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Influence of bond coat surface roughness on the structure of axial suspension plasma spray thermal barrier coatings — Thermal and lifetime performance
SCT-19701; No of Pages 9 Surface & Coatings Technology xxx (2014) xxx–xxx
Contents lists available at ScienceDirect
Surface & Coatings Technology journal homepage: www.elsevier.com/locate/surfcoat
Influence of bond coat surface roughness on the structure of axial suspension plasma spray thermal barrier coatings — Thermal and lifetime performance Nicholas Curry a,⁎, Zhaolin Tang b, Nicolaie Markocsan a, Per Nylén a a b
University West, Trollhättan, Sweden Northwest Mettech Corp., Vancouver, Canada
a r t i c l e
i n f o
Available online xxxx Keywords: Thermal barrier coatings Suspension plasma spray Thermal conductivity Thermo-cyclic fatigue Thermal shock
a b s t r a c t Suspension plasma spraying has become a very promising candidate for the production of strain tolerant coatings for the gas turbine industry. Under certain process conditions suspension plasma spraying (SPS) generates column-like structures in the produced coatings. While a mechanism for column formation has been suggested previously based on columns forming on surface asperities, the effect of modification of surface structures on SPS coating properties has not been investigated. In this study, the surface topography of bond coats within a TBC system were modified by the combination of polishing and surface grit blasting. Yttria stabilized zirconia coatings were deposited using an axial feed suspension plasma spray gun. The surface topography of the resultant coatings was characterized using striped light projection. Samples were tested for thermo-cyclic fatigue lifetime at 1100 °C during 1 hour cycles. Thermal shock performance was evaluated using the burner rig test and thermal conductivity evaluated using the laser flash analysis. The results indicate that columnar SPS coating microstructure is strongly influenced by surface topography. Test results suggest that control of surface topography may be an important factor to improve the performance of SPS coatings. © 2014 Elsevier B.V. All rights reserved.
1. Introduction Thermal barrier coatings (TBC's) have been in use for the protection of gas turbine hot section components for several decades [1,2]. For the last few decades the accepted industrial application methods for TBC coatings has been electron beam physical vapor deposition (EB-PVD) or plasma spraying (APS) [3]. EB-PVD has been favored for applications that receive significant thermal shock during service such as turbine blades and vanes [3]. Conventional porous APS ceramics tend to show poor lifetimes in applications of thermal shock. Applying APS coatings with segmented vertical cracks has been used industrially to generate strain tolerant plasma sprayed coatings [4,5]. However, these coatings tend to compromise thermal properties of the coating. Due to the relative cost of EB-PVD coatings, there is a great interest in developing coatings of similar columnar morphology using lower cost techniques [5]. Recent developments of the plasma spray process have allowed the feeding of sub-micron powder in suspension or even solution precursors to form coatings [6]. Such coatings have the advantage of much finer scale microstructure features compared to those of conventional APS coatings. For TBC applications the feature of interest is the
⁎ Corresponding author. Tel.: +43 664 5589579. E-mail address: [email protected] (N. Curry).
generation of coatings that form vertical cracks [7] or even truly columnar structures [8]. The formation of columnar structures during SPS deposition can be related to the atomization of suspension droplets and their interaction with the plasma jet. If process conditions and suspension parameters are controlled, atomization of the suspension can yield droplets in the range of a micron [9]. Berghaus et al. [10] proposed that if droplets generated by atomization in the plasma are smaller than a few microns in diameter, then their trajectory of deposition is influenced dramatically by the plasma flow close to the sample surface. This would result in an effective shallow impact angle for the particles. VanEvery et al. [11] further proposed that for such small droplets or particles, their deposition trajectory would result in the formation of columnar structures on surface asperities. Furthermore, differences in structure of SPS coatings have been noted when deposited on bond coats produced by different techniques with different inherent surface roughness [12]. Assuming that the proposed mechanism for column formation holds true, then the authors propose that column size (density) will be influenced by the surface topography of the substrate or bond coat surface in the case of a TBC system. In that case there is potential to control the structure of the SPS layer by manipulation of the interface on which it is to be deposited. This work presents initial spraying of SPS coatings on APS bond coats with modified surface roughness. Production of second stage coatings
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Please cite this article as: N. Curry, et al., Surf. Coat. Technol. (2014), http://dx.doi.org/10.1016/j.surfcoat.2014.08.067
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on HVAF bond coats is also presented. Coatings have been evaluated for thermo-cyclic fatigue life, thermal shock performance and thermal diffusivity measurement to evaluate the influence of coating structure on performance.
analysis was performed using striped light projection (MikroCAD, GFMesstechnik GmbH, Germany) following ISO standard 25178 for data analysis. The scanned coating area was 4.624 mm2, with the area chosen at random on the sample surface while avoiding the coating edge.
2. Experimental procedures 2.3. Thermo-cyclic fatigue 2.1. Sample production Samples were produced in two distinct stages. The first stage consisted of SPS YSZ coatings produced on APS sprayed bond coats. The second stage samples consisted of SPS YSZ coatings produced on bond coats manufactured using the HVAF method. Coatings from this study are summarized in Table 1. Substrates used in this study were Haynes 230 for thermo-cyclic fatigue test samples with the size of 50 mm × 30 mm × 5 mm. All other samples used Hastelloy X as a substrate material. Plates for microstructural analysis were 25 mm × 25 mm × 1.6 mm. Test buttons for thermal shock testing were 25 mm in diameter and 6 mm thick. Bond coats for the first stage were sprayed using the F4-MB gun (Sulzer Metco, Wohlen, Switzerland) and AMDRY 386-4 powder (Sulzer Metco, Wohlen, Switzerland). HVAF samples were sprayed using the Uniquecoat M3 gun (Richmond, USA) and AMDRY 386-1 powder. Powders were based on NiCoCrAlY system with additions of less than 0.7% hafnium and less than 1% silicon. Powder size distributions were chosen specific to the processing technique; an APS cut used for plasma coatings and VPS cut used for HVAF coatings. Before spraying of the ceramic SPS layer, plasma sprayed bond coats were subjected to some pre-treatment to alter the surface topography of the bond coat surface. Samples were polished using silicon carbide paper to remove surface roughness and attain a smooth surface. A second set of samples was polished in the same fashion and then grit blasted with 180 grit alumina media. A third set was grit blasted using the same 180 grit media without any other pre-treatment. The fourth set was left in its as sprayed state. The SPS ceramic coatings were produced using the Axial III system (Northwest Mettech Corp., Vancouver, Canada) and Nanofeed 350 suspension feed system. Deposition conditions used a mixture of hydrogen, nitrogen and argon as process gasses with a total gas flow of 250 SLPM and a total power of 86 kW. Samples were fixed to a rotating fixture with a spray distance of 75 mm and a surface speed of 470 cm/s. Deposition conditions were the same for all SPS coatings. Suspension used for sample production was an 8 wt.% yttria stabilized zirconia suspension (CrystalArc, Northwest Mettech Corp., Canada). The median particle size was 500 nm and solids loading of 10 wt.% powder in ethanol. Specimens for SEM analysis were sectioned using a diamond cutting blade and mounted in low viscosity epoxy based resin using a vacuum impregnation technique. Samples were subsequently polished using a well established method for TBC specimens. Gold sputtering was used to allow the ceramic layers to be observed in the SEM. 2.2. Surface analysis After deposition of the SPS layer the top surface of the samples was investigated for surface topographical parameters. Surface
Thermo-cyclic fatigue (TCF) testing is performed primarily to study the ability of a coating to resist high temperature oxidation and the stress of thermally grown oxide on cooling. While failure in TCF testing is driven by bond coat oxidation; the ability of the top coat to survive thermal shock and the stress of oxide growth will also determine the final lifetime of the coating [1]. TCF testing involves placing samples in a furnace at 1100 °C for a period of 1 h. After the heating period, the samples are removed from the furnace hot zone to a cool zone where they are quench cooled with a high flow of compressed air. Cooling below 800 °C occurs within 60 s with the remaining cooling taking longer. The main requirement of the test is that sample temperature is no more than 100 °C within 10 min of the cooling cycle start. After the cooling cycle is completed, the samples are returned to the hot zone to start another cycle. The process is fully automated and the samples are photographed immediately after leaving the hot zone. Failure is considered to have occurred when 20% of the ceramic surface has de-bonded. For each coating type, four samples were prepared for testing. It should be noted that TCF testing is an accelerated test with hot soak temperatures roughly 100 °C higher than expected for a metallic bond coat in service conditions. After samples had failed under TCF testing, the failed samples were mounted in a low viscosity epoxy based resin before sectioning and prepared for microstructural analysis to assess coating microstructure at failure. 2.4. Thermal shock testing Thermal shock testing allows the investigation of the ability of coatings to survive very rapid heating and cooling events. Such testing primarily checks the ability of the coating to cope with the stress of thermal expansion mismatch, sintering and thermal gradients. Unlike TCF testing, the influence of oxidation on the bond coat is of less importance. Thermal shock testing was conducted using a burner rig at GKN Aerospace Engine Systems (Trollhättan, Sweden) [2]. Samples were subjected to 75 second cycles with heating to surface temperatures of 1200 °C and rear face temperatures of between 960 °C and 980 °C. During testing coating temperatures are monitored using pyrometers facing both front and rear sides. Bond coat temperatures of approximately 1000 °C have been correlated to this test condition using samples instrumented with thermocouples. This was not carried out for the coatings within this test series. Samples are preheated before the test starts to 600 °C from the rear face of the sample with hot air guns. After starting the combustion, the rear heaters are switched to compressed air only for cooling of the back face of the sample. Samples are monitored during testing with a video recording system and pyrometer measurements. Failure is deemed to have occurred
