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SiC composite at elevated temperature
Composites Science and Technology 61 (2001) 1331–1338 www.elsevier.com/locate/compscitech
Notch sensitivity of fatigue life in a Sylramic composite at elevated temperature
TM
/SiC
J.C. McNulty, M.Y. He, F.W. Zok* Materials Department, University of California, Santa Barbara, CA 93106, USA Received 16 March 2000; received in revised form 20 February 2001; accepted 7 March 2001
Abstract The effects of holes and notches on the fatigue life of an advanced SylramicTM/SiC composite at 815 C have been examined. At this temperature, fracture occurs by an oxidative embrittlement mechanism, common to most SiC-based composites. In unnotched specimens, embrittlement is manifested at stresses above the matrix cracking limit, mc , leading to fracture following relatively short exposure times ( 100 h). As a consequence, a fatigue threshold is obtained at a stress, th mc . This threshold is due to the absence of an easy path for oxygen ingress when matrix cracks are not present. In center-hole specimens, an analogous threshold is obtained, at a stress, th mc =ke , where ke is the elastic stress concentration factor (= 2.5). That is, once the cracking limit is exceeded at the hole edge, embrittlement and fracture ensue. The threshold stress for center-notch specimens with stress concentration ke ¼ 7 is numerically similar to that of the center-hole specimens with ke ¼ 2:5, indicating some tolerance to local stress levels above the global matrix cracking limit in sharply notched geometries. Non-linear finite-element calculations of the stresses in the center-hole and center-notch specimens are used to infer the local conditions associated with the threshold. A key result is that the damage tolerance and notch insensitivity normally associated with inelastic straining cannot be exploited at temperatures at which the embrittlement mechanism operates. The implication is that composite structures with holes and notches must be designed extremely conservatively to ensure long lifetimes (> 100 h). # 2001 Elsevier Science Ltd. All rights reserved. Keywords: Notch sensitivity; Fatigue; SiC composite; Embrittlement
1. Introduction Continuous-fiber ceramic composites (CFCCs) based on SiC constituents have been under development for the past two decades, mainly for use in advanced gas turbine engines. The motivation for this activity is the desire to increase operating temperatures and hence improve engine efficiency and performance. Additional benefits associated with the elimination of film cooling of combustor liners and the ensuing reductions in NOx emissions have also been identified [1,2]. The selection of SiC as the main constituent in highperformance CFCCs is based on a number of attractive physical and mechanical characteristics. Notably, their low diffusivity results in good creep resistance at high temperatures. Furthermore, their high thermal conductivity and low thermal expansion coefficient lead to high thermal shock resistance. In addition, the SiC* Corresponding author. Tel.: +1-805-893-8699; fax: +1-805-8938486. E-mail address: [email protected] (F.W. Zok).
based CFCCs exhibit inelastic deformation associated with matrix cracking and interface debonding, providing damage tolerance and strength retention in the presence of holes and notches [3–5]. The most significant problem hindering SiC–CFCCs is oxidation embrittlement. The embrittlement most commonly occurs by oxygen ingress through matrix cracks, followed by reaction of the oxygen with the fibers and the fiber coatings [6–22]. The problem is particularly severe at intermediate temperatures (500–900 C) [6,15,23–34]. It is expected to be exacerbated by cyclic loading, since under such conditions the reaction gases contained within the crack are expelled during unloading and the oxidizing atmosphere drawn into the composite through matrix cracks during reloading. Cyclic loading may also accelerate fiber fracture in regions where the fibers have bonded to the surrounding matrix as a consequence of the oxidation process, thereby limiting the extent of further sliding along the fiber/matrix interface. The present paper examines the embrittlement phenomenon in an advanced SiC–CFCC which is currently of interest to both the aerospace and the land-base
0266-3538/01/$ - see front matter # 2001 Elsevier Science Ltd. All rights reserved. PII: S0266-3538(01)00032-X
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power-generation industries. Experiments are performed to assess the fatigue performance at a moderately high temperature (815 C). This temperature falls in the regime in which embrittlement has been reported for other SiCbased CFCCs, yet is also relevant to the service environments in which CFCC components are expected to operate in turbine engine applications. The sensitivity of fatigue life to stress concentrations is probed through tests on specimens with circular holes and notches as well as straight (unnotched) specimens. The study of the effects of stress concentrations is motivated by the need, in some instances, to use through-holes for either attachment or cooling in engine components. In a broader context, it might also be used to infer the extent of degradation due to foreign impact damage, assuming at the simplest level that the damaged region is unable to sustain load and hence acts similarly to a hole or notch. Some additional insights into the local conditions required for embrittlement and fracture in the notched geometries are obtained through calculations of the stresses in the notched geometries, by the use of a non-linear constitutive law appropriate to CFCCs.
