A method for evaluating the creep properties of overlay coatings

Surface and Coatings Technology 124 (2000) 13–18 www.elsevier.nl/locate/surfcoat

A method for evaluating the creep properties of overlay coatings M.P. Taylor a, H.E. Evans a, *, C.B. Ponton a, J.R. Nicholls b a School of Metallurgy and Materials, The University of Birmingham, Birmingham B15 2TT, UK b School of Industrial and Manufacturing Science, Cranfield University, Cranfield, Bedford MK43 0AL, UK Received 8 March 1999; accepted in revised form 20 October 1999

Abstract A new method is described for the evaluation of the creep properties of as-deposited overlay coatings. The method uses a composite tensile specimen consisting of a core alloy of well-characterised creep properties onto which the overlay coating is deposited. The overall strain/time response of this specimen, tested under constant stress conditions, can then be deconvoluted to obtain the creep characteristics of the coating. The method has the advantage that the coatings are evaluated for the same microstructural conditions as they will be used under, i.e. having the same porosity levels and microstructure. A limited demonstration of the method has been made for an air plasma-sprayed Ni25Cr6AlY coating deposited onto an austenitic steel core and creep tested at 900°C. The creep strength of this coating was surmised to be higher than that of equivalent monolithic samples produced by low pressure plasma spraying. © 2000 Elsevier Science S.A. All rights reserved. Keywords: Creep; Overlay coatings

1. Introduction Oxidation-resistant metallic coatings are widely used to confer protection to superalloys in high-temperature applications [1]. They may be applied as simple overlays or as a component of more complex systems such as thermal barrier coatings. In all cases, the creep properties of the coating have a substantial influence on behaviour, e.g. through short-range stress relaxation at interfacial crack tips and oxide spallation [2] or, through largescale effects, on stress levels in other components of the coating system [3]. Clearly, the successful modelling of these processes and of the endurance of coating systems in general will require knowledge of the dependence of coating creep (or relaxation) rates on temperature, time and creep strain. Previously published [4,5] measurements of creep properties of coating alloys have been obtained on monolithic samples, i.e. complete creep specimens machined from bulkier samples of the coating alloy obtained by low-pressure plasma spraying (LPPS ). This approach has the obvious advantage that direct measurements can be made on the composition of interest, but the method is impractical for those coating processes, * Corresponding author. Tel.: +44-121-414-5172; fax: +44-1452-700-730.

such as chemical or physical vapour deposition, where it is more difficult or expensive to produce bulk deposits. A disadvantage, even for sprayed coatings, is that the monolithic specimens produced will tend to have a different density and microstructure from thinner coatings sprayed directly onto an alloy substrate and may, consequently, have different creep properties from actual coatings. In response to these concerns, a new method is presented in this paper of evaluating coating creep by deconvoluting the creep response of a composite coated specimen. The experimental study will use air plasmasprayed (APS) NiCrAlY coatings, but the data presented will be limited to demonstrating the validity of the method by comparison with previous results [4] on monolithic LPPS alloys of similar composition.

2. The method of creep evaluation The approach is to consider a coated tensile creep specimen of rectangular cross-section as shown schematically in Fig. 1. This particular specimen shape is desirable in order to ensure that the total coating thickness, 2z, is comparable with the thickness, h, of the core so that both make similarly weighted contributions to the overall creep response of the composite specimen. Provided also that the specimen width, w, is much larger

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mental work to be described later in this paper, a core of austenitic steel is used partly because this can be obtained as a thin-sectioned sheet but also because extensive data are available covering both steady-state [6 ] and primary [7] creep. It is particularly important to recognise that the latter, time-dependent creep rate should be considered in the analysis since the core will enter steady-state conditions only after some 10–15% strain under constant-stress conditions whereas, at some compositions, the coating may fail before this. A particularly successful algorithm for describing creep strain, e , during primary creep was first provided c by McVetty [8] and subsequently validated for Type 304 stainless steel by Garofalo et al. [7]: e =e [1−exp(−mt)]+e˙ t. (3) c T s Here, e is a constant, t is time under stress and m is T proportional to the steady-state creep rate, e˙ , i.e. with s the same temperature and stress dependencies [7]: (4) s where b is a numerical constant. The steady-state creep rate is given typically as: m=be˙

A

Q e˙ =Asn exp − s h RT

Fig. 1. A schematic cross-section of the coated specimen.