Table 1 Experimental coating manufacturing route and layer thickness. Coating ID
Bond coat method
Surface treatment
Surface roughness Ra (μm)
Bond coat thickness (μm)
Plasma 1 Plasma 2 Plasma 3 Plasma 4 HVAF
Plasma Plasma Plasma Plasma HVAF
Polished Polished & grit blasted Grit blasted As-sprayed As-sprayed
1–2 3–4 6–8 11–12 8–9
146 161 186 183 211
μm μm μm μm μm
± ± ± ± ±
7 9 12 11 8
Please cite this article as: N. Curry, et al., Surf. Coat. Technol. (2014), http://dx.doi.org/10.1016/j.surfcoat.2014.08.067
Top coat thickness (μm) 211 201 205 220 283
μm μm μm μm μm
± ± ± ± ±
6 11 9 17 10
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TBC system. Thermal conductivity may then be calculated using the following relationship [13]:
Table 2 Material properties of TBC layers at room temperature. Material
Density (g/cm3)
Specific heat capacity (J/g·K)
Thermal diffusivity (mm2/s)
Hastelloy-X (Substance) Plasma Bond Coat HVAF bond coat Yttria stabilized zirconia
8.22 6.5 8.2 3.9–4.1
0.468 0.545 0.452 0.538
2.871 1.007 3.277 N/A
when 20% of the ceramic surface has spalled from the test sample. As samples are monitored once per revolution for the test fixture, the result is accurate to within the nearest 4 cycles of failure. Thermal shock samples were prepared from 25 mm diameter buttons 6 mm thick that were coated with the complete TBC system. Before testing, overspray was ground away from the edge of the sample and the button mounted to a carrier plate using a single spot weld. Due to constraints with testing time, only the sample on HVAF bond coat was tested in thermal shock. 2.5. Thermal conductivity analysis Thermal conductivity analysis was performed using the laser flash method that has been long accepted for analysis of coating thermal properties [3]. 10 mm diameter samples for thermal property evaluation were water-jet cut from coated plates. A graphite cover plate with an 8 mm diameter orifice is used to prevent laser light from the sample edge from influencing the signal. This also prevents measurements from the edge region where cutting may have induced microstructural changes. In this case the tested samples consist of the complete TBC system: substrate, bond coat and SPS ceramic layers. For each coating type a minimum of four samples were measured at room temperature only. Thermal diffusivity for the complete coating system was measured for the samples using a Netzsch LFA 247 (Nezsch Gerätebau GmbH, Germany). During a measurement a laser pulse is fired on the back face of the test sample. The resulting temperature increase in the rear face is measured using an infra-red detector. Thermal diffusivity is then calculated from the following formula [13]: 2 α ¼ 0:1388 L =ðtð0:5ÞÞ where α is the thermal diffusivity (mm2/s), L is the thickness of the sample and t (0.5) is the half time taken for the total temperature rise. Six laser shots were made when the sample was at a steady state condition. The result is an average for the thermal diffusivity of the complete
λ ¼ α Cp ρ where λ is thermal conductivity (W·m−1·K−1), Cp is the specific heat capacity (J/g·K), ρ is the density of the material (g/cm3) and α is the thermal diffusivity as expressed previously. As measurements are performed on a complete TBC system, the evaluation is conducted considering the sample as a 3 layer system in which the thermal conductivity of the SPS ceramic is unknown. Using Proteus LFA Analysis software (Nezsch Gerätebau GmbH, Germany), the thermal conductivity of the SPS ceramic layer can be calculated from the measured diffusivity values of the complete TBC system using measured data for the various layers presented in Table 2. Thermal diffusivity and specific heat capacity have previously been measured for the substrate and bond coat materials [15]. The specific heat capacity of the YSZ material was measured using differential scanning calorimeter DSC 404C (Nezsch Gerätebau GmbH, Germany). The influence of thermal expansion of the layers is neglected in this case as there is no significant increase in temperature during the test measurements. Thickness of all the layers within the system was measured from microstructure cross-sections of each individual plate. Thermal conductivity is then evaluated using a three layer model that accounts for thermal contact resistance between the layers and heat loss during the measurement. 3. Results 3.1. Microstructure Figs. 1–4 present the microstructure cross-section and top surface of the samples in order of increasing bond coat surface roughness. In all cases, columnar features are present within the microstructures of the SPS layers and the individual column tops can be identified from the topographic view. It should be stressed again that coatings were deposited with identical processing parameters and sample fixturing during deposition. It can be observed that the greater the surface roughness of the underlying bond coat, the wider and more irregular the columns become in the SPS layer applied above. This can be correlated to the number of columns that can be observed in the top surface images. Higher density of column tops can be observed on bond coats with lower roughness. This observation relates well to the proposed deposition mechanism suggested by VanEvery et al. [11]. It is suggested that the incoming particles (if they are in the range 1 μm in diameter) will follow the flow of the plasma parallel with the surface (bond coat) for some distance before they impact a surface asperity. Thus the individual
Fig. 1. Cross-section and top surface of Plasma 1 coating (polished) SPS coating labeled A, bond coat labeled B and substrate labeled C.
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Fig. 2. Cross-section and top surface of Plasma 2 coating (polished and grit blasted).
columns form by progressive build-up of particulates on surface peaks. Reducing surface roughness will increase the density of peaks in a specific area. As each peak is a possible initiation site for a column within the SPS coating, smoother surfaces tend to generate a greater number of columns. Rough surfaces (as with Plasma 4) have fewer peaks though they are larger in size. This results in a more uneven SPS layer dominated by larger columns. The microstructure of the SPS sample on HVAF bond coat can be observed in Fig. 5. The HVAF bond coats had a surface roughness between that of the APS as-sprayed and grit blasted coating. In this case no additional surface modification has been performed. The presence of some larger particles at the bond coat interface can be observed (arrowed). In some cases these particles appear to be the origin of the columns formed in the coating above. However the presence of such particles presents some potential problems in the coating performance. The influence of such particles has been demonstrated in both experimental and modeling work to be detrimental to the lifetime of APS TBC's due to stress generation [14]. Additionally if the particle is poorly bonded to the coating below, it presents a possible pathway for oxidation and formation of detrimental mixed oxides. A high magnification micrograph of the top surface of a column is shown in Fig. 6. The surface shows a mixture of features present within the structure. The coating is built up from a mixture of small scale splats on the order of 1–2 μm in diameter (labeled A). Present also are partially deformed droplets (labeled B) and spherical particles (labeled C). The differences in particle states may be attributed to the treatment of the particles within the plasma jet. Splats make up the majority of the structure, originating from particles treated in the plasma core. Deformed
and spherical particles are deposited from the plasma jet periphery as it passes over the surface. 3.1.1. Column measurement In order to characterize the influence of surface roughness there is an interest to know the number of columns generated in each case. However the traditional approach of counting the number of columns visually using optical or SEM microstructures is not reliable as it has proven difficult to correctly identify the boundary between adjacent columns (as observed in Figs. 1 and 2). However, the number of columns can be indicated by the surface profile of the sprayed coating. Surface analysis data from striped light projection measurements is displayed in Fig. 7. Sa is the three dimensional arithmetic mean height and Sz is the average peak to valley height. It can be noted that as the surface roughness of the bond coat increases (from left to right on the graph), so does the apparent roughness (Sa and Sz) of the SPS layer. This can be related to the larger but fewer in number columns generated as the surface roughness is increased. This can be confirmed by observations of the column tops in cross-section (Figs. 1–5). More important for describing the column size may be the use of feature parameters such as PSm which describes the mean width of profile features on the surface, the features in this case being the column tops. Fig. 8 displays the results for PSm measured for the experimental coatings. It can be observed that the rough trend of increased feature width (column width) with increasing bond coat roughness is also followed as in the previous results (see Fig. 7). An important difference in this case is that the HVAF coating shows lower feature width relative to that of standard and grit blasted coatings despite a similar level of roughness.