2. Materials and experimental measurements All experiments were performed on a composite comprising SylramicTM SiC1 fibers in a plain-weave architecture and a SiC matrix produced by a proprietary hybrid process involving chemical vapor infiltration (CVI) and reactive melt infiltration (MI). The microstructure of the composite is shown in Fig. 1. The notable features include the presence of a uniform BN coating on the fibers, an ‘overcoat’ of SiC around the fiber tows (produced by CVI) and a SiC-based two-phase matrix (produced by reactive MI). The matrix is essentially fully dense, without the large pores that are invariably present in CFCCs made by CVI alone. The absence of pores in the matrix has two consequences. First, it increases the stress at the onset of matrix cracking relative to that of the CVI CFCCs containing pores. Secondly, it is expected to improve the thermal conductivity, especially in the direction transverse to the fibers. The fiber volume fraction is 35%. To assess the effects of stress concentrators on low cycle fatigue (LCF), three types of specimens were used: (i) standard dog-bone (unnotched) specimens; (ii) straight tensile specimens of width 2W=31.8 mm, with a center hole of diameter 2a=6.35 mm ða=W ¼ 0:2Þ; and (iii) straight tensile specimens of width 2W=31.8 mm, with a center notch of length 2a=6.35 mm ða=W ¼ 0:2Þ and notch root radius 0:2 mm. The specimens were prepared by electrodischarge machining. Four specimens of each geometry were tested at a temperature of 1
Produced by Dow Corning.
Fig. 1. Scanning electron micrographs of a cross-section through the composite: (a) at low magnification, showing the longitudinal and transverse fiber tows; (b) at higher magnifications, showing the meltinfiltrated matrix surrounding a fiber tow, and (c) the interior of the tow. The micrographs were taken in backscatter electron imaging (BEI) mode, thereby revealing atomic number contrast. The coatings on the fibers are BN. The SiC within the tows was produced by a chemical vapor infiltration process, prior to melt-infiltration of the remainder of the matrix.
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815 C in ambient air. The specimens were heated using an induction heating system and a steel susceptor. Temperatures were measured using thermocouples at three locations on the specimen surface: at the center and at points located 20 mm from the center. The temperatures at these three locations were maintained within 5 C of the targeted temperature (815 C) for the duration of each test. Steel tabs were affixed to the specimen ends using a high temperature ceramic adhesive and the tabbed sections then inserted into hydraulic wedge grips. Additionally, several tests were performed at room temperature on straight (unnotched) specimens as well as center hole specimens with hole diameters of either 2a=3.18 or 6.35 mm and with a=W ¼ 0:2. The latter tests were used to assess the extent of notch sensitivity under ambient conditions. These results formed the basis on which the notch sensitivity of LCF life was subsequently assessed. The elevated temperature tests included one monotonic tension test on each specimen type at a stressing rate of d=dt ¼ 100 MPa/min, typically leading to fracture in 43 min. The remaining three specimens were subject to LCF tests with a stress ratio of R=0.01. The loading spectrum comprised relatively rapid loading and unloading (each taking 1 min), and a hold time of 2 h at the peak stress. This spectrum was chosen to simulate the loading conditions that a gas turbine engine component might encounter in service. The peak stress levels were selected on the basis of the monotonic tensile test results. In cases in which fracture did not occur, the tests were terminated after 1 week ( 160–170 h) of loading. The room temperature tests were also performed at a stressing rate of 100 MPa/min. The monotonic tensile test results on the unnotched specimens are plotted in Fig. 2. At both 20 and 815 C, the tensile response is essentially bilinear, with the transition in slope occurring at the onset of matrix cracking
Fig. 2. Tensile response of unnotched specimens at both 20 and 815 C. Note the strong similarity in the stress-strain response at the two test temperatures, the key difference being the strains at which fracture occurs. The solid lines represent a bilinear fit to the data, used for calibrating the nonlinear constitutive law.
