The determination of the delamination resistance in thermal barrier coating system by four-point bending tests

Surface & Coatings Technology 201 (2006) 744 – 754 www.elsevier.com/locate/surfcoat

The determination of the delamination resistance in thermal barrier coating system by four-point bending tests Y. Yamazaki a,⁎, A. Schmidt b , A. Scholz b a

Department of Mech. and Cont. Eng., Niigata Institute of Technology, 1719 Fujihashi, Kashiwayaki, Niigata 945-1195, Japan b Institute of Materials Technology, Darmstadt University of Technology, Grafenstr. 2, D-64283 Darmstadt, Germany Received 23 May 2005; accepted in revised form 16 December 2005 Available online 7 February 2006

Abstract The delamination resistance in a plasma-sprayed thermal barrier coating (TBC) system was evaluated by means of a modified four-point bending test. In this work, in order to study the effect of the interface roughness between the bond coating (BC) and the top coating (TC), two kinds of the powder with the different particle size were used for spraying the BC material. In addition, the influence of the isothermal aging on the delamination resistance was also investigated. The experimental results indicate that the effect of the TC/BC interface roughness wasn't significant. The energy release rate Gc, which was estimated by the modified four-point bending tests, increased with increasing aging time at 1000 °C, on the other hand, the scatter of it decreased with increasing aging time. The crack mainly propagated in the top coating for as-sprayed condition and at the top-coating/bond-coating interface after thermal aging for 2000 h. The energy release rate Gc was correlated with the fracture strength of the weakest parts of the TBC specimen. The energy release rate Gc increased with thermal aging until a critical time due to the sintering of the top-coating expired. It can be expected that Gc will decrease with the thermal aging after the critical aging time because the top-coating/ bond-coating interfacial strength decreases by the TGO growth. The critical aging time in this work is approximately 2000 h. © 2005 Elsevier B.V. All rights reserved. Keywords: Thermal barrier coating; 4-point bending test; Interface; Delamination; Energy release rate; Thermal aging

1. Introduction Industrial gas turbines are used as electrical generators by both utilities and private industrial companies [1–11]. The durability of a gas turbine is principally limited by those components operating at high temperatures in the turbine sections because those are exposed to hot gas. Thermal barrier coatings (TBC), that reduce the temperature in the underlying substrate material, are an essential requirement for the hot section components. A typical TBC system is composed of an oxidation resistant metallic bond coating (BC) on the superalloy substrate and a thermal insulating ceramic top coating (TC) attached to the BC. To take full advantage of the potential of the TBC systems, the evaluation and the prediction of the lifetime of TBC systems are important. Therefore experimental and analytical investigations in TBC systems have been performed [4–11]. As a result of these ⁎ Corresponding author. Tel.: +81 257 22 8109; fax: +81 257 22 8119. E-mail address: [email protected] (Y. Yamazaki). 0257-8972/$ - see front matter © 2005 Elsevier B.V. All rights reserved. doi:10.1016/j.surfcoat.2005.12.023

energetic studies, it has clearly been shown that the adherence of the TC is the most important parameter for the durability of TBC system [6–8]. It was also revealed by recent investigations that thermal grown oxide (TGO) grows at the TC/BC interface due to the thermal aging as well as thermal cycling. In addition, it was believed that the TGO plays an important role in the spallation of TBC systems [9–11]. Charalambides et al. proposed and discussed the evaluation of the critical energy release rate at the metal/ceramic interface by means of four-point bending tests [12]. This method has the advantage that the specimen geometry and test technique are simple. The critical energy release rate due to the delamination at the metal-ceramic interface can be determined easily by using this method. However, this method is only applicable for thickcoated materials that have relatively high fracture toughness to prevent vertical cracking. If it is applied to thin brittle coated materials such as TBC systems, the vertical cracking and the segmentation decrease the stored elastic energy and make the evaluation of the interface fracture energy significantly more difficult. Hofinger et al. proposed a simple modification of the

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Table 2 Mechanical properties of each layer [14]

1 2 3 4 5

Layer

Material

Young's modulus [GPa]

Poisson's ratio

Substrate Bond coat TGO Top coat Stiffener

CMSX-4 LCO22

135 183 310 50.0 205

0.3 0.3 0.3 0.3 0.3

8YSZ Steel

The final goal in this project is the development of a reliable life time model for the TBC systems. The object of the present work is to evaluate and investigate the delamination resistance of the TBC systems by means of the modified Charalambides four-point bending test. The effects of the thermal aging and the TC/BC interface roughness on the delamination resistance are also discussed. 2. Material

Fig. 1. (a) Cross sectional microstructure near the top-coating/bond-coating interface after the thermal aging at 1000 °C for 1000 h and (b) TGO thickness as a function of thermal aging time [14].