Fig. 3. Cross-section and top surface of Plasma 3 coating (grit blasted).
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Fig. 4. Cross-section and top surface of Plasma 4 coating (as-sprayed).
It should be noted that the use of grit blasting increases the variation in column width. This can be seen in the grit blasted coating that shows dramatically larger deviation than the as-sprayed coating on which it is based. The same deviation can be seen in the polished and grit blasted coating relative to the coating that was only polished. Fig. 9a and b compares the Plasma 1 (polished bond coat) sample surface with the HVAF sample surface respectively. The scale indicates the height difference (in microns) measured on the coating surface. It can be observed that the rougher HVAF coating has a smaller concentration of columns, though they are larger in diameter than the columns on the polished bond coat. These can be compared with the top surface images presented in Figs. 1 and 5 respectively. The surface finish of an as produced coating may be of interest as surface roughness on airfoil components is preferred to be as low as possible as it reduces temperature at surface (problematic) and reduces aerodynamic losses [1]. 3.2. Thermo-cyclic fatigue performance The results for thermo-cyclic fatigue displayed in Fig. 10 are interesting in that there is no observed trend relating column width to lifetime. Statistically speaking there is no significant difference between the tested samples apart from the coating on the polished and grit blasted surface. As stated previously, failure in the TCF test is dominated by bond coat oxidation with ceramic strain tolerance providing only a secondary influence [15]. Micrographs for all of the coatings after TCF failure are presented in Figs. 11–15. It is apparent from the micrographs that there is little
difference to single out in top coat structure to explain the difference in TCF lifetimes. A certain level of cracking is noticeable within all coatings, though this is to be expected due to the buckling failure experienced in the test similar to the failure in the EB-PVD coatings [16]. Plasma 1 samples with a polished bond displays more significant horizontal cracking (arrowed, Fig. 11) within the lower third of the SPS ceramic layer. While it's not apparent that failure resulted due to such cracks; it suggests that the region close to the bond coat may not have sufficient strain tolerance. This could be explained by the dense structure of the SPS layer close to the interface, with the column spacing not identifiable. The same dense structure can be observed in the as sprayed microstructure in Fig. 1. Overall it is apparent that oxidation is the driver of failure. It is observed from the micrographs of the failed samples (Figs. 11–14) that surface modification will in fact change the oxidation behavior of the plasma sprayed bond coat. The polished bond coats (Fig. 11) and the polished and grit blasted coatings (Fig. 12) show similar behavior. In both cases there has been a significant amount of internal oxidation during test exposure, though less in the case of the polished and grit blasted sample. Internal oxidation deep within the coating suggests that polishing the coatings before ceramic deposition may open pathways for oxygen access by exposing delaminations within the coating. Post grit blasting, as was the case in Fig. 10, may seal some of these pathways reducing the oxidation slightly. The grit blasted sample (Fig. 13) displays a different structure. There is none of the internal oxidation as seen in the other coatings. Oxides present in the bond coat layer are those that were produced during the deposition process due to in flight oxidation of the metallic powder.
Fig. 5. Cross-section and top surface of HVAF coating (as-sprayed).
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Fig. 8. Mean width of profile data for the SPS coatings.
Fig. 6. High magnification of SPS coating surface.
The grit blasting step performed in this case may have sealed any open pathways for oxygen attack due to the deformation of the coating surface. VPS/LPPS coatings are often shot peened for a similar reason [17]. Failure appears to be concentrated in the TGO though there are some locations where cracks propagated through the topcoat leaving a small amount still attached to the bond coat (arrowed). The microstructure of the as-produced bond (Fig. 14) coat resembles previously studied coating systems involving the same bond coat method after TCF exposure [15]. The layered oxide (labeled A) at asperities being observed in previous testing and reported in literature [18]. There is some internal oxidation occurring within the top half of the bond coat as oxygen penetrates along splat boundaries. There is also the presence of mixed oxides (labeled B) above the alumina TGO that are not present in the surface treated samples. Analysis of the failures is complicated by the influence of the surface modification on the oxidation process. The samples with modified interfaces tend to form a pure alumina layer whereas the as-sprayed coatings show spinel oxides as well as alumina. The same effect has been reported previously and is believed to be due to removal of seed oxides (generated during deposition) for nucleation of spinel oxides [19–21]. For the plasma sprayed bond coats, failure occurs when the TGO has reached a thickness in the range 6–10 μm. A similar thickness of TGO is
Fig. 7. Surface topography data, Sa and Sz of the experimental SPS coatings.
present in the failed HVAF bond coat samples. There are studies that propose a critical oxide thickness for failure.[22–24]. As the TGO grows in thickness, the stress level at the topcoat bond coat interface increases to a point where cracking is initiated. However the situation is complicated in this case. On smooth surfaces, such as the polished bond coat, the growth stress is minimized and critical stress is
Fig. 9. a. Plasma 1 (polished bond coat) surface topography map. b. HVAF (polished bond coat) surface topography map.
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Fig. 10. Thermo-cyclic fatigue lifetime for experimental coatings.
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Fig. 12. Cross-section of Plasma 2 coating (polished and grit blasted) after TCF testing.
influenced only by the thermal mismatch stress between TGO layer and other layers in the system [22]. As surface roughness increases then there is a greater contribution of growth stress to the overall stress state at the interface. Therefore the failure transitions from a TGO cracking to a mixed TGO/top coat failure as the roughness of the interface is increased. This transition is seen in the TCF test samples presented in this study. Stress required to generate failure may increase when the interface becomes rougher as cracks will increasingly have to be propagated through the top coat. Optimal surface roughness or preparation for SPS coatings is presently not known. Fauchais et al. proposed optimal bonding for an APS ceramic occurs where peak height is 1/2 to 1/3 the diameter of a splat [25]. If SPS coatings behave in a similar way, surface roughness requirements for SPS may be on the scale Ra = 1 μm or lower. Finally the overall TCF results can be understood based on the competing mechanisms. Smooth bond coat/top coat interface generates the lowest stress during high temperature exposure. Rough interfaces generate higher stress but crack propagation may be more difficult. Therefore the polished and grit blasted bond coat shows the lowest lifetime as it shows behavior half way between. The failed SPS coating on HVAF bond coat is displayed in Fig. 15. As with the as-sprayed microstructure (Fig. 5); the presence of a poorly
bonded interface particle can be noted in the micrograph (labeled A). The presence of some cracking within the SPS layer can be noted (labeled B) however, the failure appears to be within the TGO along the majority of the coating cross-section. It can be noted that the remaining beta phase zone within the bond coat (labeled C) is approximately 30–40 μm in width. Beta phase (NiAl) is considered to be the reservoir of aluminium for formation of alumina scale in the TGO [22]. As such there is not a significant amount of chemical lifetime left in the coating. The failure resembles that of the grit blasted plasma spray bond coat with some regions of the YSZ layer still attached to the bond coat. HVAF coatings may benefit from some form of post-treatment due to the number of larger poorly bonded particles present at the coating interface.
Fig. 11. Cross-section of Plasma 1 coating (polished) after TCF testing.
Fig. 13. Cross-section of Plasma 3 coating (grit blasted) after TCF testing.
