprayed coatings. the presence of elemental cobalt (Co) in the coating (as expected), but X-ray diffraction did not reveal any substantial cobalt-containing phases.instead, it showed a broad maxima in the two theta range, which is characteristic of microcrystalline or amorphous materials.the cobalt and excess carbon is probably present in the coatings in an amorphous state.the formed deposit is harder than Cr C 25NiCr deposit and presents a low 3 2 The Amsler tests have led to evaluate the wear rate of coatings by means of coated samples weight loss as by the analysis of the surface damage under various applied loads.fig.6 shows the evolution of the moment according to the sliding distance which was 2500 m. This evolution is the same for the two materials and permit to explain why during all tests, the samples cross a grinding period which results in a temporary increase of the moment value because of the high friction of the mating surfaces.a stability of the moment is recorded thereafter, due to the reduction in friction of the new surface qualities.with the increase of load, friction becomes increasingly significant and is accompanied by a light weight loss of deposits (Fig.7).This is due primarily to wear and fatigue of the two coating types. The worn surfaces (after the Amsler tests) are shown at high magnification in Fig.8.Micrographs reveal that on the HVOF sprayed Cr C 25NiCr there is an evidence 3 2 of particle pull-out or scratching which supports the wear by abrasion.damaged surfaces contain craters whose rate and dimension are more significant than those caused in WC Co deposits.the deposit damage appears more ductile in case of Cr C 25NiCr than in 3 2 WC Co.This is mainly due to the difference in hardness of the two materials.a matter transfer caused by the Table 3 HVOF thermal sprayed coatings characteristics Deposits Roughness Roughness Porosity Microhardness (as-deposited) (ground) (%) (HK) R a (mm) R a (mm) Cr3C 2 25%NiCr 8.38" " "4 WC 12%Co 7.98" " "5

37 312 T. Sahraoui et al. / Materials and Design 24 (2003) Fig.5.Optical micrographs of the as-sprayed coatings: (a) Cr C 25NiCr; (b) WC Co. 3 2 Fig.6.Evolution of the moment varied with sliding distance. adhesive wear is to be noted for the two deposits.in the same context counter-samples (brass discs) knew a considerable weight loss.it is important to specify that Fig.7.Evolution of the weight loss of HVOF coated samples vs.the applied loads. the wear tests of HVOF sprayed Cr C 25NiCr and 3 2 WC 12Co coatings were carried out without lubrication and with extreme condition loadings.thus WC Co coatings exhibit good wear resistance than those of Fig.8.SEM observations of worn region of samples coated with: (a) Cr C 25NiCr; (b) WC Co. 3 2

38 T. Sahraoui et al. / Materials and Design 24 (2003) Cr3C 2 25%NiCr and appear more adequate with a good performance in case of lubrication and moderated loadings. 4. Conclusion The results of the present study on the Cr C 25NiCr 3 2 and WC Co HVOF sprayed coatings can be summarized as follows: Coatings exhibit high hardness with a high volume fraction of carbides being preserved during the HVOF spraying process. Hardness and wear resistance of the WC Co coatings were better than those of the Cr C 25NiCr coatings. 3 2 Compared to HVOF sprayed Cr C 25NiCr coatings, 3 2 WC Co coatings exhibit a low porosity. Considering the economical and ecological requirements, HVOF sprayed carbides based cermets coatings can be used for a possible replacement of hard chromium plating in gas turbine shafts repair. References w1x Rastegar F, Richardson DE.Alternative to chrome: HVOF cermet coatings for high horse power diesel engines.surf Coat Technol 1997;90: w2x Replacement of chromium plating, available from: httpyy March 27, w3x Metallisation.Application data sheet, pp.6 7.Available from: ( w4x L.Hanlin.Etude de l influence des conditions de projection thermique sur les proprietes mecaniques de revetements ˆ de cermets.these ` de doctorat, Universite de Franche Comte, France, w5x Nerz J, Kushner B, Rotolico A.Microstructural evaluation of tungsten carbide cobalt coatings.j Thermal Spray Technol 1992;1(2): w6x Li CJ, Ohmori A, Harada Y.Formation of an amorphous phase in thermally sprayed WC Co.J Thermal Spray Technol 1996;5(1):69 73.

39 Electrochimica Acta 49 (2004) Corrosion and wear behaviour of HVOF cermet coatings used to replace hard chromium L. Fedrizzi a,, S. Rossi b, R. Cristel b, P.L. Bonora b a Department of ICMMPM, University of Rome La Sapienza, Via Eudossiana 12, Rome, Italy b Department of Materials Engineering and Industrial Technologies, University of Trento, Via Mesiano 77, Trento, Italy Received 9 May 2003; accepted 6 January 2004 Available online 14 April 2004 Abstract The high velocity oxygen fuel (HVOF) powder spray process represents the state-of-the-art for thermal spray metallic coatings and can result in very dense, tightly adherent coatings with little or no oxidation during the application and low residual stresses. In particular this technology is applied in the automotive industry as well as in the textile and paper machinery and can be an interesting alternative to the use of galvanic processes with high environmental impact. Substitution of hard chromium coatings with new HVOF cermet coatings has been studied in this paper. Coatings obtained from conventional and nano-powders with the chemical composition 75Cr 3 C 2 25NiCr were applied on AISI 1045 steel used for cylinders in earth moving machines. A special apparatus was used to perform tribo-corrosion tests. The applied load ranged between 5 and 80N using a rotation speed of the sample of about 200 rpm. The counterface was an alumina block working in a block on ring configuration. Electrochemical methods were used to modify and control the degradation mechanisms and then tribo-corrosion tests were carried out under different working conditions: (a) free corrosion (no applied polarisation); (b) only lubricated wear (by applying a cathodic polarisation); (c) forced corrosion (by applying an anodic current). The behaviour of the cermet coatings was compared to that of conventional hard chromium coatings. It was very interesting to observe that nano-powder coatings displayed a markedly smaller weight loss value with respect to hard chromium and conventional HVOF coatings under all the selected working conditions. This better behaviour can be related to the lower surface roughness and to the better distribution of carbides in the metal matrix and also to the lower porosity of the coating Elsevier Ltd. All rights reserved. Keywords: HVOF coatings; Hard chromium coatings; Nano-sized powders; Tribo-corrosion; Electrochemical wear test 1. Introduction Recent studies show that 80% of the total costs for the protection of metals are related to coating applications. Organic coatings cover a large part of this percentage, but also metallic ones have a relatively big market. In fact, metallic coatings possess, together with good corrosion resistance, good aesthetics, brightness, and interesting mechanical properties such as hardness and wear resistance. Nowadays, due to the rapid technological developments, metallic coatings need, in order to maintain their competitiveness and to find new possible applications, to be envi- Corresponding author. Fax: addresses: lorenzo.fedrizzi@ing.unitn.it (L. Fedrizzi), stefano.rossi@ing.unitn.it (S. Rossi). ronmentally friendly and to enhance their properties both chemical and mechanical. In particular, galvanic processes can be considered highly polluting. Undoubtedly, from an environmental point of view, one of the most critical galvanic process is the hard chromium-plating. Hard chromium-plating technology is used today in the industry to coat different types of critical mechanical components (valves, pistons, piston rings, rods, hydraulic components) due to the good wear and corrosion resistance of this coating [1 3]. However chromium-plating causes effects on human health because of the use of substances in the galvanic process whose toxicological features have been recognised. Carcinogenic effects were observed; a common example of professional disease in the galvanic field is the nasal septum perforation. Moreover, the risk is connected to the combination of high temperatures and high humidity in the galvanic treatment area with the use of electric equipment goods /$ see front matter 2004 Elsevier Ltd. All rights reserved. doi: /j.electacta