Charalambides test in order to evaluate the interfacial fracture toughness of thin brittle layers tending to separate by vertical cracks [13]. The modification by Hofinger is the bonding of a stiffener on the top of the thin brittle surface layer. This stiffening layer suppresses the segmentation of the brittle layer and increases the stored energy in the layer and therefore the driving force for the delamination. Another advantage of this method is that an analytical solution is possible. This modified Charalambides test is suitable to evaluate the delamination resistance in TBC systems.

The 8 wt.% yttria partially stabilized zirconia, 8YSZ, was used as top coating (TC) layer and was deposited on the bond coating by air plasma spraying. The single crystal Ni-base superalloy CMSX-4 prismatic bar of 5 mm in thickness and 10 mm in width was used as the substrate. The CoNiCrAlY alloy, LCO22, was selected as the bond coating (BC) layer between TC and substrate. In this work, in order to study the effect of the interface roughness between BC and TC, two kinds of the LCO22 powder with different particle size were used. The specimens that were sprayed with fine grain LCO22 powder are called fine BC specimens while those that were sprayed with coarse powder are called coarse BC specimens. In this work, the effect of the TGO thickness on the delamination resistance of the TBC system was also studied. Typical microstructures near the TC/BC interface after thermal exposure at 1000 °C are shown in Fig. 1(a) [14]. The TGO thickness was measured and compared with the aging time. The TGO thickness as a function of the aging time is shown in Fig. 1 (b) [14]. In Fig. 1(b), the TGO thickness increased with increasing aging time. The parameters of the specimens are summarized in Table 1. In this table, the TGO thicknesses estimated from the relationship as shown in Fig. 1(b) are also listed. The mechanical properties of substrate, bond coating, top coating, stiffener and TGO are summarized in Table 2.

Table 1 Parameters of the specimens used Specimen no.

Bond coating powder size (LCO22)

Average surface roughness of bond coating, Ra [μm]

Thickness of bond coating [μm]

Thickness of APS-YSZ [μm]

Isothermal oxidation at 1000 °C

Estimated TGO thickness [μm]

C01-C04 C05, C06 C07, C08 C09, C10 C18-C20 C12, C13 C14, C15 C16, C17

Fine Fine Fine Fine Coarse Coarse Coarse Coarse

8.71 8.71 8.71 8.71 9.52 9.52 9.52 9.52

360 360 360 360 350 350 350 350

340 340 340 340 330 330 330 330

0h 200 h 1000 h 2000 h 0h 200 h 1000 h 2000 h

0 3 5 7 0 3 5 7

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3. Specimen preparation The prismatic notched bar specimens that were 130 mm in length, 10 mm in width and approximately 9.9 mm in thickness (including stiffener), as shown in Fig. 2 were prepared. The thicknesses of the top coating and the bond coating were listed in Table 1. The mechanical notch with a width of 2 – 3 mm was induced in the top coating of the specimen's centre after the thermal exposures. In this work, in order to prevent vertical crack in top coating, namely to inhibit the segmentation of top coating, the stiffener with the sized of 60 × 10 × 4.2 mm were bonded on top coating, as shown in Fig. 2. The material of the stiffener was structural steel. The heat curing type epoxy adhesive, Araldite AT1 (CIBAGEIGY Ltd.), was selected and used for bonding of the stiffener. However, it can be considered that the adhesive infiltrates into the TC during the curing process because the TC has porous microstructure. If the adhesive infiltrates until the TC/BC interface, it has undesirable influence on experimental results. Therefore the infiltration depth of the adhesive must be small or almost zero. The microstructures of the TC after bonding are shown in Fig. 3. It can be observed in Fig. 3 that the pores near the TC/AT1 interface were filled with the dark material. The results of the EDX analysis are shown in Fig. 4. Peaks of carbon and oxygen were found in the filled pores. Because AT1 consists of carbon and oxygen, these pores were filled with the adhesive. In particular, the carbon and oxygen peaks can be found in the portion of approximately 100 μm from TC/AT1 interface, but there is no evidence of oxygen in deeper than 100 μm. These results indicate that the adhesive infiltrated approximately 100 μm in depth into TC (less than one third of TC thickness). Therefore, the influence of the adhesive infiltration on the test results can be neglected. The detail of bonding procedure was reported in Ref. [15].