3.3. Thermal shock testing Thermal shock testing of SPS samples produced in this study has been limited to those produced on an HVAF bond coat. The test was performed on multiple samples and continued up to 9903 cycles. At this point the testing was discontinued without sample failure. As a reference the average lifetime for a conventionally sprayed TBC system in
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Fig. 14. Cross-section of Plasma 4 coating (as-sprayed) after TCF testing. Layered oxide labeled A and mixed oxide labeled B.
Fig. 16. Cross-section of HVAF coating (as-sprayed) after thermal shock testing. Interface particle labeled A and top coat cracking labeled B.
this test is 1800 cycles. The SPS samples exceed the best lifetime by a factor of 5 without failure. The microstructure of the SPS coating with HVAF bond coat after 9903 thermal shock cycles can be seen in Fig. 16. The bond coat exhibits a well defined two phase structure with lighter shaded gamma phase and darker shaded beta phase. There is very little depletion of the beta-phase at either the bond coat interface region or at the substrate interface. This can be expected due to the total exposure time at high temperature being around 197 h and the bond coat temperature being approximately 100 °C lower compared to that in the TCF test. Two poorly bonded particles can be noted at the bond coat interface (labeled A). As discussed previously, such feature may be detrimental to long term coating lifetime. Some cracks within the SPS coating can be observed (arrowed B) though not significant enough yet to cause delamination of the ceramic. A higher magnification image of the bond coat/ceramic interface is displayed in Fig. 17 with key features labeled. The alumina layer generated at the interface during high temperature exposure can be observed together with clear gamma and beta phases within the bond coat. There are also some areas of mixed oxide present above the alumina layer. The
structure of the SPS layer close to the interface shows no obvious signs of sintering as the pore shape is still irregular. Overall the coating appears to have a significant amount of life left to it in the thermal shock test and the HVAF coating seems promising as a bond coat for the SPS TBC system. Further improvement may be achieved if the oxide formed during testing consisted of alumina only in order to reduce interface stress.
Thermal conductivity results are displayed in Fig. 18. Overall the results demonstrate that as sprayed thermal conductivities for SPS coatings are somewhat lower than the average for a conventional APS ceramic where 1 W/m∙K is more common [26]. Overall results would suggest an increase in thermal conductivity for increasing surface roughness. This can be explained by the reduction in amount of porosity due to fewer column boundaries. The coating on the standard APS bond coat does not fit the trend. However, due to the highly irregular bond coat surface, column boundaries tend to be oriented at an angle rather
Fig. 15. Cross-section of HVAF coating (as-sprayed) after TCF testing. Interface particle labeled A, topcoat cracking labeled B and beta-phase zone labeled C.
Fig. 17. High magnification image of the HVAF coating topcoat/bond coat interface after thermal shock testing.
3.4. Laser flash analysis
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Acknowledgments Thanks to Stefan Björklund for assistance with spraying of the HVAF bond coat samples. Thanks go to Wyszomir Janikowski for his work on sample evaluation. Thanks to Jönköping University for access to the LFA equipment. Thanks also to Lars Östergren and Nicholas Erb at GKN for help with the thermal shock rig testing. References
Fig. 18. Thermal conductivity data for experimental coatings.
than perpendicular to the substrate surface. This may increase the insulating effect of the boundary regions.
4. Conclusions The results of this study demonstrate that SPS coating microstructure is influenced by the topography of the surface on which it is deposited. While the microstructure influence on performance properties is masked in this case by bond coat effects; in principle SPS microstructure could be partially designed by controlling the surface topography on which it is deposited. Surface modification of APS bond coats is not recommended due to dramatic changes in oxidation behavior. However, HVAF or HVOF coatings may be suitable for further development of surface modification. The fact that SPS coatings adhere to reasonably smooth surfaces means that diffusion aluminide bond coats could be attempted as used with EB-PVD coatings. The deposition requirements for SPS coatings are not yet well defined as they are for APS or EB-PVD coatings. More work is required in order to fully understand the optimal surface requirements for an SPS TBC system to achieve the best combination of properties. In reality it is likely that the surface requirements for good bonding and lifetime are somewhat more complicated than simple factors such as Ra.
[1] F.O. Soechting, J. Therm. Spray Technol. 8 (4) (Dec. 1999) 505–511. [2] R. Miller, J. Therm. Spray Technol. 6 (1) (Mar. 1997) 35–42. [3] U. Schulz, C. Leyens, K. Fritscher, M. Peters, B. Saruhan-Brings, O. Lavigne, J.-M. Dorvaux, M. Poulain, R. Mévrel, M. Caliez, Aerosp. Sci. Technol. 7 (1) (Jan. 2003) 73–80. [4] M. Karger, R. Vaßen, D. Stöver, Surf. Coat. Technol. 206 (1) (Oct. 2011) 16–23. [5] A. Feuerstein, J. Knapp, T. Taylor, A. Ashary, A. Bolcavage, N. Hitchman, J. Therm. Spray Technol. 17 (2) (Jun. 2008) 199–213. [6] L. Pawlowski, Surf. Coat. Technol. 203 (19) (Jun. 2009) 2807–2829. [7] A. Guignard, G. Mauer, R. Vaßen, D. Stöver, J. Therm. Spray Technol. 21 (3–4) (Jun. 2012) 416–424. [8] Z. Tang, H. Kim, I. Yaroslavski, G. Masindo, Z. Celler, D. Ellsworth, Proceedings of the International Thermal Spray Conference, Hamburg, 2011. [9] J. Fazilleau, C. Delbos, V. Rat, J.F. Coudert, P. Fauchais, B. Pateyron, Plasma Chem. Plasma Process. 26 (4) (Aug. 2006) 371–391. [10] J.O. Berghaus, S. Bouaricha, J.-G. Legoux, C. Moreau, T. Chráska, Proceedings of the International Thermal Spray Conference, Basel, Switzerland, 2005. [11] K. VanEvery, M. Krane, R. Trice, H. Wang, W. Porter, M. Besser, D. Sordelet, J. Ilavsky, J. Almer, J. Therm. Spray Technol. 20 (4) (2011) 817–828. [12] N. Curry, K. VanEvery, T. Snyder, N. Markocsan, Coatings 4 (Aug. 2014) 630–650. [13] R.E. Taylor, Mater. Sci. Eng. A 245 (2) (May 1998) 160–167. [14] M. Gupta, K. Skogsberg, P. Nylén, J. Therm. Spray Technol. 23 (1–2) (Jan. 2014) 170–181. [15] N. Curry, N. Markocsan, L. Östergren, X.-H. Li, M. Dorfman, J. Therm. Spray Technol. 22 (6) (Aug. 2013) 864–872. [16] A.G. Evans, D.R. Mumm, J.W. Hutchinson, G.H. Meier, F.S. Pettit, Prog. Mater. Sci. 46 (5) (2001) 505–553. [17] J.F. Loersch, J.W. Neal, Peened Overlay Coatings, Apr 30, 1985. 4514469. [18] O. Trunova, T. Beck, R. Herzog, R.W. Steinbrech, L. Singheiser, Surf. Coat. Technol. 202 (20) (Jul. 2008) 5027–5032. [19] N. Czech, M. Juez-Lorenzo, V. Kolarik, W. Stamm, Surf. Coat. Technol. 108–109 (Oct. 1998) 36–42. [20] A. Gil, V. Shemet, R. Vassen, M. Subanovic, J. Toscano, D. Naumenko, L. Singheiser, W.J. Quadakkers, Surf. Coat. Technol. 201 (7) (Dec. 2006) 3824–3828. [21] D. Mercier, C. Kaplin, G. Goodall, G. Kim, M. Brochu, Surf. Coat. Technol. 205 (7) (Dec. 2010) 2546–2553. [22] H.E. Evans, Surf. Coat. Technol. 206 (7) (Dec. 2011) 1512–1521. [23] M. Tamura, M. Takahashi, J. Ishii, K. Suzuki, M. Sato, K. Shimomura, J. Therm. Spray Technol. 8 (1) (Mar. 1999) 68–72. [24] H. Aleksanoglu, A. Scholz, M. Oechsner, C. Berger, M. Rudolphi, M. Schütze, W. Stamm, Int. J. Fatigue 53 (Aug. 2013) 40–48. [25] P. Fauchais, R. Etchart-Salas, V. Rat, J.F. Coudert, N. Caron, K. Wittmann-Ténèze, J. Therm. Spray Technol. 17 (1) (Mar. 2008) 31–59. [26] N. Curry, J. Donoghue, Surf. Coat. Technol. 209 (Sep. 2012) 38–43.