40 2804 L. Fedrizzi et al. / Electrochimica Acta 49 (2004) Since some years, the hydraulic pistons which need repair are also coated with one of the thermal spray processes, the high pressure high velocity oxygen fuel (HP-HVOF) process. The used coatings are mainly chromium carbide in a nickel chromium metal matrix. The applied thickness of the coating, as sprayed, is usually about 150 m thick and after diamond grinding minimum 100 m. The coatings obtained in this way show a much longer life combined with a better corrosion resistance than the usual galvanic chromium-plating [4 7]. The thermal spray coating process is by far the most versatile modern surfacing method with regard to economics, range of materials, and scope of applications. The thermal spray process permits rapid applications of high-performance materials in thickness from some microns to more than 25 mm. Thermal spraying requires minimal base-metal preparation (substrate surfaces should be grit blasted at least to a near white metal finish or Swedish standard SA 2 1/2), can be applied in the field, and is a low temperature (>95 C) method compared with techniques such as weld overlay. The high velocity oxygen fuel powder spray process represents the state-of-the-art of thermal spray metallic coatings. It uses a combination of a fuel-like kerosene, which burns with large quantities of oxygen to produce a flame which combines a relatively low melting temperature of about 3000 C with an extremely high speed. This speed may reach values of 2000 m/s as opposed to about 100 m/s (flame spraying) to 1000 m/s (plasma spraying). This can result in very dense, tightly adherent coatings with little or no oxidation during the application and low residual stresses. The HVOF technique is most used for hard, abrasion resistant coatings where adhesion is the main concern, but it can be used for any material available in powder form. The low residual coating stress produced in the HVOF process allows significantly greater thicknesses to be applied compared to the plasma method, while providing lower porosity, lower oxide content, and higher coating adhesion. The powder used to produce these coatings has a typical grain size in the range m with carbide grain size of about 1 5 m. The produced coatings are corrosion resistant, if the coating thickness after the grinding process is more than 100 m. However a finishing process, like grinding, is necessary to obtain a pre-defined surface roughness then at least 150 m coating must be applied to obtain a functional coating. This makes this kind of coatings expensive for application in the industry. To reduce costs, it is necessary to reduce the consumption from the powder, the needed application time as well as the time (costs) for finishing. A reduction of the consumption of the powder and thereby the thickness is only possible if the coating is not pervious to corrosive products, causing under-corrosion. Laboratory tests have shown that fine powders, with fine carbides, are less dense than coarse powders with coarse carbides. On the other hand, it is proved that the density from the coating is increasing with increasing particle speed. The surface roughness after spraying is decreasing with increasing particle speed. In order to combine the performances of the new coatings, developed using new deposition techniques, with the request of an optimum surface finishing (very smooth and brilliant appearance) as expected in the production of pistons for earth movement machines, it is important to further develop the spray equipments in particular in the direction of the utilisation of very fine or nano-powders (nano-structured and mainly nano-sized) [8,9]. The HP-HVOF technique can solve the classic hard chromium coating difficulties to adapt the properties of the coating to special demands. In some cases the environment that the piston rods are exposed to both corrosion and wear is very demanding, for example when used off-shore and at mining. The HVOF technique permits to develop coatings for piston rods which show improved corrosion resistance and are not classified carcinogenic. The use of nano-powders will permit to obtain a very smooth surface finish so avoiding a final grinding step. The aim of this work, developed in the frame of an European Project (no. GRD ), is to analyse the tribo-corrosion behaviour of HVOF deposited coatings and to compare their behaviour to that of hard chromium ones. 2. Experimental Discs with a diameter of 30 mm and a thickness of 17 mm of AISI 1045 steel were machined from industrial production bars and used as substrate; the discs were drilled in the centre, to obtain toroidal samples. The chemical composition is shown in Table 1. Galvanic as well as HVOF coatings were then applied. Electrodeposition of hard chromium was made using a chromic acid bath containing sulphuric acid as catalyst (sulphuric acid concentration was about 1/100 of the chromic acid concentration), with a conductivity of 480 ms/cm. The temperature of the bath was about C, the current density was 15 A/dm 2 and the deposition time was about 2.5 h. The coating final roughness, in the range of m, was obtained by using a taping process and the nominal coating thickness was of about 25 m. In the case of 75Cr 3 C 2 25NiCr (80%Ni 20%Cr) powders the typical grain sizes for HVOF spraying varies in the range between 10 and 45 m. The quantity of carbide is 75%. After the spray process, the coatings are mostly finished by grinding and polishing procedures. The main aim of this work is to spray Cr 3 C 2 -based cermet powders by HVOF resulting in coating as thin as possible and within a surface roughness region, so that a grinding process will not be necessary. To reach this aim, it was necessary to use powders of grain sizes below 10 m, as the surface roughness mainly depends on the maximum particle size of the sprayed powder. The roughness will be approx-

41 L. Fedrizzi et al. / Electrochimica Acta 49 (2004) Table 1 Chemical composition of the AISI 1045 steel C (%) Mn (%) Si (%) P (%) S (%) Cr (%) Ni (%) Mo (%) Sn (%) Cu (%) N (%) Fe (%) Balance imately 10% of the diameter of the largest particles in the powder. Taking a standard Cr 3 C 2 HVOF powder for plasmadensifying, the particles will be molten. During the rapid resolidification, very fine Cr 3 C 2 grains are precipitated in the metallic matrix. Their size depends on the cooling rate. This powder is still too coarse for applications in the field of hard chromium replacement. Therefore it has to be crushed down by attritor milling. After classifying with an air classifier the powder should have the required grain size distribution. In this way it was possible to obtain a powder with a grain size distribution of 50% smaller than 5 m. For wear corrosion tests, an apparatus constructed in the laboratory was used [10,11]. The apparatus is shown schematically in Fig. 1. The disc samples for testing were mounted on a rotating transmission shaft, connected to an electrical motor. Only the circumferential area remained in contact with the testing electrolyte. A rod was used to hold the fixed alumina counterface in the shape of a parallelepiped (with the dimensions in the range of 25 mm 10 mm), and to impose the external load. A block on ring sliding type wear system was then obtained. Samples rotation speeds were 150, 200 and 250 rpm, while loads applied to the counterface ranged between 5 and 80N; the testing time was in the range of 6 15 h. The block on ring tribological system was introduced for working in a polymethylmethacrylate (PMMA) cell containing NaCl testing electrolyte (0.6 wt.% for hard chromium and 3.5 wt.% for HVOF coatings). Electrochemical characterisation, including corrosion potential (E corr ) versus time measurements and electrochemical impedance measurements, was carried out. These last data were obtained every hour in a 100 khz to 4 mhz frequency range using a PAR 273A potentiostat connected to a Solartron SI 1255 frequency response analyser. A platinum counterelectrode and an Ag/AgCl (+207 mv versus SHE) reference electrode were used. Electrochemical impedance measurements allowed to obtain polarisation resistance (R p ) trends even during tribo-corrosion tests. Corrosion tests were performed without a load application. Tribo-corrosion tests were performed under free immersion or applying an anodic current. Lubricated wear was obtained under cathodic protection. Surface degradation due to the mechano-chemical damage was also measured by calculation of samples weight loss after the tests. A high precision analytical balance was used (10 4 g). Moreover a profilometer analysis was performed on the sample surface, by using a Dektat3 instrument, in order to obtain the depth of the wear track at the end of the test. The microstructural analysis of samples, before and after wear corrosion tests, was obtained by using optical and scanning electron microscopes, on external surfaces and metallographic cross sections. 3. Results 3.1. Characterisation of hard chromium-coated samples The corrosion behaviour of hard chromium coatings is well known [12,13] and the good corrosion resistance depends on the passive behaviour of this metal. However chromium coatings are invariably porous and a thickness increase results in a cracked deposit, where cracks are caused by the formation of unstable chromium hydrides during plating. For many applications of hard chromium coatings it is more important to know the behaviour under wear condition. Fig. 1. Sketch of the apparatus used for tribo-corrosion tests.