Fig. 3. SEM image of the cross sections of the bonded TBC/Stiffener specimens.

can apply only a constant displacement which is 7 mm/min, a spring unit was used in order to convert the constant displacement speed into the constant loading speed. Most of four-point bending tests were carried out under constant loading speed at 5 N/s. Some tests were performed at 0.6 N/s to study the effect of the loading speed on the experimental results. All tests were conducted at room temperature. The detail, the accuracy and the reproducibility of the test equipment were reported in Ref. [15]. 5. Analytical considerations The specimen was loaded in the four-point bending as shown in Fig. 2. In this case, the delamination occurs under the combined influence of residual stresses and those generated by the application of an external mechanical load. It can be shown in Ref. [16] that the energy release rate, G, is given by, G ¼ Gp þ Gpr þ Gr

ð1Þ

4. Equipment and test conditions The schematic illustration of test equipment is shown in Fig. 5. This test equipment consists of a load cell with a load capacity of 2 kN, the spring unit with the four point bending device and an extensometer. In this work, a creep machine was used to apply the loading to the specimen. Since the machine

The first term in Eq. (1), Gp, is the standard linear elastic fracture mechanics expression and the residual stress-independent contribution to the crack driving force. The third term in Eq. (1), Gr, is the release rate for the relaxation of the residual stresses. This is the driving force available for crack extension with no applied load. The second term in Eq. (1), Gpr, represents

Fig. 2. Specimen geometry in this work, specimen width is 10 mm (unit in mm).

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Fig. 4. EDX analysis results near TBC/Stiffener interface (a) SEM image, (b) Zr map, (c) C map, (d) O map.

the interaction between the aplied load and the residual stresses distribution. In the four-point bending delamination test shown in Fig. 2, as long as the crack is located between the inner loading spans, the energy release rate for the specimen without

residual stress, GP, can be calculated as a modification of Hofinger et al. [13],
 M2 1 1 − Gp ¼ 2 ð2Þ 2b Ec;3 T Ic;3 T Ec;5 T Ic;5 T with the constant bending moment M ¼ PL=2

ð3Þ

and the second moments of area per unit width T¼ Ic;k

k n t þ t o X 1 c;i c;i−1 ¼ −Y0;k Þ2 EiT Ii þ ti Ec;k T 2 i¼1

ð4Þ

with k X

Yo;k ¼

k X

2 2 EiTðtc;i −tc;i−1 Þ

i¼1

2

k X

;

Ec;k T¼

EiTti

i¼1

Ii ¼

Fig. 5. Schematic illustration of a test equipment (front view).

ti3 ; 12

tc;i ¼

ti EiT

i¼1 k X

; ti

i¼1 i X i¼1

ti ;

EiT ¼

Ei ð1−mi Þ

ð5Þ

where t is the thickness of the layer, b denotes the width of the specimen, E and ν represent the Young's modulus and Poisson's ratio, respectively, and the subscripts 1, 2,..,5 refer

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with R R R R DTf Ei y2 dy Ei ai dy− Ei ydy Ei ai ydyg ; A¼ R R R Ei dy Ei y2 dy−ð Ei ydyÞ2 B¼

Fig. 6. Effects of the Young's modulus of TC and the stiffener thickness on the energy release rate; E4,as-spray = 50 GPa (see Table 2).

to the substrate, the BC layer, the TGO layer, the TC layer, and the stiffener. Using Eq. (2) with the mechanical properties of the materials as shown in Table 2, the effects of Young's modulus of the TC and the thickness of the stiffener on the energy release rate for the specimen without residual stress, Gp, were studied analytically. The normalized energy release rate as a function of the Young's modulus of the TC is shown in Fig. 6 for different stiffener thickness. From Fig. 6, the energy release rate increased essentially under constant substrate loading conditions through the stiffening. The calculations also show that the influence of the Young's modulus of the TC on the energy release rate for the specimen without residual stress, Gp, decreases with increasing thickness of the stiffener. If the stiffener thickness is 4.2 mm (actual stiffener thickness in this work), the effect of Young's modulus of the TC on the energy release rate for the specimen without residual stress, Gp, is negligible even if the Young's modulus of the TC is changed during the thermal aging. The third term in Eq. (1), Gr, can be calculated by [16], Gr ¼