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Contents lists available at ScienceDirect
Surface & Coatings Technology journal homepage: www.elsevier.com/locate/surfcoat
Influence of bond coat surface roughness on the structure of axial suspension plasma spray thermal barrier coatings — Thermal and lifetime performance Nicholas Curry a,⁎, Zhaolin Tang b, Nicolaie Markocsan a, Per Nylén a a b
University West, Trollhättan, Sweden Northwest Mettech Corp., Vancouver, Canada
a r t i c l e
i n f o
Available online xxxx Keywords: Thermal barrier coatings Suspension plasma spray Thermal conductivity Thermo-cyclic fatigue Thermal shock
a b s t r a c t Suspension plasma spraying has become a very promising candidate for the production of strain tolerant coatings for the gas turbine industry. Under certain process conditions suspension plasma spraying (SPS) generates column-like structures in the produced coatings. While a mechanism for column formation has been suggested previously based on columns forming on surface asperities, the effect of modification of surface structures on SPS coating properties has not been investigated. In this study, the surface topography of bond coats within a TBC system were modified by the combination of polishing and surface grit blasting. Yttria stabilized zirconia coatings were deposited using an axial feed suspension plasma spray gun. The surface topography of the resultant coatings was characterized using striped light projection. Samples were tested for thermo-cyclic fatigue lifetime at 1100 °C during 1 hour cycles. Thermal shock performance was evaluated using the burner rig test and thermal conductivity evaluated using the laser flash analysis. The results indicate that columnar SPS coating microstructure is strongly influenced by surface topography. Test results suggest that control of surface topography may be an important factor to improve the performance of SPS coatings. © 2014 Elsevier B.V. All rights reserved.
1. Introduction Thermal barrier coatings (TBC's) have been in use for the protection of gas turbine hot section components for several decades [1,2]. For the last few decades the accepted industrial application methods for TBC coatings has been electron beam physical vapor deposition (EB-PVD) or plasma spraying (APS) [3]. EB-PVD has been favored for applications that receive significant thermal shock during service such as turbine blades and vanes [3]. Conventional porous APS ceramics tend to show poor lifetimes in applications of thermal shock. Applying APS coatings with segmented vertical cracks has been used industrially to generate strain tolerant plasma sprayed coatings [4,5]. However, these coatings tend to compromise thermal properties of the coating. Due to the relative cost of EB-PVD coatings, there is a great interest in developing coatings of similar columnar morphology using lower cost techniques [5]. Recent developments of the plasma spray process have allowed the feeding of sub-micron powder in suspension or even solution precursors to form coatings [6]. Such coatings have the advantage of much finer scale microstructure features compared to those of conventional APS coatings. For TBC applications the feature of interest is the
⁎ Corresponding author. Tel.: +43 664 5589579. E-mail address: [email protected] (N. Curry).
generation of coatings that form vertical cracks [7] or even truly columnar structures [8]. The formation of columnar structures during SPS deposition can be related to the atomization of suspension droplets and their interaction with the plasma jet. If process conditions and suspension parameters are controlled, atomization of the suspension can yield droplets in the range of a micron [9]. Berghaus et al. [10] proposed that if droplets generated by atomization in the plasma are smaller than a few microns in diameter, then their trajectory of deposition is influenced dramatically by the plasma flow close to the sample surface. This would result in an effective shallow impact angle for the particles. VanEvery et al. [11] further proposed that for such small droplets or particles, their deposition trajectory would result in the formation of columnar structures on surface asperities. Furthermore, differences in structure of SPS coatings have been noted when deposited on bond coats produced by different techniques with different inherent surface roughness [12]. Assuming that the proposed mechanism for column formation holds true, then the authors propose that column size (density) will be influenced by the surface topography of the substrate or bond coat surface in the case of a TBC system. In that case there is potential to control the structure of the SPS layer by manipulation of the interface on which it is to be deposited. This work presents initial spraying of SPS coatings on APS bond coats with modified surface roughness. Production of second stage coatings
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on HVAF bond coats is also presented. Coatings have been evaluated for thermo-cyclic fatigue life, thermal shock performance and thermal diffusivity measurement to evaluate the influence of coating structure on performance.
analysis was performed using striped light projection (MikroCAD, GFMesstechnik GmbH, Germany) following ISO standard 25178 for data analysis. The scanned coating area was 4.624 mm2, with the area chosen at random on the sample surface while avoiding the coating edge.
2. Experimental procedures 2.3. Thermo-cyclic fatigue 2.1. Sample production Samples were produced in two distinct stages. The first stage consisted of SPS YSZ coatings produced on APS sprayed bond coats. The second stage samples consisted of SPS YSZ coatings produced on bond coats manufactured using the HVAF method. Coatings from this study are summarized in Table 1. Substrates used in this study were Haynes 230 for thermo-cyclic fatigue test samples with the size of 50 mm × 30 mm × 5 mm. All other samples used Hastelloy X as a substrate material. Plates for microstructural analysis were 25 mm × 25 mm × 1.6 mm. Test buttons for thermal shock testing were 25 mm in diameter and 6 mm thick. Bond coats for the first stage were sprayed using the F4-MB gun (Sulzer Metco, Wohlen, Switzerland) and AMDRY 386-4 powder (Sulzer Metco, Wohlen, Switzerland). HVAF samples were sprayed using the Uniquecoat M3 gun (Richmond, USA) and AMDRY 386-1 powder. Powders were based on NiCoCrAlY system with additions of less than 0.7% hafnium and less than 1% silicon. Powder size distributions were chosen specific to the processing technique; an APS cut used for plasma coatings and VPS cut used for HVAF coatings. Before spraying of the ceramic SPS layer, plasma sprayed bond coats were subjected to some pre-treatment to alter the surface topography of the bond coat surface. Samples were polished using silicon carbide paper to remove surface roughness and attain a smooth surface. A second set of samples was polished in the same fashion and then grit blasted with 180 grit alumina media. A third set was grit blasted using the same 180 grit media without any other pre-treatment. The fourth set was left in its as sprayed state. The SPS ceramic coatings were produced using the Axial III system (Northwest Mettech Corp., Vancouver, Canada) and Nanofeed 350 suspension feed system. Deposition conditions used a mixture of hydrogen, nitrogen and argon as process gasses with a total gas flow of 250 SLPM and a total power of 86 kW. Samples were fixed to a rotating fixture with a spray distance of 75 mm and a surface speed of 470 cm/s. Deposition conditions were the same for all SPS coatings. Suspension used for sample production was an 8 wt.% yttria stabilized zirconia suspension (CrystalArc, Northwest Mettech Corp., Canada). The median particle size was 500 nm and solids loading of 10 wt.% powder in ethanol. Specimens for SEM analysis were sectioned using a diamond cutting blade and mounted in low viscosity epoxy based resin using a vacuum impregnation technique. Samples were subsequently polished using a well established method for TBC specimens. Gold sputtering was used to allow the ceramic layers to be observed in the SEM. 2.2. Surface analysis After deposition of the SPS layer the top surface of the samples was investigated for surface topographical parameters. Surface
Thermo-cyclic fatigue (TCF) testing is performed primarily to study the ability of a coating to resist high temperature oxidation and the stress of thermally grown oxide on cooling. While failure in TCF testing is driven by bond coat oxidation; the ability of the top coat to survive thermal shock and the stress of oxide growth will also determine the final lifetime of the coating [1]. TCF testing involves placing samples in a furnace at 1100 °C for a period of 1 h. After the heating period, the samples are removed from the furnace hot zone to a cool zone where they are quench cooled with a high flow of compressed air. Cooling below 800 °C occurs within 60 s with the remaining cooling taking longer. The main requirement of the test is that sample temperature is no more than 100 °C within 10 min of the cooling cycle start. After the cooling cycle is completed, the samples are returned to the hot zone to start another cycle. The process is fully automated and the samples are photographed immediately after leaving the hot zone. Failure is considered to have occurred when 20% of the ceramic surface has de-bonded. For each coating type, four samples were prepared for testing. It should be noted that TCF testing is an accelerated test with hot soak temperatures roughly 100 °C higher than expected for a metallic bond coat in service conditions. After samples had failed under TCF testing, the failed samples were mounted in a low viscosity epoxy based resin before sectioning and prepared for microstructural analysis to assess coating microstructure at failure. 2.4. Thermal shock testing Thermal shock testing allows the investigation of the ability of coatings to survive very rapid heating and cooling events. Such testing primarily checks the ability of the coating to cope with the stress of thermal expansion mismatch, sintering and thermal gradients. Unlike TCF testing, the influence of oxidation on the bond coat is of less importance. Thermal shock testing was conducted using a burner rig at GKN Aerospace Engine Systems (Trollhättan, Sweden) [2]. Samples were subjected to 75 second cycles with heating to surface temperatures of 1200 °C and rear face temperatures of between 960 °C and 980 °C. During testing coating temperatures are monitored using pyrometers facing both front and rear sides. Bond coat temperatures of approximately 1000 °C have been correlated to this test condition using samples instrumented with thermocouples. This was not carried out for the coatings within this test series. Samples are preheated before the test starts to 600 °C from the rear face of the sample with hot air guns. After starting the combustion, the rear heaters are switched to compressed air only for cooling of the back face of the sample. Samples are monitored during testing with a video recording system and pyrometer measurements. Failure is deemed to have occurred