42 2806 L. Fedrizzi et al. / Electrochimica Acta 49 (2004) In this paper, the corrosion behaviour of the chromium coatings, under wear conditions, was first studied by monitoring the corrosion potential. Such parameter can be affected by the dynamic evolution of the metal surface under mechanical damage. In fact many mechanisms can interfere with the stabilisation of the electrochemical parameter such as the continuous removal (caused by the contact with counterface) and reformation (contact with the electrolyte) of the passive chromium oxide layer on the sample surface; the accumulation of the corrosion products in some coating defects and their removal caused by the sliding counterface; the activation of localised corrosion phenomena (such as pitting) on the chromium coating; the removal of coating fragments caused by the wear process, promoting the metal substrate exposure to the electrolyte. These mechanisms cause the formation of surface areas, continuously changing, with different electrochemical activity. However, in our experiments a potential trend was always measurable, as an average value of the different local corrosion potentials, then obtaining useful information about surface evolution and degradation. As shown in Fig. 2, all the potential versus wear path curves show an almost instantaneous decrease of the corrosion potential just after load application; such a potential trend was previously observed in an other work [10]. The typical E corr values of the chromium coatings measured without any surface mechanical damage range between 200 and 350 mv Ag/AgCl [13]. When the sliding counterface is applied on the coating surface, the corrosion potential suddenly lowers to values in the range of 600 to 700 mv Ag/AgCl. This potential jump highlights the surface damage suffered by the system which passes from a passive state to an active one induced by wear. These values of E corr are typical of the active state of chromium: in fact we confirmed in laboratory that a sample of metallic chromium without the surface oxide layer goes to potential of 800 mv Ag/AgCl. Impedance data were also collected during immersion in the testing electrolyte. As discussed in our previous work [13], Nyquist impedance spectra obtained under pure corrosion conditions (no mechanical wear) mainly show a large capacitive loop (Fig. 3a). Interpretation of these experimental data by using an equivalent electrical circuit should include more than one time constant (more than one capacitive loop is present in the impedance spectra). However electrochemical impedance spectroscopy, was here mainly used to evaluate mechanical effects on the cell impedance and in particular on the total resistance (R tot ) parameter, which can be measured on the real axis by extrapolation of the impedance diagram at the lower frequencies [13,14]. These values, as a first approximation, were analysed as a polarisation resistance (R p ), which is inversely proportional to the corrosion rate. When no mechanical damage was superposed on corrosion degradation, polarisation resistance values in the range of 100 k cm 2 were obtained. When wear corrosion tests were performed, impedance spectra maintained a similar shape (Fig. 3b), but total resistance (R tot ) drastically changed, decreasing to a range between 8 and 12 k cm 2. The decrease of the polarisation resistance by about one order of magnitude, confirms, as suggested by the corrosion potential changes, that the coating surface passes from the passive state (typical of the pure corrosion tests) to the active condition of the tribo-corrosion, caused by the modifications of the chromium oxide layer. R tot was monitored during wear corrosion tests, changing some mechanical parameters such as the applied load or rotation speed. Fig. 4 shows that R tot decreases with the Fig. 2. Trend of E corr during tribo-corrosion tests of hard chromium coatings.

43 L. Fedrizzi et al. / Electrochimica Acta 49 (2004) Fig. 3. Impedance diagrams (Nyquist representation) after 3 h of immersion obtained under the following testing conditions: (a) 200 rpm and 0N; (b) 200 rpm and 5N. Fig. 4. Trend of R tot with applied load.

44 2808 L. Fedrizzi et al. / Electrochimica Acta 49 (2004) Table 2 Weight loss (g) under different working conditions Working conditions Corrosion Polarisation: OCP Rotation speed: 200 rpm Applied load: none Lubricated wear Polarisation: 1500 mv (Ag/AgCl) Rotation speed: 200 rpm Applied load: 5N Tribo-corrosion Polarisation: OCP Rotation speed: 200 rpm Applied load: 5N a Weight loss obtained measuring the corrosion current. Weight loss (g) a increase of the applied load. A similar behaviour of R tot was also observed with increasing rotational speed. This means that there is an increase of the metal activity due to the fact that higher rotation speeds and higher applied loads produce bigger material loss and higher surface stresses, reducing the repassivation process. In order to evaluate the relative importance of wear and corrosion mechanisms on the total damage produced on coated samples, weight loss measurements after corrosion or tribo-corrosion tests were carried out according to the following procedures: 1. wear corrosion test was carried out at open circuit potential (OCP) in the selected electrolyte (0.6 wt.% NaCl solution) using 200 rpm rotation speed and 5N applied load; 2. corrosion test was performed at open circuit potential in the same electrolyte without any applied load; sample rotation speed was 200 rpm in order to have a similar oxygen supply to the metal surface as under wear corrosion test; 3. lubricated wear test was obtained in the same electrolyte using 200 rpm rotation speed and 5N applied load; corrosion phenomena were inhibited by applying a cathodic potential ( 1500 mv versus Ag/AgCl). The duration of tests was 15 h in all cases. As shown in Table 2, weight loss due to tribo-corrosion at open circuit potential is about 27 mg; weight loss due to lubricated wear under cathodic polarisation is about 23 mg; weight loss due to free corrosion is only about mg. This last value was also calculated from the knowledge of the corrosion current density (obtained by impedance data), because the measurement obtained by using an analytical balance could be affected by errors due to the very small weight difference observed before and after the corrosion test. A first interesting observation is that the weight loss obtained by summing separate lubricated wear and corrosion degradation is lower than the weight loss obtained under tribo-corrosion alone. Then synergy in the combined degradation phenomenon is clearly visible and the observed difference could be higher if we consider the mechanism of lubricated wear under our working conditions. In fact surface degradation under cathodic polarisation could be affected by a possible reduction of the surface metal oxide; in this way the sample counterface contact can be modified favouring an adhesive wear. Corrosion degradation appears to be important under wear conditions, in fact chromium corrosion is activated by the continuous surface abrasion as shown by electrochemical measurements (see Fig. 3a and b). But even if the selected loads were rather small, our results indicate that wear of adhesive type [15] appears to be a major degradation mechanism Characterisation of HVOF-coated samples Some microstructural characteristics were first analysed in order to compare coatings obtained using standard powders with those obtained using smaller powders (almost nano-sized). The thickness of the samples, obtained using magnetic measurements (ASTM B 499) with a Fisher deltascope, is shown in Table 3. Conventional samples are thick, which is usual for this kind of application technology. The average thickness of these samples is in the range of 400 m with a very large dispersion. Coatings obtained with the nano-powders are thinner, with an average thickness in the range of 200 m. Nevertheless both types of samples seem to have a thickness sufficient to protect the metal substrate with a barrier effect. The average roughness (R a ), measured with a proper profilometer, is about 5.7 m for conventional Cr 3 C 2 coatings and about 2.2 m for nano-powder coatings, showing the clear improvements in surface smoothing using nano-sized powders, even if the results are not as good as requested for some particular applications. However this result is quite important because it could avoid grinding proce- Table 3 Thickness of the HVOF coatings Material Conventional 75Cr 3 C 2 25NiCr (80 20) Nano-75Cr 3 C 2 25NiCr (80 20) Minimum thickness m (S.D.=10.46) m (S.D.=13.65)

45 L. Fedrizzi et al. / Electrochimica Acta 49 (2004) dures after spraying. Grinding in fact has two negative effects: costs increase and possible cracking of the brittle coating. Another very important parameter to assess corrosion protection of the HVOF coatings is the porosity level in the deposited layer. Microscopical observations clearly show the different morphology of the produced coating as a function of the employed powder. Micrographs of the top view of the surface are shown in Fig. 5a and b. The finer microstructure of the coating obtained using nano-sized powder is well evident, whereas some pores can be observed in the coatings obtained using standard powders. Image analysis was performed on a cross section of the coating using a 100 magnification. This analysis reveals a pores percentage of about 5.5% (standard deviation of 1.49) for the conventional and in the range of 2.5% (standard deviation of 1.07) for the nano-powder coatings. The electrochemical behaviour and the resistance to tribo-corrosion were then evaluated. Unfortunately it was not possible to use the same working conditions utilised with the hard chromium coatings because of the different chemistry of the HVOF coatings which are mainly composed of hard carbides in a small percentage of metal matrix. Then, for the tribo-corrosion test a bigger external load of 80N was necessarily employed and consequently, in order to enhance the corrosion degradation too, a more concentrated NaCl solution (3.5 wt.%) was used. The trend with time of the corrosion potential, without load application, is shown in Fig. 6. The two HVOF coatings after immersion in the testing electrolyte show a more or less rapid decrease of the corrosion potential which can be related to the presence of porosity in coating. The standard coating shows a slower decrease of the corrosion potential because of the higher thickness, however this sample reaches Fig. 5. (a) Top view of the surface morphology of the HVOF standard coating. (b) Top view of the surface morphology of the nano-sized HVOF coating.