than its thickness, edge effects may be neglected. Under creep testing, the specimen will be subjected to a tensile gross stress, s , which partitions between core and g coating according to the balance of forces across the specimen cross-section: 2s zw+s hw=s (h+2z)w. (1) coat h g Here, s is the average uniaxial tensile stress in each coat of the coating layers and s is the corresponding stress h in the core of thickness h. At the creep temperature, it is unlikely that significant stress gradients exist across either the coating or the core. It is assumed, and will be justified later, that the composite specimen retains adherence whilst deforming uniformly in the tensile direction. Rearranging Eq. (1) gives s

coat

s (h+2z)−s h h . = g 2z

(2)

The coating stress, corresponding to the specimen strain rate, can then be determined provided that the core stress at the same strain rate is known. It follows that the key to establishing the coating properties by this method is to use an alloy for the core whose creep behaviour is well established. In the experi-

B

(5)

where A is a constant, n (≥1) is the creep index, Q is the activation energy for creep, R is the gas constant and T is absolute temperature. Since values for these parameters for the core will be known, as will the overall specimen creep rate and time under stress, their insertion into Eq. (3) will, in principle, provide the stress, s , h which must exist within the core to produce the observed creep rate at that time. This substitution leads to the specimen creep strain being described by:

G

C A A B

B DH

Q e =e 1−exp −bAsn exp − t c T h RT Q +Asn exp − t. h RT

(6)

At short times, i.e. in primary creep, this equation can only be solved for the core stress, s , by numerical or h graphical methods; the latter approach was used in the present work. This value of core stress can then be used in Eq. (2) to obtain the corresponding stress in the coating and this, of course, can then be obtained as a function of strain throughout the apparent primary creep of the composite specimen. At large values of t, however, the first term of Eq. (6) approaches e and T the core creep rate is then given by Eq. (5), i.e. steadystate conditions exist. In this case, an analytical solution of the equations can be obtained for the stress in the core and, by insertion into Eq. (2), the coating stress

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M.P. Taylor et al. / Surface and Coatings Technology 124 (2000) 13–18 Table 1 Composition (wt.%) of the austenitic steel core Cr

Ni

Mn

Si

Nb

(C+N )

Fe

19.2

25.3

0.6

0.56

0.7

≤0.045

Bal.

as:

s

coat

=

1 2z

G

G

s (h+2z)−h g

e˙

A

A exp −

Q RT

B

HH

1/n .

(7)

3. Experimental approach In order to demonstrate the above method for evaluating coating creep properties, a core material of a 20Cr25Ni austenitic steel was used. This had the composition given in Table 1 and was obtained as a coldworked sheet of 0.38 mm thickness. Tensile specimens, of gauge dimensions 25.4 mm by 6.4 mm, were cold-stamped from this material and subsequently air plasma-coated on each of the major surfaces with an alloy of composition (wt.%) 68Ni, 25Cr, 6Al, 0.4Y. The coating thickness on each of these surfaces was approximately 0.15 mm so that, as discussed above, the total coating thickness was comparable to the core thickness. On all specimens, coating thickness was evaluated prior to testing using a micrometer but, because of the irregularity of the coating surface, this procedure tended to overestimate the actual thickness. Scoping calculations, using the procedure described

above, demonstrated, however, that the strain rate/stress behaviour reported below was not significantly affected by errors in coating thickness of around 5%. Accordingly, these macroscopic measurements of coating thickness were used throughout and are detailed in Table 2. Prior to testing, the samples were annealed in argon for 1 h at 930°C to effect recrystallisation of the steel core and to produce a sensibly uniform grain size of some 10 to 15 mm. Uniaxial creep tests were undertaken under constant stress conditions in air at 900°C. These were short-term, relatively high-stress, tests designed to produce strain rates of around 10−5 s−1, typical of those local to interface cracks during cooling [2]. Specimen length was monitored during the tests using capacitance transducers.

4. Results and discussion A representative creep curve for a coated specimen tested at a gross stress of 42.4 MPa is shown in Fig. 2. This response of the composite specimen is a combination of the creep behaviour of both the coating and the steel core. It is unlikely that each component will have the same intrinsic strain/time characteristics over the length of the test, and so the actual stress in the coating and in the core will vary with strain even though the composite specimen, as a whole, experiences a constant stress. As can be seen, the overall effect is to produce an initial period of decreasing creep rate extending to approximately 14.5% creep strain. A pseudo-steady state then precedes a short tertiary stage and failure at about 18.5% strain.