Table 1 Experimental coating manufacturing route and layer thickness. Coating ID
Bond coat method
Surface treatment
Surface roughness Ra (μm)
Bond coat thickness (μm)
Plasma 1 Plasma 2 Plasma 3 Plasma 4 HVAF
Plasma Plasma Plasma Plasma HVAF
Polished Polished & grit blasted Grit blasted As-sprayed As-sprayed
1–2 3–4 6–8 11–12 8–9
146 161 186 183 211
μm μm μm μm μm
± ± ± ± ±
7 9 12 11 8
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Top coat thickness (μm) 211 201 205 220 283
μm μm μm μm μm
± ± ± ± ±
6 11 9 17 10
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TBC system. Thermal conductivity may then be calculated using the following relationship [13]:
Table 2 Material properties of TBC layers at room temperature. Material
Density (g/cm3)
Specific heat capacity (J/g·K)
Thermal diffusivity (mm2/s)
Hastelloy-X (Substance) Plasma Bond Coat HVAF bond coat Yttria stabilized zirconia
8.22 6.5 8.2 3.9–4.1
0.468 0.545 0.452 0.538
2.871 1.007 3.277 N/A
when 20% of the ceramic surface has spalled from the test sample. As samples are monitored once per revolution for the test fixture, the result is accurate to within the nearest 4 cycles of failure. Thermal shock samples were prepared from 25 mm diameter buttons 6 mm thick that were coated with the complete TBC system. Before testing, overspray was ground away from the edge of the sample and the button mounted to a carrier plate using a single spot weld. Due to constraints with testing time, only the sample on HVAF bond coat was tested in thermal shock. 2.5. Thermal conductivity analysis Thermal conductivity analysis was performed using the laser flash method that has been long accepted for analysis of coating thermal properties [3]. 10 mm diameter samples for thermal property evaluation were water-jet cut from coated plates. A graphite cover plate with an 8 mm diameter orifice is used to prevent laser light from the sample edge from influencing the signal. This also prevents measurements from the edge region where cutting may have induced microstructural changes. In this case the tested samples consist of the complete TBC system: substrate, bond coat and SPS ceramic layers. For each coating type a minimum of four samples were measured at room temperature only. Thermal diffusivity for the complete coating system was measured for the samples using a Netzsch LFA 247 (Nezsch Gerätebau GmbH, Germany). During a measurement a laser pulse is fired on the back face of the test sample. The resulting temperature increase in the rear face is measured using an infra-red detector. Thermal diffusivity is then calculated from the following formula [13]: 2 α ¼ 0:1388 L =ðtð0:5ÞÞ where α is the thermal diffusivity (mm2/s), L is the thickness of the sample and t (0.5) is the half time taken for the total temperature rise. Six laser shots were made when the sample was at a steady state condition. The result is an average for the thermal diffusivity of the complete
λ ¼ α Cp ρ where λ is thermal conductivity (W·m−1·K−1), Cp is the specific heat capacity (J/g·K), ρ is the density of the material (g/cm3) and α is the thermal diffusivity as expressed previously. As measurements are performed on a complete TBC system, the evaluation is conducted considering the sample as a 3 layer system in which the thermal conductivity of the SPS ceramic is unknown. Using Proteus LFA Analysis software (Nezsch Gerätebau GmbH, Germany), the thermal conductivity of the SPS ceramic layer can be calculated from the measured diffusivity values of the complete TBC system using measured data for the various layers presented in Table 2. Thermal diffusivity and specific heat capacity have previously been measured for the substrate and bond coat materials [15]. The specific heat capacity of the YSZ material was measured using differential scanning calorimeter DSC 404C (Nezsch Gerätebau GmbH, Germany). The influence of thermal expansion of the layers is neglected in this case as there is no significant increase in temperature during the test measurements. Thickness of all the layers within the system was measured from microstructure cross-sections of each individual plate. Thermal conductivity is then evaluated using a three layer model that accounts for thermal contact resistance between the layers and heat loss during the measurement. 3. Results 3.1. Microstructure Figs. 1–4 present the microstructure cross-section and top surface of the samples in order of increasing bond coat surface roughness. In all cases, columnar features are present within the microstructures of the SPS layers and the individual column tops can be identified from the topographic view. It should be stressed again that coatings were deposited with identical processing parameters and sample fixturing during deposition. It can be observed that the greater the surface roughness of the underlying bond coat, the wider and more irregular the columns become in the SPS layer applied above. This can be correlated to the number of columns that can be observed in the top surface images. Higher density of column tops can be observed on bond coats with lower roughness. This observation relates well to the proposed deposition mechanism suggested by VanEvery et al. [11]. It is suggested that the incoming particles (if they are in the range 1 μm in diameter) will follow the flow of the plasma parallel with the surface (bond coat) for some distance before they impact a surface asperity. Thus the individual
Fig. 1. Cross-section and top surface of Plasma 1 coating (polished) SPS coating labeled A, bond coat labeled B and substrate labeled C.
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Fig. 2. Cross-section and top surface of Plasma 2 coating (polished and grit blasted).
columns form by progressive build-up of particulates on surface peaks. Reducing surface roughness will increase the density of peaks in a specific area. As each peak is a possible initiation site for a column within the SPS coating, smoother surfaces tend to generate a greater number of columns. Rough surfaces (as with Plasma 4) have fewer peaks though they are larger in size. This results in a more uneven SPS layer dominated by larger columns. The microstructure of the SPS sample on HVAF bond coat can be observed in Fig. 5. The HVAF bond coats had a surface roughness between that of the APS as-sprayed and grit blasted coating. In this case no additional surface modification has been performed. The presence of some larger particles at the bond coat interface can be observed (arrowed). In some cases these particles appear to be the origin of the columns formed in the coating above. However the presence of such particles presents some potential problems in the coating performance. The influence of such particles has been demonstrated in both experimental and modeling work to be detrimental to the lifetime of APS TBC's due to stress generation [14]. Additionally if the particle is poorly bonded to the coating below, it presents a possible pathway for oxidation and formation of detrimental mixed oxides. A high magnification micrograph of the top surface of a column is shown in Fig. 6. The surface shows a mixture of features present within the structure. The coating is built up from a mixture of small scale splats on the order of 1–2 μm in diameter (labeled A). Present also are partially deformed droplets (labeled B) and spherical particles (labeled C). The differences in particle states may be attributed to the treatment of the particles within the plasma jet. Splats make up the majority of the structure, originating from particles treated in the plasma core. Deformed
and spherical particles are deposited from the plasma jet periphery as it passes over the surface. 3.1.1. Column measurement In order to characterize the influence of surface roughness there is an interest to know the number of columns generated in each case. However the traditional approach of counting the number of columns visually using optical or SEM microstructures is not reliable as it has proven difficult to correctly identify the boundary between adjacent columns (as observed in Figs. 1 and 2). However, the number of columns can be indicated by the surface profile of the sprayed coating. Surface analysis data from striped light projection measurements is displayed in Fig. 7. Sa is the three dimensional arithmetic mean height and Sz is the average peak to valley height. It can be noted that as the surface roughness of the bond coat increases (from left to right on the graph), so does the apparent roughness (Sa and Sz) of the SPS layer. This can be related to the larger but fewer in number columns generated as the surface roughness is increased. This can be confirmed by observations of the column tops in cross-section (Figs. 1–5). More important for describing the column size may be the use of feature parameters such as PSm which describes the mean width of profile features on the surface, the features in this case being the column tops. Fig. 8 displays the results for PSm measured for the experimental coatings. It can be observed that the rough trend of increased feature width (column width) with increasing bond coat roughness is also followed as in the previous results (see Fig. 7). An important difference in this case is that the HVAF coating shows lower feature width relative to that of standard and grit blasted coatings despite a similar level of roughness.