46 2810 L. Fedrizzi et al. / Electrochimica Acta 49 (2004) Fig. 6. Trend of the corrosion potential of the HVOF coatings during immersion in the testing electrolyte (3.5 wt.% NaCl) without external load. a lower plateau value (from 50 till 450 mv (Ag/AgCl)). It is possible to suppose that the higher porosity of the conventional coating allows the electrolyte to better reach the steel substrate. The corrosion potential does not reflect the passive behaviour of the noble NiCr metal matrix. This fact could be related to the many interfaces between metal matrix and carbides which can favour the presence of crevices favouring pitting attack. However in our opinion the presence of macrodefects, together with the porosity, allows the substrate to be reached by the electrolyte. Polarisation curves also confirm the absence of a clear passivity in the HVOF coatings. Anodic current tends to be relatively high and the less stable behaviour of the conventional Cr 3 C 2 coating is shown in Fig. 7. During the anodic polarisation it is possible to observe an increase of the current density which takes account for a pitting corrosion attack, less evident for the nano-coating. The polarisation curves seems to show that from a corrosion point of view the two HVOF coatings are not a perfect barrier. Electrochemical impedance measurements, performed in the same testing solution, confirm the above statements. As shown in Fig. 8, in the Nyquist representation, impedance decreases from about cm 2 to less than cm 2 after 8 h of immersion for the conventional Cr 3 C 2 coating. The starting values are almost in the same range of those measured on hard chromium coatings (Fig. 3a), but the rapid decrease of the impedance can be due to corrosion developing inside the coatings pores. In fact, at the end of the test some rusted spots were visible on the coating surface. Fig. 7. Polarisation curves of the HVOF coatings during immersion in the testing electrolyte (3.5 wt.% NaCl) without external load.

47 L. Fedrizzi et al. / Electrochimica Acta 49 (2004) Fig. 8. Trend of impedance diagrams in the Nyquist representation of the HVOF standard coating during immersion in the testing electrolyte (3.5 wt.% NaCl) without external load. Table 4 Working parameters of tribo-corrosion tests Material Sliding speed (rpm) Covered distance (m) Electrochemical conditions Conventional 75Cr 3 C 2 25NiCr E = 1.5 V vs. Ag/AgCl Free corrosion i = A/cm 2 Nano-75Cr 3 C 2 25NiCr E = 1.5 V vs. Ag/AgCl Free corrosion i = A/cm 2 Wear and corrosion behaviour was finally tested with the same apparatus shown in Fig. 1. The applied load is much higher to that used for the hard chromium (80N) because of the very high wear resistance of these coatings. The rotation speed of the sample is 200 rpm. The counterface was an alumina block working in a block on ring configuration. Tribo-corrosion tests were carried out under three different working conditions (Table 4): free corrosion (no applied polarisation); only lubricated wear (by applying a cathodic polarisation); forced corrosion (by applying an anodic current). With respect to the study carried out on the hard chromium coating, some experimental conditions were changed be- cause the tribo-corrosion behaviour is mainly affected by the mechanical properties of the HVOF coatings. In fact, to obtain (under free immersion conditions) a similar weight loss to that obtained with the hard chromium coatings (compare Tables 2 and 5) the applied load must be increased from 5 to 80N. Then in order to increase the effect of the corrosion attack, an anodic current was also applied as shown in Table 4. Moreover it was no more possible to estimate the weight loss due to pure corrosion from the impedance data because they are largely affected by the electrochemical response of the steel substrate and by corrosion mechanisms affected by the presence of crevices. The corrosion potential measured during the wear test quickly stabilised in a range between 400 and 500 mv (Ag/AgCl) which is very similar to that measured without Table 5 Tribo-corrosion data under different working conditions Sample Electrochemical conditions Weight loss (g) Volume loss (mm 3 ) Wear rate, W (m 2 ) Specific wear rate, K a (m 2 /N) Conventional 75Cr 3 C 2 25NiCr E = 1.5 V vs. Ag/AgCl Free corrosion i = A/cm Nano-75Cr 3 C 2 25NiCr E = 1.5 V vs. Ag/AgCl Free corrosion i = A/cm

48 2812 L. Fedrizzi et al. / Electrochimica Acta 49 (2004) Fig. 9. Trend of the corrosion potential of the HVOF coatings during immersion in the testing electrolyte (3.5 wt.% NaCl) during tribo-corrosion. mechanical damage. This behaviour is completely different to that shown by the hard chromium coatings (see Fig. 2) where the passive state of chromium was disturbed by the surface mechanical damage. In this case it seems that the corrosion potential is mainly affected by the coating structure (porosity and coating through defects) that by the mechanical activation of the metal matrix at the surface. The two HVOF coatings show a similar behaviour at beginning and only tend to differ after some kilometres of wear, probably because of the smaller thickness of the nano-coating which allow the steel substrate to be easily reached by the solution (Fig. 9). Table 5 shows the results obtained under all the selected working conditions. It is clear that for both coatings, tribo-corrosion under free immersion is slightly higher than under cathodic polarisation (only wear). The observed variation could range in the experimental error in particular for the conventional coating. However it is interesting to observe that the difference is always in the direction of an increase of weight loss eliminating the cathodic protection. This result indicates the presence of synergy between the chemical and mechanical damages. This is particularly true if we consider that the weight loss due to pure corrosion is almost zero. Such an effect is even more evident when applying an anodic current which favours the degradation of the coating by corroding the metal matrix around the carbides. The weight loss under forced corrosion appears to be very high. By comparing the behaviour of the nano-powder coatings with that of the conventional ones, with the same chemical composition, it is very interesting to observe that nano-powder coatings display a remarkably smaller volume loss value with respect to the conventional coatings under all the selected working conditions (Fig. 10). The better behaviour can be related to the lower surface roughness and to the better distribution of carbides in the metal matrix and also to the lower porosity of the coating. The volume loss data were obtained considering the following formula for the calculation of the composite density: d Cr3 C 2 NiCr = 0.75 d Cr3 C (0.8d Ni + 0.2d Cr ) Considering d Cr3 C 2 = 6.50 g/cm 3, d Cr = 7.19 g/cm 3, and d Ni = 8.91 g/cm 3, the density of the composite material is 7.02 g/cm 3. In this way it was then possible to evaluate the wear rate (W) by dividing the volume loss by the total distance and in particular it was possible to calculate the specific wear rate K a (m 2 /N) as K a = W F N where F N is the applied load. To be more precise in this work the specific wear rate should be considered as a specific wear corrosion rate, even if in the HVOF coatings the contribution of corrosion is almost negligible. The data presented in Table 5 are in the range of and m 2 /N and are typical for a ceramic ceramic contact. Such a calculation allowed to finally compare the tribo-corrosion behaviour of the HVOF coatings with respect to that of the hard chromium coatings. To obtain this data we selected the best performances of the hard chromium coatings studied in the previous part of this work. The working conditions are summarised in Table 6 when 22.1 mg of the hard chromium coating were lost during the test. The K a value obtained for the hard chromium coating is m 2 /N, which is quite higher with respect to that of the nano-hvof-coating ( m 2 /N). The better behaviour of the composite coating obtained by HVOF is due to the high ceramic percentage present in