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at a stress of 150–165 MPa. There was no apparent effect of temperature on this response, except that the fracture strain at 815 C was slightly lower than that at 20 C. Optical microscopy of polished cross-sections through the fractured test specimens revealed periodic matrix cracks, as commonly observed in this class of CFCC. The effects of holes on the net-section strength at room temperature are shown in Fig. 3. The strength decreases with increasing hole size, to 70% of the unnotched strength at 2a=6.3 mm. Similar reductions have been reported previously for SiC/SiC, SiC/MAS [5] and all-oxide ceramic composites [35]. These effects can be rationalized on the basis of the stress redistribution due to inelastic straining around the hole as well as the size-scale dependence of the local conditions needed to precipitate fracture, as described in Section 4. The LCF behavior of 815 C is shown in Fig. 4. Several notable features emerge. (i) The monotonic tension tests on the center-hole and center-notch geometries yield strengths similar to one another, c 160–165 MPa,
Fig. 3. Effects of center holes on room-temperature tensile strength. The solid line represents the prediction of the point stress fracture model for a characteristic distance of d=0.75 mm.
Fig. 4. Summary of LCF life at 815 C. Arrows pointing to the right indicate runout; arrows pointing to the left indicate test results obtained from monotonic tension tests.
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despite large differences in the elastic stress concentration factors: ke=2.5 vs. 7.1 (see Appendix). These strengths are 60% of the unnotched monotonic strength, u 270 MPa, measured at 815 C. Evidently the degree of notch sensitivity at this temperature is only slightly greater than that at 25 C. (ii) In the unnotched specimens, a fatigue threshold (defined by tf 5 160h) was obtained at a stress O level, th 165 MPa, corresponding closely to the matrix cracking limit: mc 150–160 MPa (Fig. 2). Similar correlations between the threshold stress and the matrix cracking stress had been established previously for static loading of a CVI SiC/SiC at a temperature of 900 C [31]. However, the cracking stress of the latter material was considerably lower (60–70 MPa), a consequence of the matrix porosity. (iii) In the center-hole and center-notch specimens, the threshold was considerably lower, th 60–85 MPa, representing 40–50% of the unnotched threshold. Essentially the same behavior was obtained with notches and holes. (iv) For all three specimen geometries, the stress rupture curves were extremely shallow, with the exception of the regime at very short fracture times (tf << 1 h). Evidently fracture occurs very rapidly at stresses above the threshold level. The implication is that life prediction at stresses above the threshold is likely to be exceedingly difficult.
3. Fractography Representative fractured test specimens were examined in a scanning electron microscope (SEM). Typical features are shown in Figs. 5 and 6. Specimens subjected to monotonic tensile loading at elevated temperature exhibited little fiber pullout (Fig. 5), suggesting a relatively high interfacial strength. However, there was no indication of oxidation on the fracture surfaces following the short exposure times ( 3 min) of these test specimens. Similarly low levels of pullout were observed on the fracture surfaces of specimens tested at room temperature. By contrast, the LCF specimens exposed to elevated temperature for many hours exhibited significant levels of oxidation on the fracture plane. Fig. 6 shows micrographs of a center-hole specimen, tested at a peak stress, p ¼85 MPa. The regions of the fracture surface near the hole edge exhibited particularly severe oxidation, as manifest by the presence of a glassy layer on the axial fiber tows. Similar features were also observed on the center-notch specimens. The inference from the presence of the glassy layer on the fiber fracture surfaces is that the fibers are failing progressively, starting near the tip of the notch or hole (where the stress concentration is at a maximum) and proceeding through the remaining (unnotched) section of the material. This sequence is consistent with that observed in other SiC-based CFCCs, as detailed in [15].
Fig. 5. Fracture surface of a center hole specimen tested under monotonic tension at 815 C.