  1 I4 2 Ic;3 2 t4 2 tc;3 2 jtc þ jsb þ rtc þ rsb 2 E4 T Ic;3T E4T Ec;3T

R R R R DT f Ei dy Ei ai ydy− Ei ydy Ei ai dyg R R R Ei dy Ei y2 dy−ð Ei ydyÞ2

ð8Þ

where, αi is the thermal expansion coefficient and is listed in Ref. [15] and y is the distance from the rear surface of the substrate. ΔT is the temperature difference from the neutral temperature (in which the residual stresses become zero) to the room temperature (the experimental condition). However, it is much difficult to measure the ΔT directly in the APSTBC systems because the temperature distribution in the TBC system is unstable during the APS process. In addition, it was reported [19,20] that the residual stress distribution in the APS-TBC system was easily changed due to the thermal aging; the average residual stress was almost 50 MPa (tension) in the TC of the as-spray specimens, however almost − 100 MPa (compression) in the TC after thermal ageing at 900 °C for 100 h. And the average residual stress in the TC was almost independent to the thermal aging time [20]. From these evidence, in this work, the ΔT in Eqs. (7) and (8) were determined using the average residual stress in the TC; 50 MPa for as-spray condition on the other hand, − 100 MPa for the thermal aged condition. The residual stress distributions in the TBC specimens estimated from Eq. (7) are shown in Fig. 7. The residual stress distribution in the TBC specimens was changed due to the thermal aging. The second term in Eq. (1), Gpr, can be calculated by [16], Gp ¼

P dur P Dur ¼ 2 da 2 Da

ð9Þ

where, ur is the residual displasment, ur = u(P = 0) and can be four from a load-displacement curve [16,21]. Fig. 8 shows how the three – Gp, Gpr, Gr – vary with the applied load during delamination testing of as-sprayed and aged specimens. The difference of the residual stress distribution

ð6Þ

where, κ is the elastic stress gradient, relaxed by allowing the layer to bend freely, σ is the average axial elastic stress in the layer to extend or contract freely. The subscript tc and sb refer to the TC and the composite of substrate and BC, respectively. Since the residual stress distribution in the specimen can't be measured easily, in this work, the residual stress distribution in the specimen was estimated from [17,18], rres;i ¼ −

Ei ai Ei DT þ ð A þ ByÞ 1−mi 1−mi

ð7Þ Fig. 7. Estimated distribution of residual stress within TBC specimens.

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Fig. 9. Typical load – deflection curve in the four-point bending test.

Fig. 8. Variations of G, Gp, Gr, and Gpr with applied load during 4-point bending test, (a) as-sprayed, fine BC specimen, (b) aged at 1000 °C for 2000 h, fine BC specimen. From the critical load at which interfacial crack propagation occurred, Pc, the critical energy release rate, Gc, can be obtained.

slope of the load – deflection curve decreased with increasing deflection. At stage IV the same behavior can be observed. In these stages the vertical cracks were initiated at the edge of the stiffener and propagated through the TC to the TC/BC interface. The plateaus of the curve can be observed at stage III and at stage V. At these stages, the delamination crack propagated rapidly until the crack tip approached the inner loading lines. At stage VI the deflection increased again in proportion to load. The energy release rate for the delamination was calculated from the critical load value corresponding to the plateaus. A typical deflection–time curve is shown in Fig. 10. The deflection rate of the specimen (in other words, the slope of the deflection–time curve) is also shown in this figure. The peak deflection rate may correspond to the delamination crack growth rate. The influence of the thermal aging on the deflection rate will be discussed later.