Fig. 3. Cross-section and top surface of Plasma 3 coating (grit blasted).
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Fig. 4. Cross-section and top surface of Plasma 4 coating (as-sprayed).
It should be noted that the use of grit blasting increases the variation in column width. This can be seen in the grit blasted coating that shows dramatically larger deviation than the as-sprayed coating on which it is based. The same deviation can be seen in the polished and grit blasted coating relative to the coating that was only polished. Fig. 9a and b compares the Plasma 1 (polished bond coat) sample surface with the HVAF sample surface respectively. The scale indicates the height difference (in microns) measured on the coating surface. It can be observed that the rougher HVAF coating has a smaller concentration of columns, though they are larger in diameter than the columns on the polished bond coat. These can be compared with the top surface images presented in Figs. 1 and 5 respectively. The surface finish of an as produced coating may be of interest as surface roughness on airfoil components is preferred to be as low as possible as it reduces temperature at surface (problematic) and reduces aerodynamic losses [1]. 3.2. Thermo-cyclic fatigue performance The results for thermo-cyclic fatigue displayed in Fig. 10 are interesting in that there is no observed trend relating column width to lifetime. Statistically speaking there is no significant difference between the tested samples apart from the coating on the polished and grit blasted surface. As stated previously, failure in the TCF test is dominated by bond coat oxidation with ceramic strain tolerance providing only a secondary influence [15]. Micrographs for all of the coatings after TCF failure are presented in Figs. 11–15. It is apparent from the micrographs that there is little
difference to single out in top coat structure to explain the difference in TCF lifetimes. A certain level of cracking is noticeable within all coatings, though this is to be expected due to the buckling failure experienced in the test similar to the failure in the EB-PVD coatings [16]. Plasma 1 samples with a polished bond displays more significant horizontal cracking (arrowed, Fig. 11) within the lower third of the SPS ceramic layer. While it's not apparent that failure resulted due to such cracks; it suggests that the region close to the bond coat may not have sufficient strain tolerance. This could be explained by the dense structure of the SPS layer close to the interface, with the column spacing not identifiable. The same dense structure can be observed in the as sprayed microstructure in Fig. 1. Overall it is apparent that oxidation is the driver of failure. It is observed from the micrographs of the failed samples (Figs. 11–14) that surface modification will in fact change the oxidation behavior of the plasma sprayed bond coat. The polished bond coats (Fig. 11) and the polished and grit blasted coatings (Fig. 12) show similar behavior. In both cases there has been a significant amount of internal oxidation during test exposure, though less in the case of the polished and grit blasted sample. Internal oxidation deep within the coating suggests that polishing the coatings before ceramic deposition may open pathways for oxygen access by exposing delaminations within the coating. Post grit blasting, as was the case in Fig. 10, may seal some of these pathways reducing the oxidation slightly. The grit blasted sample (Fig. 13) displays a different structure. There is none of the internal oxidation as seen in the other coatings. Oxides present in the bond coat layer are those that were produced during the deposition process due to in flight oxidation of the metallic powder.
Fig. 5. Cross-section and top surface of HVAF coating (as-sprayed).
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Fig. 8. Mean width of profile data for the SPS coatings.
Fig. 6. High magnification of SPS coating surface.
The grit blasting step performed in this case may have sealed any open pathways for oxygen attack due to the deformation of the coating surface. VPS/LPPS coatings are often shot peened for a similar reason [17]. Failure appears to be concentrated in the TGO though there are some locations where cracks propagated through the topcoat leaving a small amount still attached to the bond coat (arrowed). The microstructure of the as-produced bond (Fig. 14) coat resembles previously studied coating systems involving the same bond coat method after TCF exposure [15]. The layered oxide (labeled A) at asperities being observed in previous testing and reported in literature [18]. There is some internal oxidation occurring within the top half of the bond coat as oxygen penetrates along splat boundaries. There is also the presence of mixed oxides (labeled B) above the alumina TGO that are not present in the surface treated samples. Analysis of the failures is complicated by the influence of the surface modification on the oxidation process. The samples with modified interfaces tend to form a pure alumina layer whereas the as-sprayed coatings show spinel oxides as well as alumina. The same effect has been reported previously and is believed to be due to removal of seed oxides (generated during deposition) for nucleation of spinel oxides [19–21]. For the plasma sprayed bond coats, failure occurs when the TGO has reached a thickness in the range 6–10 μm. A similar thickness of TGO is
Fig. 7. Surface topography data, Sa and Sz of the experimental SPS coatings.
present in the failed HVAF bond coat samples. There are studies that propose a critical oxide thickness for failure.[22–24]. As the TGO grows in thickness, the stress level at the topcoat bond coat interface increases to a point where cracking is initiated. However the situation is complicated in this case. On smooth surfaces, such as the polished bond coat, the growth stress is minimized and critical stress is
Fig. 9. a. Plasma 1 (polished bond coat) surface topography map. b. HVAF (polished bond coat) surface topography map.
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Fig. 10. Thermo-cyclic fatigue lifetime for experimental coatings.
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Fig. 12. Cross-section of Plasma 2 coating (polished and grit blasted) after TCF testing.
influenced only by the thermal mismatch stress between TGO layer and other layers in the system [22]. As surface roughness increases then there is a greater contribution of growth stress to the overall stress state at the interface. Therefore the failure transitions from a TGO cracking to a mixed TGO/top coat failure as the roughness of the interface is increased. This transition is seen in the TCF test samples presented in this study. Stress required to generate failure may increase when the interface becomes rougher as cracks will increasingly have to be propagated through the top coat. Optimal surface roughness or preparation for SPS coatings is presently not known. Fauchais et al. proposed optimal bonding for an APS ceramic occurs where peak height is 1/2 to 1/3 the diameter of a splat [25]. If SPS coatings behave in a similar way, surface roughness requirements for SPS may be on the scale Ra = 1 μm or lower. Finally the overall TCF results can be understood based on the competing mechanisms. Smooth bond coat/top coat interface generates the lowest stress during high temperature exposure. Rough interfaces generate higher stress but crack propagation may be more difficult. Therefore the polished and grit blasted bond coat shows the lowest lifetime as it shows behavior half way between. The failed SPS coating on HVAF bond coat is displayed in Fig. 15. As with the as-sprayed microstructure (Fig. 5); the presence of a poorly
bonded interface particle can be noted in the micrograph (labeled A). The presence of some cracking within the SPS layer can be noted (labeled B) however, the failure appears to be within the TGO along the majority of the coating cross-section. It can be noted that the remaining beta phase zone within the bond coat (labeled C) is approximately 30–40 μm in width. Beta phase (NiAl) is considered to be the reservoir of aluminium for formation of alumina scale in the TGO [22]. As such there is not a significant amount of chemical lifetime left in the coating. The failure resembles that of the grit blasted plasma spray bond coat with some regions of the YSZ layer still attached to the bond coat. HVAF coatings may benefit from some form of post-treatment due to the number of larger poorly bonded particles present at the coating interface.
Fig. 11. Cross-section of Plasma 1 coating (polished) after TCF testing.
Fig. 13. Cross-section of Plasma 3 coating (grit blasted) after TCF testing.
3.3. Thermal shock testing Thermal shock testing of SPS samples produced in this study has been limited to those produced on an HVAF bond coat. The test was performed on multiple samples and continued up to 9903 cycles. At this point the testing was discontinued without sample failure. As a reference the average lifetime for a conventionally sprayed TBC system in
Please cite this article as: N. Curry, et al., Surf. Coat. Technol. (2014), http://dx.doi.org/10.1016/j.surfcoat.2014.08.067
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Fig. 14. Cross-section of Plasma 4 coating (as-sprayed) after TCF testing. Layered oxide labeled A and mixed oxide labeled B.