49 L. Fedrizzi et al. / Electrochimica Acta 49 (2004) Fig. 10. Volume loss of the HVOF coatings under the different tribo-corrosion conditions. Table 6 Working parameters of tribo-corrosion tests of hard chromium and nano-hvof coatings Sample Sliding rate (m/s) Load (N) Time of test (h) Hard chromium Nano-HVOF the coating which offer a very good resistance to wear also because of the toughness of the metal matrix. 4. Conclusions The research involves very high benefits for the environment, as the proposed HVOF technique allows to substitute some highly polluting surface treatment technologies, such as chromium-plating, with a perfectly clean process from an environmental point of view. The replacement should bring some important benefits such as sensible reduction in wastewater pollution caused by chromium-plating processes and the increase of the performance (corrosion and wear resistance) with respect to chromium plated. Tribo-corrosion phenomena involving mechano-chemical degradation were studied using electrochemical and weight loss measurements. The apparatus used to study wear corrosion has been very effective because the combination of both electrochemical and mechanical analyses allowed to analyse the degradation mechanisms. Hard chromium degradation was found to be determined mainly by a wear of an adhesive type mechanism. But weight loss measurements clearly showed a synergistic effect due to the combined wear and corrosion degradation. Electrochemical data suggested that the corrosion rate of chromium coatings is increased by almost one order of magnitude by the mechanical damage. Degradation mechanisms of the HVOF coatings appeared to be quite different. In this case the presence of a large ceramic component in the composite coating made the corrosion degradation less important. The active passive behaviour is really important for the hard chromium coating and is no more fundamental in the case of the HVOF coating even if the metal matrix is NiCr made. The use of nano-sized powders improves the good behaviour of the conventional powders mainly because of a decrease of the interconnected porosity, a lower roughness, and a better distribution of the chromium carbides in the metal matrix. Acknowledgements The authors would like to take this opportunity to thank Dr. Susane Becker and the European Commission, which founded this work through the Competitive and Sustainable Growth Programme ( ). Project: The replacement of hard chromium coatings for mechanical components through High Pressure Nano-structure HVOF coatings, under Contract G5RD-CT Also the authors gratefully acknowledge the partners that have actively contributed to these project results: Dr. Passerini from Sapes, coordinator of the project C. and R. Bastoni from Osvat V. Cantelli from Fiat Research Centre A. Kirsten and R. Moll from Woka A. Igartua from Tekniker G. Matthaeus from Thermico A. Horsewell from Denmark

50 2814 L. Fedrizzi et al. / Electrochimica Acta 49 (2004) Technical University J. Landa from Tarabusi Antonio Forn and Josep Picas from Universitat Politècnica de Catalunya and F. Rabezzana from Metec. References [1] H. Silman, G. Isserlis, A.F. Averill, Protective and Decorative Coatings for Metals, Finishing Publications, Teddington, England, [2] K.N. Strattfford, P.K. Datta, C.G. Googan, Coatings and Surface Treatment for Corrosion and Wear Resistance, Ellis Horwood, Chichester, UK, [3] Department of Trade and Industry, Wear Resistant Surfaces in Engineering: A Guide to Their Production, Properties and Section, HMSO Books, London, UK, [4] S. Zimmermann, H. Kreye, in: Proceedings of the Ninth National Thermal Spray Conference, ASM International, Materials Park, OH, USA, 1996, p [5] K.J. Stein, B.S. Schorr, A. Marder, Wear 224 (1999) 153. [6] Barbezat, A.R. Nicoll, A. Sickinger, Wear 529 (1993) 162. [7] L. Russo, M. Dorfmann, Thermal Spraying: Current Status and Future Trends, High Temperature Society of Japan, 1995, p [8] B.H. Kear, L.E. McCandlish, Nanostr. Mater. 3 (1993) 19. [9] K. Jia, T.E. Fischer, Wear 200 (1996) 206. [10] S. Rossi, L. Fedrizzi, F. Deflorian, Werkstoffe und Korrosion 51 (2000) 552. [11] S. Rossi, L. Fedrizzi, M. Leoni, P. Scardi, Y. Massiani, Thin Solid Films 350 (1999) 161. [12] H. Silman, in: L.L. Shreir (Ed.), Corrosion, vol. 2, Newnes Butterworths, London, UK, 1979, p [13] L. Fedrizzi, S. Rossi, F. Deflorian, P.L. Bonora, in: Proceedings of the Eurocorr 98 Conference (EFC), Utrecht, The Netherlands, September, [14] J. Ross Macdonald, Impedance Spectroscopy, Wiley, New York, [15] L. Fedrizzi, S. Rossi, F. Bellei, F. Deflorian, Wear 253 (2003) 1173.

51 Wear 252 (2002) Wear characteristics of electrolytic hard chrome and thermal sprayed WC 10 Co 4 Cr coatings sliding against Al Ni bronze in air at 21 C and at 40 C P.L. Ko, M.F. Robertson National Research Council, Innovation Centre, 3250 East Mall, Vancouver, BC, Canada, V6T 1W5 Received 9 August 2001; received in revised form 25 February 2002; accepted 12 March 2002 Abstract Hexavalent chromium is carcinogenic and the disposing of solutions for electroplating chromium can create serious health and environmental hazards. Alternative methods of depositing other hard facing materials to replace the chrome hard coating used in aircraft landing gear are being sought and evaluated. One of these is the high-velocity-oxy-fuel (HVOF) thermal sprayed coating of tungsten carbide (WC). Sliding wear tests of HVOF WC 10 Co 4 Cr and electrolytic hard chrome (EHC) coatings sliding against Al Ni bronze were performed in a purpose-built multi-site reciprocating test rig in air at room temperature and in an environmental chamber at 40 C. The effects of several parameters, which included coating thickness, surface finish, sliding velocity and accumulated sliding distance, on the performance of these coatings were studied. The results found that the Al Ni bronze sustained higher wear from chrome plated rods than from HVOF coated rods. The Al Ni bronze also sustained much higher mass losses at low sliding speed than at high sliding speed. SEM and EDX revealed the existence of a uniformly distributed oxide layer on the Al Ni bronze specimens from high-speed tests. The metallographic examination also revealed substantial mass transfer of Al Ni bronze to the EHC coating and of tungsten carbide from the HVOF coating to the softer Al Ni bronze resulting in the observed weight gain on the EHC specimens and weight loss from the HVOF specimens. Overall, the mass losses of the Al Ni bronze were lower at 40 C than at room temperature and the thickness and surface finish of the coatings as tested appeared to have very little effect on the mass loss of Al Ni bronze Elsevier Science B.V. All rights reserved. Keywords: High-velocity-oxy-fuel; Electrolytic hard chrome; Al Ni bronze; Wear; Temperature effect; Material transfer 1. Introduction Hexavalent chromium is carcinogenic and the disposing of solutions for electroplating chromium can create serious health and environmental hazards [1]. Since the early 90s, there have been many studies to seek and to evaluate alternative methods of depositing other hard facing materials. A thermal sprayed tungsten carbide (WC) coating is one that has received much attention [2,3]. Voyer and Marple [4] found that for sliding against carbon materials, more porous coatings sprayed by a high power plasma spraying (HPPS) process exhibited better wear resistance than coatings produced by the high-velocity-oxy-fuel (HVOF) process. On the other hand, Slavin and Nerz [5] found very little difference between these two thermal spraying processes. They concluded that the performance of the HVOF process Corresponding author. address: pak.ko@nrc.ca (P.L. Ko). might have a slight edge over the HPPS process. Other studies also found that among several thermal-spraying processes, HVOF spraying is preferred for the manufacturing of coatings containing WC, for this process produces denser, harder and more wear-resistant coatings [6,7]. Furthermore, HVOF thermal spraying is a dry process that does not require chemical baths, and therefore, is more environmentally friendly. There is a considerable amount of published data on the friction and wear characteristics of HVOF thermal sprayed WC coatings [8 10], as well as data on the performance of these coatings in comparison with other materials and electro-chrome plated surfaces [2,11]. However, almost all of these studies were carried out at room temperature with some at high temperature [12]. None were tested at below freezing, particularly at 40 C, temperatures. The present wear test series was designed to compare the friction and wear characteristics of HVOF thermal sprayed coatings to electrolytic hard chrome (EHC) plating used in /02/$ see front matter 2002 Elsevier Science B.V. All rights reserved. PII: S (02)

52 P.L. Ko, M.F. Robertson / Wear 252 (2002) aircraft landing gear. These components are exposed to fairly extreme environments, such as temperature. In the present study, dry sliding tests were carried out in air at room temperature and at 40 C. The effects of several parameters, which include coating thickness, surface finish, sliding velocity and accumulated sliding distance, on the performance of these coatings sliding against an Al Ni bronze specimen, were studied. 2. Experimental 2.1. Reciprocating sliding wear test rig The test apparatus shown in Fig. 1, was designed and constructed in the Tribology Laboratory of the National Research Council of Canada. It consists of three identical reciprocating test devices arranged in parallel on an aluminum Fig. 1. Reciprocating sliding test rig: (a) overall view; (b) close-up view, showing the clamped rod specimen and the exposed bushing assembly.