Table 2 Details of the creep tests Gross stress (MPa)

Total coating thickness (mm)

Creep strain (%)

Time (ks)

Creep rate (s−1)

Core stress (MPa)

Coating stress (MPa)

28.0

0.271

42.4

0.266

56.0

0.315

13 10 8 6 4 2 13 10 8 6 4 2 13 10 8 6 4 2

49.36 29.06 18.03 9.62 4.03 1.09 8.32 5.72 4.26 2.98 1.80 0.78 2.60 1.87 1.38 0.89 0.52 0.15

1.15×10−6 1.75×10−6 2.23×10−6 3.28×10−6 4.26×10−6 6.83×10−6 1.05×10−5 1.30×10−5 1.47×10−5 1.66×10−5 1.98×10−5 2.25×10−5 3.74×10−5 4.44×10−5 4.52×10−5 5.36×10−5 5.59×10−5 5.81×10−5

29.2 27.8 28.2 29.8 30.4 32.9 44.2 45.0 44.4 43.4 43.1 42.5 58.5 57.5 55.5 54.6 52.2 50.5

26.3 28.3 27.7 25.4 24.5 20.9 39.6 38.5 39.4 40.8 41.3 42.1 53.0 54.2 56.6 57.7 60.5 62.5

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M.P. Taylor et al. / Surface and Coatings Technology 124 (2000) 13–18

Fig. 2. The creep curve for a composite specimen tested under a gross constant stress of 42.4 MPa at 900°C.

A metallographic cross-section of this failed specimen is shown in Fig. 3. It is quite clear that bonding between coating and core has been preserved apart from regions in the vicinity of the failure site where a crack has formed through the thickness of the coating and some interfacial decohesion is also present. This maintenance of a good coating/core bond over the period when the composite specimen is deforming uniformly is an implicit requirement of the present analytical approach. Estimates of the specimen strain rates for the three tests undertaken in this demonstration were obtained from the appropriate creep curve and these are given for various creep strains in Table 2. The corresponding values of stress in the austenitic steel core were obtained through use of Clay’s [6 ] equation for steady-state creep: e˙ =210s5 exp s h

A

−42096 T

B

where stress, s , is in megapascals and temperature, T, h in kelvin and, as appropriate, Garofalo’s [7] value of b=37 in Eqs. (3) and (4). The stress in the core for various creep times was calculated by solving Eq. (6) graphically or, at higher strains, by using Eq. (8) in Eq. (7). These calculated values of core stress are given in Table 2 together with the corresponding values of coating stress obtained from the force balance of Eq. (2). The dependence of coating strain rate with coating stress for strains ≥8% is shown in Fig. 4. Also shown as broken lines is the steady-state creep behaviour of monolithic specimens of a Ni16Cr6AlY and of a Ni35Cr6AlY coating alloy obtained by Brindley and Whittenberger [4] by a combination of stress relaxation and compressive

(8)

Fig. 3. A metallographic cross-section of a composite specimen creep tested to fracture at a stress of 42.4 MPa at 900°C.

Fig. 4. The dependence of steady-state creep rate of the present Ni25Cr6AlY coating with stress, s , at 900°C. Also shown as broken coat lines are the corresponding data obtained by Brindley and Whittenberger [4] on monolithic samples of Ni16Cr6AlY and Ni35Cr6AlY.

M.P. Taylor et al. / Surface and Coatings Technology 124 (2000) 13–18

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steady-state conditions exist. The curvature associated with these plots is consistent with the presence of a threshold stress, s , such that the steady-state creep rate o can be described by:

Fig. 5. Steady-state creep rates (i.e. for strains of 8–13%) at 900°C in the present coatings can be described by a reduced-stress formulation as shown by the solid line. This corresponds to a threshold stress of 12.6 MPa. At lower strains and low stresses, creep rates are higher than expected from this formulation.

creep. The present results on the Ni25Cr6AlY coating coincide with their best-estimate line for the Ni16Cr6AlY alloy. It might be expected, however, that if the current alloy with 25% Cr were to follow the trend shown in the figure of a reduction in creep strength with increasing chromium content, its intrinsic behaviour should lie somewhere between the two curves shown. Although no direct comparison can be made with a monolithic alloy of the same composition, it is tempting to suggest that the increased volume fraction of dispersed oxides produced in the present coating by air plasma spraying compared with LPPS has, indeed, produced a strengthening effect in subsequent creep. The dependence of creep rate on stress obtained in the present coatings is insensitive to creep strain for values ≥8%, as shown in Fig. 4, and indicates that