(Fig. 7) are reflected in the higher value of Gr for the aged TBC specimen (Fig. 8b). As shown in Fig. 8, from the critical applied load, Pc, which interfacial crack propagation occurred, the critical energy release rate (the interfacial fracture energy), Gc, can be obtained. 6. Experimental results and discussions A typical load–deflection curve during the four-point bending test is shown in Fig. 9, with the analytical curve calculated by the beam theory without the delamination. In this figure, the load–deflection curve consists of 6 stages. At stage I, the load – deflection curve of the specimen was linear and the experimental curve was equivalent to the analytical one. After the linear stage in the load–deflection curve, at stage II, the

Fig. 10. Typical deflection – time and deflection rate – time curves.

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Y. Yamazaki et al. / Surface & Coatings Technology 201 (2006) 744–754 Table 3 The interfacial fracture energy of various interfaces within multilayer coating systems [13,21–23] System

Interfacial Interface fracture energy [N/m]

Note

ZrO2/NiCrAlY/ Ni-alloy ZrO2/Al2O3/ NiCrAlY/ Ni-alloy NiCrAlY/ Ni-alloy Al2O3/NiCrAlY/ Ni-alloy Al2O3/Ni-alloy

≈40

ZrO2/NiCrAlY

≈35

ZrO2/Al2O3

4-point [21,22] bending test 4-point [21,22] bending test

≈150

NiCrAlY/ 4-point [21,22] Ni-alloy bending test Al2O3/NiCrAlY 4-point [21,22] bending test Al2O3/Ni-alloy 4-point [21,22] bending test Cr2O3/Ni-alloy Peel test [23] ZrO2/Ni-alloy 4-point [13] bending test

Cr2O3/Ni-alloy ZrO2/Ni-alloy

≈110 ≈90 ≈190-250 ≈60

Reference

Fig. 11. Effect of the loading speed on the Gp.

In order to study the effect of the loading speed on the Gp the bending tests were also performed under slow loading speed (approximately 0.6 N/s, this value was almost ten times lower than a normal test speed). The average value of Gp as a function of the loading speed (also the deflection speed) is shown in Fig. 11 with the scatter of the Gp indicated by an error bar. In this figure, the average value of Gp under the slow loading condition was almost equivalent to those under the normal condition; the data of the former were within the scatter band of the latter. Therefore, there is no significant influence of the loading speed on the delamination resistance in this work. The average critical energy release rate, Gc, as a function of the thermal aging time at 1000 °C is shown in Fig. 12. Where Gc was evaluated from Pc with Eq. (1). In this figure, the average critical energy release rate, Gc, was indicated by each dot together with the scatter of the energy release rate. The number

Fig. 12. Average energy release rate, Gc, as a function of the aging time at 1000 °C.

close to the scatter bar indicates the number of specimens tested for each condition, respectively. Comparing critical energy release rates between fine and coarse BC specimens in Fig. 12, the influence of the TC/BC interface roughness on the critical energy release rate wasn't significant. Table 3 shows the interfacial fracture energy of various interfaces within multilayer coating systems [13,21–23]. The critical energy release rate in present TBC specimen is slightly higher than the interfacial fracture toughness of other TBC system, however almost comparable to the other ceramic/metal interface strength. From Fig. 12 the average critical energy release rate increased with increasing aging time at 1000 °C. On the other hand, the scatter of the Gc decreased with increasing aging time. It was reported by the “Subcommittee on Superalloys and Coatings” in the Society of Materials Science, Japan (JSMS) that the adhesion strength of an APS-TBC, which was evaluated by means of tensile tests, increased with increasing thermal aging time at 900 °C [24]. Although test method and aging temperature are different, the trends by tensile tests were quite similar to our results. The fracture toughness KIc of the free-

Fig. 13. Fracture toughness of free-standing TBC film as a function of the aging time at 1316 °C [17].

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standing APS-TBC film and also the GIc of it, which was 2 calculated from KIc /E, are shown in Fig. 13 [25]. Comparing Fig. 13 with Fig. 12, the influence of the thermal aging of the GIc of the free-standing APS-TBC film was similar to the results in this work. The GIc of the free-standing APS-TBC film was almost equivalent to the average critical energy release rate in this work not only for the as-sprayed specimen but also for the aged specimens. Typical microphotographs of the crack propagation path after the test are shown in Fig. 14. It is revealed from Fig. 14(a)

Fig. 15. Typical fracture surfaces after the tests (a) as-sprayed, coarse BC specimen and (b) aged, fine BC specimen.

that the crack mainly propagated in the TC for the as-sprayed specimen. On the other hand, after the thermal aging for 2000 h, the crack mainly propagated at the TC/BC interface as shown in Fig. 14(b) and (c). It can be also observed from these figures

Fig. 14. Cross sectional microphotographs of the crack propagation path; (a) assprayed, fine BC specimen, (b) aged at 1000 °C for 2000 h, fine BC specimen and (c) aged at 1000 °C for 2000 h, coarse BC specimen.