4. Modeling of notch sensitivity In light of the extremely shallow stress rupture curves, the modelling activity focused predominantly on the effects of notches and holes on the threshold stress levels. A secondary priority was the notch sensitivity of the monotonic tensile strength at ambient temperature. 4.1. Notch sensitivity of monotonic tensile strength The notch sensitivity of strength under ambient (nonoxidizing) conditions is dictated by two factors: (i) the extent to which inelastic straining mitigates the stress concentration at the notch tip, and (ii) the dependence of the fracture condition on the stress gradients that exist ahead of the notch. The former effects can be calculated by finite element methods using a non-linear constitutive law appropriate to CFCCs. In the present study, the law developed by Genin and Hutchinson [36] has been used for such calculations. This constitutive law is based on a phenomenological description of the development of inelastic strain in cross-ply CFCCs under biaxial stressing. The pertinent functions are
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calibrated using the results of tensile tests performed both along one of the fiber axes (in the 0 /90 orientation) and at 45 to the fiber axes. The constitutive law is then implemented in a finite-element code for the purpose of calculating stress and strain fields in non-uniform specimen or component geometries. The capability of the model to predict the latter response has been validated through comparisons with strain measurements on specimens containing notches and circular holes [5,35]. The model tacitly assumes that the material response is scale-independent. This presents some limitations on the size-scale of features that can be accurately simulated, as discussed in Section 4.2. Additional details of the implementation and calibration procedures are described in Refs. [36] and [5]. The effects of stress gradients on fracture are less well understood. Nevertheless, a simple fracture criterion based on the attainment of a critical stress (taken as the unnotched strength) over a characteristic distance d along the incipient fracture plane has been found to adequately describe notched strength of CFCCs, provided the effects of inelastic straining on the stress distribution are taken
into account [5,35]. Furthermore, previous studies have revealed that the characteristic distance inferred from the model and the experiment is rather insensitive to the details of the composite microstructure, falling in the narrow range of 0.5–0.75 mm for SiC/SiC [5], SiC/ MAS [5] and an all-oxide CFCC [35]. This approach has been used to model the strength of the center-hole specimens of the SylramicTM/SiC composite tested at room temperature, using a characteristic distance, d=0.75 mm. The effects of inelastic straining on the stress distribution ahead of a hole are illustrated by the results plotted in Fig. 7(a). The strength predictions are plotted in Fig. 3, for comparison with the experimental measurements. Excellent agreement between experiment and theory is obtained. Furthermore, the inferred characteristic distance is remarkably similar to the values obtained in other CFCCs, indicating a commonalty in the factors controlling fracture of the various composite systems. The same approach has been used to rationalize the effects of both notches and holes on the strength at 815 C. In this case, the characteristic distance was inferred
Fig. 6. Fracture surface of a center hole specimen tested in fatigue at p ¼ 85 MPa at 815 C.
Fig. 7. Typical stress distributions for (a) center-hole and (b) centernotched specimens of the SylramicTM/SiC composite, calculated using the nonlinear constitutive law of Genin and Hutchinson [36]. Also shown for comparison are the corresponding elastic predictions. The results demonstrate the role of inelastic straining in mitigating stress concentrations.
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Fig. 8. Comparisons of the measured and predicted monotonic tensile strengths of specimens with holes and notches at 815 C. The strength is underestimated using the elastic stress concentration factor in combination with a critical stress fracture criterion, i.e. n ¼ u =ke . In contrast, the model based on the inelastic stress distribution coupled with the point-stress criterion provides an adequate representation, for d 0.3–0.5 mm.
from comparisons of the measured and predicted notched strengths, assuming various values for d in the model. The comparisons are presented in Fig. 8. The inferred characteristic distance is d 0.3–0.5 mm, considerably smaller than the value obtained from the room temperature tests (d 0.75 mm). This reduction in d reflects a degradation in notch-sensitivity. It is unknown whether this degradation is intrinsic to the fracture mechanism at elevated temperature, it reflects the onset of oxidative embrittlement, or it is associated with scatter in the experimental data. 4.2. Notch sensitivity of the threshold stress An analogous stress-based fracture model has been used to rationalize the threshold stress levels in the center-notch and center-hole specimens. The threshold stress was taken to be that for which the longitudinal stress, , along the incipient fracture plane exceeds the o unnotched threshold, th 165 MPa, over a characteristic distance, dth, ahead of the hole or notch. As before, the characteristic distance was obtained by comparing the calculated results with the experimental measurements. Such comparisons are plotted in Fig. 9. For the center-hole specimens, the inferred distance spans the range dth 00:2 mm: the variation arising from the difference between the lowest stress at which fracture had occurred, 85 MPa, and the highest stress at which run-out was obtained, 60 MPa. Such low values of dth are indicative of extremely strong notch sensitivity. Indeed, in the limit of dth=0, the threshold corresponds to the case where the maximum stress (at the hole edge) is equal to the unnotched threshold stress. Since the unnotched threshold is dictated by the matrix cracking stress and the material response up to the cracking limit is elastic, a conservative estimate of the threshold can be obtained using the elastic stress concentration factor ke
Fig. 9. Comparisons of measured and predicted thresholds at 815 C for (a) center-hole and (b) center-notch specimens, based on the nonlinear finite element model and the point-stress fracture criterion.