e˙ =A(s−s )p (9) s o where A is a constant for a given temperature and the exponent, p, has typical values between 2 and 4 [9]. For the present results, agreement is obtained, for strains ≥8%, with p=3, A=5.876×10−10 MPa−3 s−1 and s =12.6 MPa, as shown in Fig. 5. The quality of fit is o good with a correlation coefficient, r2, of 0.984 and a standard deviation in s of only ±1.6 MPa. o Various models have been developed to account for such creep/threshold effects [9–11]. They differ substantially in detail but all relate the threshold condition to the minimum stress required for a dislocation to bypass a dispersed particle by climb and then to break away from the particle. In alloys intentionally strengthened by second-phase particles, threshold stresses are much higher than deduced here, e.g. 50 MPa in an austenitic steel containing TiN particles [9] and 100 MPa for a c∞-strengthened superalloy (IN738LC ) [12]. Insofar that the threshold stress value found in the present coatings is much less than these, the results are consistent with the properties of a relatively weak alloy strengthened in part by an adventitious dispersion of oxide particles produced during the spraying process. At strains <8%, departures from the reduced-stress formulation of Eq. (9) arise at low stresses as shown in Fig. 5 for creep strains of 2 and 4%. It seems likely that this represents primary creep of the coating alloy. By interpolation of straight lines on the log–log plot of Fig. 5 drawn for given creep strains <8%, the dependence of creep rate on strain can be obtained for a given coating stress. Examples of this variation at stresses of 30 and 50 MPa are shown in Fig. 6 where, in addition for strains ≥8%, steady-state creep was assumed according to Eq. (9). It is clear that, at the lower stress, there appears to be a substantial primary creep stage; at the higher stress, there is little reduction of creep rate with strain and, to a reasonable approximation, steady-state conditions can be thought to exist from strains as small as 2%. One consequence of this stress dependence of the duration of primary creep is that at these lower strains, appropriate to the modelling of thermal loading in coated systems, a creep algorithm of the simple form of Eq. (5) may be used.

5. Conclusions

Fig. 6. The variation of strain rate with strain in the present coatings at 30 MPa and 50 MPa showing a more marked primary creep stage at the lower stress.

A new method has been described for evaluating the creep behaviour of overlay coatings. Composite tensile specimens are used which consist of a core alloy of wellcharacterised creep properties onto which the overlay coating is deposited. The overall strain/time response of this specimen, tested under constant stress conditions,

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M.P. Taylor et al. / Surface and Coatings Technology 124 (2000) 13–18

can then be deconvoluted to obtain the creep characteristics of the coating. The method has the advantage, over testing of monolithic samples of the coating alloy, that the overlay is evaluated for the same microstructural conditions as it will be used under. A limited demonstration of the method has been made for an air plasma-sprayed Ni25Cr6AlY coating deposited onto an austenitic steel core and creep tested at 900°C. The creep strength of this coating was surmised to be higher than that of equivalent monolithic samples produced by low-pressure plasma spraying, but the general stress/strain rate behaviour was similar. This comparison provides a limited validation of the present method. Acknowledgement This work was funded by the Engineering and Physical Sciences Research Council under Grants GR/K97660 and GR/K97882.

References [1] S.R.J. Saunders, J.R. Nicholls, Mater. Sci. Technol. 5 (1989) 780. [2] H.E. Evans, A. Strawbridge, R.A. Carolan, C.B. Ponton, Mater. Sci. Eng. A 225 (1997) 1. [3] W.J. Brindley, Workshop on Thermal Barrier Coatings, NASA, Cleveland, OH, 1995. [4] W.J. Brindley, J.D. Whittenberger, Mater. Sci. Eng. A 163 (1993) 33. [5] M.G. Hebsur, R.V. Miner, Thin Solid Films 147 (1987) 143. [6 ] B.D. Clay, Mechanical Behaviour and Nuclear Applications of Stainless Steels at Elevated Temperatures, Metals Society, London, 1982, p. 122. [7] F. Garofalo, O. Richmond, W.F. Domis, F. von Gemmingen, Joint International Conference on Creep, Institute of Mechanical Engineers, London, 1963, p. 1. [8] P.G. McVetty, Mech. Eng. 56 (1934) 149. [9] H.E. Evans, I. Beden, R.C. Ecob, Proc. Roy. Soc. London, Ser. A 404 (1986) 339. [10] E. Arzt, D.S. Wilkinson, Acta Metall. 34 (1986) 1893. [11] R.S. Mishra, T.K. Nandy, G.W. Greenwood, Philos. Mag. 69 (1994) 1097. [12] P.J. Henderson, M. McLean, Acta Metall. 31 (1983) 1203.