Fig. 16. Tensile strength of free-standing APS-TBC film as a function of the area ration of inter-splats fracture surface; H1–H3 used the hollow powder, FC1– FC3 used the fused crashed powder [18].

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inter-splats fracture of the free-standing APS-TBC film was reported in Ref. [26]. Fig. 16 represents the tensile strength of the free-standing APS-TBC film as a function of the area ratio of the inter-splats fracture surface [26]. From Fig. 16 the tensile strength increased with decreasing inter-splats fracture mode. These results indicate that the inter-splats strength of the TBC was increased by the sintering. It can also be considered in this work that the TC sintered during the thermal aging. The crack path in the specimen by the four-point bending test is summarized in Fig. 17 schematically. For the as-sprayed condition, because the TC has a lower fracture strength compared with the TC/BC interfacial strength, the crack mainly propagated in the TC. After aging at 1000 °C for 2000 h, the fracture strength of the TC became high due to the sintering. On the other hand, the TC/BC interfacial strength decreased with the aging time due to the TGO nucleation and growth. Therefore the crack mainly propagated at the TC/BC interface. The crack mainly propagated in the TC after the thermal aging for 200 h and 1000 h because the TC/BC interface maintained high strength in comparison to the TC. From the above results, the effect of the thermal aging on the energy release rate is summarized in Fig. 18 schematically. In as-sprayed TBC specimen, the TC/BC interface has the higher strength compared with the TC. The fracture strength of the TC increases with the thermal aging by the sintering, on the other hand, the TC/BC interfacial strength decreases due to TGO growth. When the thermal aging time becomes longer than the critical time, the TC/BC interfacial strength becomes lower than the fracture strength of the TC. The energy release rate Gc evaluated by the four-point bending test is correlated with the fracture strength of the weakest parts of the TBC specimen. Therefore it increases with the thermal aging time until the critical time. In addition, it is easily expected that the Gc decreases with the thermal aging after the critical aging time. The critical aging time in this TBCs is approximately 2000 h. The thermal aging up to 10 000 h is now being applied to TBC specimens, which will be presented elsewhere.

Fig. 17. Schematic illustration of fracture mode in TBC specimen by the 4-point bending tests; aging temperature = 1000 °C, (a) as-sprayed, (b) aged for 200 h and 1000 h and (c) aged for 2000 h.

that the TGO grew at the TC/BC interface after the thermal aging. In the 200 h and 1000 h aged specimens, although the TGO also grew at the TC/BC interface, the crack however mainly propagated in the TC. Typical fracture surfaces for the as-sprayed and the aged specimens are represented in Fig. 15. The fracture surface of the as-sprayed specimen exhibits mainly inter-splats fracture mode, on the other hand, the aged specimen takes the mixed mode of the TC/BC interface and inter-splats fractures. The relationship between the fracture strength and the

Fig. 18. Schematic illustration of the energy release rate of the TBC specimen evaluated by the 4-point bending tests.

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(3)

(4) (5)

(6)

Fig. 19. Deflection rate correlated with aging time at 1000 °C.

Since the observation during the four-point bending tests, it was seen that the delamination crack growth rate for the as-sprayed specimens was much lower than that for the aged specimens. Fig. 19 shows the relationship between the peak deflection rate (i.e. the crack growth rate) and the thermal aging time. In this figure, the scatter of the peak deflection rate was also shown by an error bar. From Fig. 19 the peak deflection rate, i.e. the crack growth rate, increased with increasing thermal aging time. The mechanism of this phenomenon was not clear yet. However, following possibilities can be considered. a) The delamination of the aged specimen occurred at higher critical load compared to the as-sprayed specimen. Therefore the higher applied load may induce the higher crack propagation rate. b) If the connection of the splats for the as-sprayed TBC was looser than that for the aged TBC, the crack propagation may be disturbed by the shielding effects produced by the micro-crack nucleation near the crack-tip. However, there is no evidence for these considerations, therefore further investigations, such as metallographical ones, must be performed. 7. Conclusions In this study, the evaluation of the delamination resistance in the TBC systems was carried out by means of modified Charalambides four-point bending tests. The effects of the thermal aging and the top-coating/bond-coating interface roughness on the delamination resistance were also investigated. The following conclusions were made: (1) From SEM and EDX analyses, the adhesive infiltrated approximately 100 μm in depth into the top coating, however there is no significant influence of the adhesive infiltration on the test results. (2) From analytical considerations, the effect of the Young's modulus of the top coating on the energy release rate is