53 882 P.L. Ko, M.F. Robertson / Wear 252 (2002) platform. Each test device is driven by an eccentric linkage connected to a common drive shaft, which is coupled to a variable speed, 1-hp dc motor and a reduction gear unit, Fig. 1(a). The eccentricity gives a stroke length (peak to peak amplitude) of 25.4 mm. The reduction gear unit provides a range for cycling rates from 25 to 175 cycles/min. The sliding assembly, which holds the rod specimen, is supported on a linear roller bearing. The stationary assembly holds a half-bushing specimen, which is placed above the sliding rod and is connected to an anchored force transducer for continuous friction force measurement during a test, Fig. 1(b). Each test unit was designed to withstand a maximum normal load of 650 N. The normal load is applied to the stationary assembly through a loading mechanism mounted underneath the test platform. The mechanism has a three to one leverage to help reduce the applied deadweight to one-third of the designated normal load of 640 N for easier handling Environmental chamber The test program includes two series of dry tests; one in room air and the other in air at 40 C. The room-air tests were conducted in an open-air environment with an ambient temperature of C. The 40 C tests were conducted in a Thermotron SE-Series Cascade Environmental Chamber, Fig. 2. The cooling/heating system uses hydro-fluoro-carbon (HFCs) refrigerants to cool the dehumidified air to a controlled temperature within the environmental chamber. The system provides a temperature range from 180 to 70 C and is capable of controlling the temperature to ±0.3 C. For this series of tests, the entire test platform and structure was housed in the environmental chamber. A sealed opening allows the shaft of the eccentric drive units to be connected to the dc motor and reduction gear unit, which are stationed outside the chamber Specimens As noted in Section 2.1 the configuration of the specimens is that of a circular rod and a half-bushing. They were machined to diametrical tolerances of 2.5 and 25 m, respectively. Two alloy steels, AerMet100 (MMS-217) and 300 M (AMS6419E), were used as the substrates for the rod specimens for both coatings. The bushing specimen material was an Al Ni bronze (AMS 4640/C63000), which has a nominal chemical composition of Cu 10% Al 5% Ni 3% Fe (wt.%) Coatings The HVOF coating used in the present studies has a nominal chemical composition of WC 10% Co 4% Cr (wt.%). Both EHC and HVOF deposited WC Co Cr were deposited directly, without an interfacial layer, on the 300 M and Aermet100 substrate materials. In this paper, these coatings are referred to as chrome and HVOF for short Test parameters In addition to the two air temperatures, the test parameters also included sliding speed ( and m/s); coating thickness (76 and 254 m ± 12.5 m after grinding) and surface finish (0.1 and 0.2 m Ra). The sliding distance was nominally 1830 m (36,000 cycles at 25.4 mm stroke length) for all tests. However, some tests were repeated with twice the sliding distance (3660 m), to evaluate its effect on the wear rates. For the second series of tests at 40 C, only the coating surface finish of 0.1 m Ra was tested Test procedures Before and after each test, the specimens were cleaned with ethanol in an ultrasonic bath, dried with compressed air, and then placed in a dessicator for a minimum of 24 h. Fig. 2. View of the cold chamber showing the complete test rig inside.

54 P.L. Ko, M.F. Robertson / Wear 252 (2002) The mass loss of a specimen was obtained by weighing the specimen before and after a test on an analytical balance with a resolution of ±0.1 mg for the coated rod specimens, and another analytical balance with a higher resolution to ±10 g for the lighter bronze specimens. Both analytical balances are housed in an environment-controlled room. After the specimens were installed, deadweight load was applied to the stationary Al Ni bronze specimen by way of a loading mechanism described in Section 2.1. The dc motor was started and run up to the designated sliding speed. The force transducer monitored the frictional resistance between the sliding rod and the stationary bushing at regular intervals. The data was stored in a computer and analysed. A maximum friction force limit and the designated number of test cycles were set at the computer controller, which could automatically stop the dc motor and terminate the test when either the set force limit was exceeded by the friction force or when the set number of cycles has been reached. If the test was interrupted due to high friction before completing the required number of cycles, the unaffected test sites were restarted manually to complete the remainder of the test cycles. The screen of the control computer shown in Fig. 3 displays simultaneously the force/time traces of the triplicate-tests and the control functions, which include the friction limit, the run time limit, and the number of cycles for the test. It also displays the instantaneous read out of the maximum friction forces at the three test sites, the elapsed time and the number of cycles run. For the cold temperature tests, the chamber temperature was allowed to stabilize at the operating temperature before a test is started. Thermocouples were installed on the top surface of each of the stationary specimen assemblies. The signals of these thermocouples were monitored at regular intervals. A fourth thermocouple was attached to the test platform for temperature control. At the end of a test, the chamber was allowed to warm up to 20 C before it is opened for the removal of the specimens. 3. Results 3.1. Wear The wear characteristics of the HVOF WC 10 Co 4 Cr and the EHC coatings versus the Al Ni bronze were evaluated in terms of the designated test parameters, namely: temperature, coating thickness, coating finish, sliding distance and sliding velocity. The mass loss results used in the evaluation and presented in this paper were the averaged mass loss rates from three replicate-tests performed simultaneously. These triplicate-tests have been found to be consistent with very little deviation in the mass loss results. As noted earlier, two alloy steels were used as rod substrates for the coated specimens. Thus, two identical sets of tests were performed; one with the Aermet 100 for the rod substrate and the other with the 300 M. The results show that these alloy steels, when used as the base material for Fig. 3. Screen of the control computer showing example of the force traces and the control functions.

55 884 P.L. Ko, M.F. Robertson / Wear 252 (2002) Fig. 4. Mass losses on test coatings and counter surface bushings (Al Ni bronze) in room air: (a) reciprocating sliding speed, m/s; (b) reciprocating sliding speed, m/s. the coated rods and tested in the environments used in the present study, have no effect on the mass losses of the coatings and the Al Ni bronze. The two sets of results are almost identical. Subsequently, the corresponding data from these two sets of results were summed and averaged for presentation in this paper without further identifying the substrate materials. Furthermore, preliminary results of the room temperature tests (not shown) also reveal that the designated thickness (76 and 254 m) and surface finish (0.1 and 0.2 m Ra) of the chrome coating have little to no effect at all on the mass loss of Al Ni bronze. While the respective results of the HVOF coating tests are less consistent and do show some slight variations, particularly with the thinner of the two thicknesses, nevertheless the difference in mass loss from the two surface finishes is very small. For the room temperature results, therefore, the corresponding data from the two surface finishes were summed and averaged. In the case of 40 C tests, only the specimens with 0.1 m Ra surface finish were tested. In the following, the effects of speed of sliding, temperature and coating thickness on the wear of Al Ni bronze and its counterpart coating are presented At room temperature For both chrome and HVOF coatings, the mass loss of the Al Ni bronze at the low sliding speed (25.4 mm/s), Fig. 4(a), is about two to three times higher than the corresponding mass loss at the high sliding speed (76.2 mm/s), Fig. 4(b). While the thickness of the chrome coating (254 and 76 m) caused no difference in the mass loss of the Al Ni bronze, the thicker HVOF coating caused substantially higher mass losses of the Al Ni bronze than did the thinner HVOF coating. The results in Fig. 4 also reveal that there are distinct differences in the wear characteristics of EH chrome and HVOF WC 10 Co 4 Cr coatings. The chrome plated rods caused higher wear on the Al Ni bronze than the HVOF coated rods did. On the other hand, the chrome plated rod specimens show weight gains (negative on the bar charts), whereas the HVOF coated specimens show some weight losses At 40 C The mass-loss results of the coated rods versus Al Ni bronze in air at 40 C are shown in Fig. 5. In general, the wear characteristics of these combinations at 40 C are very similar to those observed in the room air tests. That is, the chrome coatings caused higher wear on the Al Ni bronze than that caused by the HVOF coatings; the chrome plated rods had weight gains while the HVOF coated rods had weight losses. Again, as in the room temperature tests, the mass loss rates of the Al Ni bronze were higher at the low sliding speed. With the exception of sliding against the chrome coating at the higher sliding speed, the mass loss rates of the Al Ni bronze were lower at 40 C than at room temperature. On the other hand, the mass loss rates of the HVOF coated rods were much higher at 40 C than at room temperature. In fact, they were almost as high as the mass loss rates of their Al Ni bronze counter parts. At the same time, the weight gains of the chrome plated rods also increased at 40 C Effects of temperature, sliding distance and sliding speed The bar charts in Figs. 4 and 5 show that the results of the 76 and 254 m thick coatings, with the exception of the HVOF coatings at room temperature, which showed a 20 40% difference in mass losses of the Al Ni bronze, are very close, that is within ±15%. The corresponding data for the two coating thicknesses were summed, averaged (that is B1 and B2, and R1 and R2 in each sector of Figs. 4 and 5), and plotted in terms of temperature and types of coating