Fig. 16. Cross-section of HVAF coating (as-sprayed) after thermal shock testing. Interface particle labeled A and top coat cracking labeled B.
this test is 1800 cycles. The SPS samples exceed the best lifetime by a factor of 5 without failure. The microstructure of the SPS coating with HVAF bond coat after 9903 thermal shock cycles can be seen in Fig. 16. The bond coat exhibits a well defined two phase structure with lighter shaded gamma phase and darker shaded beta phase. There is very little depletion of the beta-phase at either the bond coat interface region or at the substrate interface. This can be expected due to the total exposure time at high temperature being around 197 h and the bond coat temperature being approximately 100 °C lower compared to that in the TCF test. Two poorly bonded particles can be noted at the bond coat interface (labeled A). As discussed previously, such feature may be detrimental to long term coating lifetime. Some cracks within the SPS coating can be observed (arrowed B) though not significant enough yet to cause delamination of the ceramic. A higher magnification image of the bond coat/ceramic interface is displayed in Fig. 17 with key features labeled. The alumina layer generated at the interface during high temperature exposure can be observed together with clear gamma and beta phases within the bond coat. There are also some areas of mixed oxide present above the alumina layer. The
structure of the SPS layer close to the interface shows no obvious signs of sintering as the pore shape is still irregular. Overall the coating appears to have a significant amount of life left to it in the thermal shock test and the HVAF coating seems promising as a bond coat for the SPS TBC system. Further improvement may be achieved if the oxide formed during testing consisted of alumina only in order to reduce interface stress.
Thermal conductivity results are displayed in Fig. 18. Overall the results demonstrate that as sprayed thermal conductivities for SPS coatings are somewhat lower than the average for a conventional APS ceramic where 1 W/m∙K is more common [26]. Overall results would suggest an increase in thermal conductivity for increasing surface roughness. This can be explained by the reduction in amount of porosity due to fewer column boundaries. The coating on the standard APS bond coat does not fit the trend. However, due to the highly irregular bond coat surface, column boundaries tend to be oriented at an angle rather
Fig. 15. Cross-section of HVAF coating (as-sprayed) after TCF testing. Interface particle labeled A, topcoat cracking labeled B and beta-phase zone labeled C.
Fig. 17. High magnification image of the HVAF coating topcoat/bond coat interface after thermal shock testing.
3.4. Laser flash analysis
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Acknowledgments Thanks to Stefan Björklund for assistance with spraying of the HVAF bond coat samples. Thanks go to Wyszomir Janikowski for his work on sample evaluation. Thanks to Jönköping University for access to the LFA equipment. Thanks also to Lars Östergren and Nicholas Erb at GKN for help with the thermal shock rig testing. References
Fig. 18. Thermal conductivity data for experimental coatings.
than perpendicular to the substrate surface. This may increase the insulating effect of the boundary regions.
4. Conclusions The results of this study demonstrate that SPS coating microstructure is influenced by the topography of the surface on which it is deposited. While the microstructure influence on performance properties is masked in this case by bond coat effects; in principle SPS microstructure could be partially designed by controlling the surface topography on which it is deposited. Surface modification of APS bond coats is not recommended due to dramatic changes in oxidation behavior. However, HVAF or HVOF coatings may be suitable for further development of surface modification. The fact that SPS coatings adhere to reasonably smooth surfaces means that diffusion aluminide bond coats could be attempted as used with EB-PVD coatings. The deposition requirements for SPS coatings are not yet well defined as they are for APS or EB-PVD coatings. More work is required in order to fully understand the optimal surface requirements for an SPS TBC system to achieve the best combination of properties. In reality it is likely that the surface requirements for good bonding and lifetime are somewhat more complicated than simple factors such as Ra.
[1] F.O. Soechting, J. Therm. Spray Technol. 8 (4) (Dec. 1999) 505–511. [2] R. Miller, J. Therm. Spray Technol. 6 (1) (Mar. 1997) 35–42. [3] U. Schulz, C. Leyens, K. Fritscher, M. Peters, B. Saruhan-Brings, O. Lavigne, J.-M. Dorvaux, M. Poulain, R. Mévrel, M. Caliez, Aerosp. Sci. Technol. 7 (1) (Jan. 2003) 73–80. [4] M. Karger, R. Vaßen, D. Stöver, Surf. Coat. Technol. 206 (1) (Oct. 2011) 16–23. [5] A. Feuerstein, J. Knapp, T. Taylor, A. Ashary, A. Bolcavage, N. Hitchman, J. Therm. Spray Technol. 17 (2) (Jun. 2008) 199–213. [6] L. Pawlowski, Surf. Coat. Technol. 203 (19) (Jun. 2009) 2807–2829. [7] A. Guignard, G. Mauer, R. Vaßen, D. Stöver, J. Therm. Spray Technol. 21 (3–4) (Jun. 2012) 416–424. [8] Z. Tang, H. Kim, I. Yaroslavski, G. Masindo, Z. Celler, D. Ellsworth, Proceedings of the International Thermal Spray Conference, Hamburg, 2011. [9] J. Fazilleau, C. Delbos, V. Rat, J.F. Coudert, P. Fauchais, B. Pateyron, Plasma Chem. Plasma Process. 26 (4) (Aug. 2006) 371–391. [10] J.O. Berghaus, S. Bouaricha, J.-G. Legoux, C. Moreau, T. Chráska, Proceedings of the International Thermal Spray Conference, Basel, Switzerland, 2005. [11] K. VanEvery, M. Krane, R. Trice, H. Wang, W. Porter, M. Besser, D. Sordelet, J. Ilavsky, J. Almer, J. Therm. Spray Technol. 20 (4) (2011) 817–828. [12] N. Curry, K. VanEvery, T. Snyder, N. Markocsan, Coatings 4 (Aug. 2014) 630–650. [13] R.E. Taylor, Mater. Sci. Eng. A 245 (2) (May 1998) 160–167. [14] M. Gupta, K. Skogsberg, P. Nylén, J. Therm. Spray Technol. 23 (1–2) (Jan. 2014) 170–181. [15] N. Curry, N. Markocsan, L. Östergren, X.-H. Li, M. Dorfman, J. Therm. Spray Technol. 22 (6) (Aug. 2013) 864–872. [16] A.G. Evans, D.R. Mumm, J.W. Hutchinson, G.H. Meier, F.S. Pettit, Prog. Mater. Sci. 46 (5) (2001) 505–553. [17] J.F. Loersch, J.W. Neal, Peened Overlay Coatings, Apr 30, 1985. 4514469. [18] O. Trunova, T. Beck, R. Herzog, R.W. Steinbrech, L. Singheiser, Surf. Coat. Technol. 202 (20) (Jul. 2008) 5027–5032. [19] N. Czech, M. Juez-Lorenzo, V. Kolarik, W. Stamm, Surf. Coat. Technol. 108–109 (Oct. 1998) 36–42. [20] A. Gil, V. Shemet, R. Vassen, M. Subanovic, J. Toscano, D. Naumenko, L. Singheiser, W.J. Quadakkers, Surf. Coat. Technol. 201 (7) (Dec. 2006) 3824–3828. [21] D. Mercier, C. Kaplin, G. Goodall, G. Kim, M. Brochu, Surf. Coat. Technol. 205 (7) (Dec. 2010) 2546–2553. [22] H.E. Evans, Surf. Coat. Technol. 206 (7) (Dec. 2011) 1512–1521. [23] M. Tamura, M. Takahashi, J. Ishii, K. Suzuki, M. Sato, K. Shimomura, J. Therm. Spray Technol. 8 (1) (Mar. 1999) 68–72. [24] H. Aleksanoglu, A. Scholz, M. Oechsner, C. Berger, M. Rudolphi, M. Schütze, W. Stamm, Int. J. Fatigue 53 (Aug. 2013) 40–48. [25] P. Fauchais, R. Etchart-Salas, V. Rat, J.F. Coudert, N. Caron, K. Wittmann-Ténèze, J. Therm. Spray Technol. 17 (1) (Mar. 2008) 31–59. [26] N. Curry, J. Donoghue, Surf. Coat. Technol. 209 (Sep. 2012) 38–43.
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