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negligible even if it was changed during the thermal aging. The energy release rate Gc, which was estimated by the four-point bending tests, was independent of the loading rate and the top-coating/bond-coating interface roughness. Gc increased with increasing aging time at 1000 °C, on the other hand, the scatter of it decreased. The crack mainly propagated in the top coating for the assprayed condition, however at the top-coating/bondcoating interface after the thermal aging for 2000 h. Gc was correlated with the fracture strength of weakest parts of the TBC specimen. The Gc increases with the thermal aging until the critical time due to the sintering of the top-coating. It can be expected that Gc decreases with the thermal aging after the critical aging time because the top-coating/bond-coating interfacial strength decreases by the TGO growth. The critical aging time in the present work is approximately 2000 h.

Acknowledgements Authors thank Siemens PG, ALSTOM power, MTU Aero Engines GmbH, MAN Turbomaschine AG, Rolls-Royce Deutschland Ltd. and CO KG and COTEC GmbH for providing the specimens and the coating. References [1] H.L. Bernstein, T.S. Grant, R.C. McClung, J.M. Allen, ASTM STP 1186 (1993) 212. [2] R.A. Miler, J. Therm. Spray Technol. 6 (1997) 35. [3] B.B. Seth, Superalloy 2000, TMS, Warrendale, PA, 2000, p. 3. [4] K. Schneider, H.W. Grunling, Thin Solid Films 107 (1983) 395. [5] P.K. Wright, Mater. Sci. Eng., A Struct. Mater.: Prop. Microstruct. Process. 245 (1998) 191. [6] H. Echosler, D. Renusch, M. Schüüze, Proceedings of Turbomat Symposium, 2002, p. 17. [7] D. Renusch, H. Echsler, M. Schütze, Conference Proceedings in Life Time Modeling of High Temperature Corrosion Process, 2001, p. 324. [8] H. Echosler, D. Renusch, M. Schütze, Mater. Sci. Eng. 20 (2004) 307. [9] T. Xu, S. Faulhaber, C. Mercer, M. Maloney, A. Evans, Acta Mater. 52 (2004) 1439. [10] M.S. Ali, S. Song, P. Xiao, J. Mater. Sci. 37 (2002) 2097. [11] A.N. Khan, J. Lu, H. Liao, Mater. Eng. A 359 (2003) 129. [12] P.G. Charallambides, J. Lund, A.G. Evans, R.M. McMeeking, J. Appl. Mech. 56 (1989) 77. [13] I. Hofinger, M. Oechsner, H. Barhr, M.V. Swain, Int. J. Fract. 92 (1998) 213. [14] H. Echsler, W. Przybilla, M. Schütze, Proceedings of EUROCORR 2000, 2000, CD-ROM. [15] Y. Yamazaki, “Four Point Bending Test in Thermal Barrier Coating System”, IfW-Darmstadt, 2005. [16] S.J. Howard, Y.C. Tsui, T.W. Clyne, Acta Metall. Mater. 42 (1994) 2823. [17] M. Toyoda, et al., IIW Doc. X-1161-88, 1988. [18] M. Toyoda, et al., Q. J. Jpn. Weld. Soc. 7 (1989) 543 (in Japanese). [19] Subcommittee on Superalloys and Coatings JSMS, “Reports of Phase II Activities — Basic Physical Properties of Air Plasma Sprayed Thermal Barrier Coating —“, The Society of Materials Science, Japan, 2005 (in Japanese). [20] M. Arai, et al., Proc. of the 43rd Symposium on Strength of Materials at High Temp., 2005, p. 166 (in Japanese).

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