56 P.L. Ko, M.F. Robertson / Wear 252 (2002) Fig. 5. Mass losses on test coatings and counter surface bushings (Al Ni bronze) in air at 40 C: (a) sliding speed, m/s; (b) sliding speed, m/s. Fig. 6. Effects of sliding speed and temperature on mass loss rates of coating and Al Ni bronze. in Fig. 6. The bar charts in Fig. 6 serve to illustrate more clearly the effects of temperature and sliding speed noted in the previous sections. To further investigate the effect of sliding speed on the mass loss rates, four additional tests with HVOF rods sliding against Al Ni bronze at speeds of , , and m/s were carried out at 40 C. The graphs in Fig. 7 show a definitive trend of decreasing mass loss rates with increasing sliding speeds within the speed range tested. The bar charts in Fig. 8(a) for 40 C and Fig. 8(b) for room temperature show that the corresponding mass loss rates after 1830 m (36,000 cycles) and 3660 m (72,000 cycles) are very close. The results of the 3660 m tests further substantiate the wear characteristics of the chrome and HVOF coatings and the effect of sliding speed on the mass loss of Al Ni bronze as noted in Sections and Friction characteristics in room air and at 40 C Fig. 9(a) and (b) show the friction characteristics of the chrome and HVOF coated rods sliding against Al Ni bronze Fig. 7. Effect of sliding speed on mass loss rates of HVOF coating vs. Al Ni bronze at 40 C. in air at room temperature and at 40 C, respectively. In room air, the coefficients of friction are in the range of It is believed that the friction between the chrome/al Ni bronze pairs was primarily due to contacts between Al Ni bronze and the transferred Al Ni bronze

57 886 P.L. Ko, M.F. Robertson / Wear 252 (2002) Fig. 8. Comparison of mass loss rates at two sliding lengths: (a) HVOF coating vs. Al Ni bronze at 40 C; (b) chrome coating vs. Al Ni bronze at 20 C. Fig. 9. Friction characteristics of coatings sliding against Al Ni bronze in air: (a) at 21 C; (b) at 40 C. on the coated surface. The combination of HVOF coated rods versus Al Ni bronze appears to be less stable. It is shown in a later section that unlike the chrome coated surface, which was almost completely covered by the transferred Al Ni bronze, only scattered patches of transferred Al Ni bronze material were observed on the HVOF coated surface, which may explain the undulating nature of the friction curves of Fig. 9. At 40 C, the coefficients of friction are significantly higher than the corresponding values at room temperature. Indeed, the computer records revealed some individual readings of friction coefficients from the HVOF coated pairs had exceeded unity. The curves in Fig. 9(b) show that the coefficients of friction of the HVOF coated pairs are around 0.8 from low-speed sliding but around 0.6 from high-speed sliding. The higher coefficient of friction at low sliding speed happens to coincide with the significantly higher mass losses of the HVOF under similar conditions noted in an earlier section. The coefficients of friction of the chrome plated pairs appear to be more stable. They settled down to about during the second half of the tests at both sliding speeds. 4. Metallographic examinations 4.1. Coatings Prior to testing, the characteristics of the chrome and HVOF WC 10 Co 4 Cr coatings were examined using scratch tests, during which a normal load was applied at a steady rate of 10 N/mm while the diamond tip was moving at a rate of 3.2 mm/min. For the HVOF coating, which is slightly harder than the chrome (Rc70 and Rc66, respectively), the load was steadily increased from 0 to 80 N whereas for the chrome coating, the load was increased from0to50n. The photo-micrographs in Fig. 10 show that the chrome appears more brittle, exhibiting a network of surface cracks with some long cracks emanating from the edge of the scratch. The scratch mark on the HVOF coating exhibits a smeared appearance with powdery debris scattered at the edge of the scratch. These seem to be the result of ductile micro-fracture [13]. Further examinations of the cross-sections of the chrome and HVOF coatings reveal many micro-cracks within the chrome coating. These cracks

58 P.L. Ko, M.F. Robertson / Wear 252 (2002) Fig. 10. SEM of scratch marks: (a) HVOF WC 10 Co 4 Cr; (b) chrome. Fig. 11. SEM of cross-sections of chrome and HVOF WC Co Cr coatings: (a) EH chrome coating; (b) HVOF WC 10 Co 4 Cr coating.

59 888 P.L. Ko, M.F. Robertson / Wear 252 (2002) are about m in length and <1 m wide, and are oriented perpendicular to the rod surface. In one area, a crack appears to run from the coating surface through the depth of the coating to the substrate, Fig. 11(a). The same photo-micrograph also shows good bonding between the chrome coating and the substrate. The cross-sections of HVOF coatings in Fig. 11(b) show a more coarse structure with a few cracks that are oriented parallel to the surface. The appearance of these cracks is random and rare, unlike those occurring in the chrome coating. Based on these observations, it would appear that the HVOF coating could break away in small particles or clusters of particles as illustrated in the right-hand-side micrograph of Fig. 11(b), but it may be more forgiving than the chrome coating in view of its better ductility. Under severe loading conditions the chrome could fail badly due to the brittle cracks Examination of the worn coated rods Both visual observation and metallographic examination in the SEM showed a bronze-coloured layer covering the worn area of the chrome coating. EDX examination of the worn area revealed strong peaks of Cu and Al indicating the layer is transferred material from the Al Ni bronze, Fig. 12(a). On the other hand, the worn surface of HVOF was generally bare with only small patches of transferred material. The corresponding EDX charts show some low-level peaks of Cu in some areas, Fig. 12(b), while no Cu peaks at all in others, Fig. 12(c). The micrographs of a sectioned chrome coating are shown in Fig. 13(a) and (b). The top view shows layers of transferred Al Ni bronze spread across the whole scanned area whereas the cross-section reveals the thickness of the built-up layer of Al Ni bronze, varying from 10 to 15 m. This transferred layer appears to be well adhered to the coating surface for it seems to follow the surface contour and pits. The appearance of this thick layer explains the higher weight gains on the chrome plated rods from the 40 C tests. On the other hand, the photo-micrographs of the HVOF coated rod from 40 C tests in Fig. 14(a) and (b), fail to show any transferred Al Ni bronze particles on the surface of the coating. Instead, some wear tracks and pits were observed on the worn surface of the HVOF coating. As noted earlier in Fig. 11(b), some small cracks near the surface of the coating could form particles and detach from the surface leaving behind pits. Some of these detached wear particles are shown, in the next Section, to have been embedded on the softer bronze acting as abrasive particles and causing the wear tracks on the HVOF coating. Fig. 12. EDX of worn rod surfaces: (a) EDX of worn chrome surface reveals strong Cu peaks; (b) EDX of worn HVOF WC 10% Co 4% Cr surface.

60 P.L. Ko, M.F. Robertson / Wear 252 (2002) Fig. 13. Top and sectional views of a worn chrome plated rod (tested in air at 40 C): (a) Appearance of the top surface with transferred Al Ni bronze layer; (b) Cross-sectional view, showing the Cr coating and the transferred Al Ni bronze layer on top. Fig. 14. SEM micrographs of a worn HVOF WC Co Cr surface (tested in air at 40 C) Examination of the worn Al Ni bronze specimens A number of worn Al Ni bronze specimens were sectioned and examined. No notable changes in surface hardness, for example due to work hardening, were detected. However, the SEM of cross-sections, Fig. 15(a) and (b), show that only very thin and sparsely located oxide layers exist on the surface of specimens from the low-speed tests whereas uniformly thick oxide layers (about 10 m) are observed on specimens from the high-speed tests. The corresponding EDX examinations along a line on the cross-section confirm the presence and the extent of these oxide layers on the Al Ni bronze surface. The spread (width) of the high oxide peaks on the EDX corresponds to the thickness of the observed oxide layer (dark stripes) on the SEM micrograph. Other sectioned Al Ni bronze specimens from both room temperature and 40 C tests were further examined with EDX. In the case of Al Ni bronze versus chrome, EDX only detected small traces of Cr (minor peaks) from the room temperature specimen, Fig. 16(a), while no discernable Cr peaks were found from the 40 C specimen, Fig. 16(b). EDX of the worn surfaces of the Al Ni bronze from the pairs with HVOF at room temperature are shown in Fig. 17. They display more or less the base material itself, i.e. elements of Cu, Ni, Al and Fe. No discernable traces of transferred materials from the HVOF coating were found on the specimen from a high-speed test while only small traces of transferred W, Cr, Co, C elements were found on the specimen from a low-speed test. However, examination of Al Ni bronze specimens from 40 C tests revealed strong evidence of transferred materials from the HVOF coating,

61 890 P.L. Ko, M.F. Robertson / Wear 252 (2002) Fig. 15. EDX and SEM of worn Al Ni bronze surfaces showing the oxide peaks and the corresponding oxide layers (vs. chrome coating): (a) low speed (25 mm/s) test, partial thin oxide layer and narrow peaks; (b) high speed (76 mm/s) test, thick oxide layer and peaks. Fig. 16. EDX of worn Al Ni bronze (against chrome coated rods): (a) at room temperature, very small trace of Cr; (b) at 40 C, no discernable Cr peaks. Fig. 17. EDX of worn Al Ni bronze (against HVOF WC Co Cr rods at room temperature): (a) at 76 mm/s, no trace of W; (b) at 25 mm/s, small peaks of transferred W.

62 P.L. Ko, M.F. Robertson / Wear 252 (2002) Fig. 18. EDX of worn Al Ni bronze (against HVOF WC Co Cr in air at 40 C): (a) at 76 mm/s, showing peaks of W; (b) at 25 mm/s, strong peaks of Tungsten. Fig. 19. EDX analysis of worn Al Ni bronze specimens. that is W, C, Cr and Co, Fig. 18. In particular, strong peaks of these transferred elements were found on the specimen from a low-speed test, Fig. 18(b). The atomic and elemental percentages of a spectrum from different spots on the worn surface were obtained from four Al Ni bronze specimens tested at four different sliding speeds. Taking into consideration that some spots would record a higher percentage of transferred W, Co and Cr than others from the same specimen, the results shown in Fig. 19 are those of the highest observed percentage-values from the respective specimens, except the lowest speed one, from which the average of two higher percentage numbers was used. The graph in Fig. 19 reveals a definite trend that the extent of transferred W, Co and Cr is greater when the sliding speed is lower. 5. Discussion The results have revealed significant differences of friction and wear characteristics between the EH chrome coating and the HVOF WC 10 Co 4 Cr coating when they slided against Al Ni bronze in unlubricated (dry) conditions. The Al Ni bronze appears to be readily adhered to the chrome surface, particularly in a dehumidified cold chamber environment, whereas, only mild to negligible transfer of copper was found on the HVOF. On the contrary, with the latter combination it was found that some WC Co Cr particles were transferred and embedded on the soft Al Ni bronze surface. As a result of these two different mechanisms of material transfer, there was never any apparent mass loss on the chrome coated specimens but mild to quite severe mass losses on the HVOF coating, even though the HVOF coating is much harder than the Al Ni bronze. At 40 C, in particular, the HVOF coating wore almost as much as its Al Ni bronze counterpart. It is noted earlier from the scratch tests, that the structure and coherence of the HVOF coating differs greatly from the EH chrome coating. The former is porous and, as the scratch tests and metallography of the cross-sections have shown, small particles can break away from the surface. Some of these small particles, which contain WC, became embedded on the softer Al Ni bronze surface. This transfer phenomenon has been substantiated by the EDX and SEM element maps shown in Figs. 19 and 20. These embedded hard particles acted like a polishing medium adhered to the softer Al Ni bronze surface, causing wear on the HVOF coating and high sliding friction. At the same time, the fine debris fills up the voids on the Al Ni bronze surface providing some wear protection as well as compensating for the mass loss of the Al Ni bronze resulting in lower apparent mass losses of the latter. The photo-micrograph in Fig. 13 has revealed a thick bronze layer on the chrome coated surface from the 40 C tests. It would appear that for the Al Ni bronze/chrome pairs, except at the beginning of a test, contacts were primarily between the Al Ni bronze and the transferred Al Ni bronze layer. The mass-loss results showed a distinct effect of sliding-speed on the wear of the Al Ni bronze and, to some extent, the HVOF WC coating. The change in mass loss rates of the Al Ni bronze has a reverse relationship with the change in sliding speeds. Liu and Bahadur [14] ob-

63 892 P.L. Ko, M.F. Robertson / Wear 252 (2002) Fig. 20. SEM of worn areas of Al Ni bronze specimens: (a) fine particle layer contain W Co Cr (10.36% atomic); (b) large particle rich in W Co Cr (45% atomic). served a considerable decrease in the wear rates of Cu 20% Nb from to 0.28 m/s and above. They attributed this phenomenon to the formation of a sufficiently thick CuO film on the surface, helped by the higher sliding speed. It is known that the rate of oxide formation is very sensitive to temperature. They also found substantial strain hardening of the copper surface layer at higher sliding speeds, which also helped to reduced the wear rate. On the other hand, an earlier work by Hirst and Lancaster [15], showed that there was neither sufficient strain hardening nor significant sub-surface deformation to explain the observed change in wear rate of brass on hardened steel with sliding speed. They considered a time-dependent contact area, which could be related to the surface shear rate and hence could affect the size of the transferred fragments. Following this consideration they suggested that the shear strength of welded junctions formed between sliding surfaces diminishes with increasing sliding speed. In the present studies, no change of surface hardness was detected. The photo-micrographs in Fig. 20 reveal tungsten-rich fine particles (transferred fragments) filled craters on the worn surface of Al Ni bronze from high-speed tests (Fig. 20 (a)), and large particles also rich in Tungsten from the worn surface of low-speed tests (Fig. 20 (b)). Another plausible explanations for the lower wear at higher sliding speed is the presence of an effective oxide layer as a consequence of higher surface temperature as noted in an earlier section and also by Liu and Bahadur [14]. The difference in mass losses with speed was less significant for tests at 40 C. While oxidation might be somewhat affected due to the lower interface temperature, perhaps humidity is a more significant parameter. The tests at 40 C were conducted under a dehumidified condition whereas the room air tests were not. The formation of surface film or any form of oxide film would be likely hindered in the dehumidified environment. In the case of chrome versus Al Ni bronze this would translate into higher surface adhesion and higher transfer of the latter to the chrome. In the case of HVOF versus Al Ni bronze, this could mean higher bulk material contacts and wear. At the same time, the HVOF coating might be more prone to break away into small particles at 40 C, they are then trapped and embedded in the Al Ni bronze. Preliminary tests at room temperatures revealed that several parameters, namely, substrate rod materials, coating surface finish and coating thickness, had very minor effects on the mass loss of Al Ni bronze. However, it is also noted that, for practical reasons, these parameters, that is the thickness and surface finish, were chosen with very limited ranges. Realistically, there is reason to believe that the surface finish, particularly that of the HVOF coating, could have some effect on the mass loss of Al Ni bronze and the coating itself. 6. Conclusion The Al Ni bronze sustained higher mass loss from chrome plated rods than from HVOF coated rods. The mass loss rates of Al Ni bronze against both chrome and HVOF coatings from the low sliding speed tests (25.4 mm/s) are more than doubled the corresponding mass loss rates from the high sliding speed tests (76.2 mm/s). Due to two different mechanisms of material transfer, the chrome plated rods always show weight gains whereas the HVOF WC 10 Co 4 Cr coated rods always show some weight losses, particularly heavier from tests at 40 C. Examinations of worn specimens using SEM and EDX reveal the existence of thicker and more uniformly dis-