in combination with the matrix cracking stress, whereupon the predicted threshold becomes o th ¼ th =ke ¼ mc =ke
ð1Þ
This represents a lower-bound estimate of the threshold stress. For the center-hole specimen with a=W ¼ 0:2, ke=2.5 [37] (Appendix) and thus th 66 MPa [also plotted on Fig. 9(a)]. This prediction is in reasonable agreement with the range obtained from the experiments (60–85 MPa), reaffirming the poor notch sensitivity associated with LCF fracture at the test temperature. Furthermore, the extreme sensitivity of the fracture time to the presence of matrix cracks is consistent with the oxidation embrittlement characteristics of other SiC-based CFCCs at intermediate temperatures [12,15,17,22,23,27,38]. For the center-notch specimens, the inferred characteristic distance is somewhat larger, dth 0:30:4 mm, indicating improved damage tolerance and lower notch sensitivity. Furthermore, the lower bound estimate of the threshold stress, th 23 MPa, based on
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Eq. (1) and the elastic stress concentration factor, ke=7.1 [39], is significantly below the measured value, 70–85 MPa. This difference is believed to be associated with the steep stress gradients that exist in the notch tip region coupled with the intrinsic volume dependence of the matrix strength. That is, the high stress levels exist over only small volumes ahead of the notch (over distances ), reducing significantly the probability of encountering a matrix flaw to initiate cracking. Finite element calculations for the center-notched geometry used in the present study (with 2a=6.35 mm and =0.2 mm) reveal that the elastic stress diminishes rapidly with distance from the notch tip, by 50% at a distance of 0.1 mm [Fig. 7(b)]. (By contrast, the reduction in elastic stress in the center-hole specimen with a hole diameter of 6.35 mm is only 6% at the same distance from the hole edge.) It is surmised that even a modest volume dependence of strength might lead to a significant elevation in the maximum local stress at the onset of cracking. This elevation in the local cracking stress may account for the rather high threshold of the notched specimens. If so, the beneficial effects of inelastic straining on the stress concentration may be overestimated by the non-linear constitutive law used in this study, because, in essence, the volume dependence of the cracking stress is neglected. A further implication is that the inferred characteristic distance (dth 0.3–0.4 mm) may be anomalously high. It remains to be established which of these effects is dominant.
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times in the post-cracking regime are not likely to be useful, because of the precipitous drop in fracture time to unacceptably low levels. The LCF threshold for center-hole specimens is similarly dictated by the matrix cracking limit. Once the cracking stress is exceeded locally (even over a very small distance ahead of the hole), fracture ensues. A reasonably accurate estimate of the threshold is obtained through the elastic stress concentration factor in combination with the matrix cracking stress [Eqn. (1)]. This highly notch-sensitive behavior suggests the need for an extremely conservative design approach for CFCC structures containing holes. Furthermore, it precludes the exploitation of the inelastic deformation mechanisms that operate in these materials and impart the requisite damage tolerance in non-oxidizing environments. The LCF threshold for center-notch specimens is comparable to that of the center hole specimens. Recognizing that the stress concentration factor is higher in the notched geometry, the composite appears to exhibit superior damage tolerance in the notched geometry. However, this apparent damage tolerance may be a result of the volumedependence of the matrix cracking stress, not a consequence of local inelastic straining. Further experiments are needed to determine the statistical parameters associated with the matrix strength and to assess their role in the conditions for fracture ahead of a notch. Additional experiments are also needed to assess the extent of scatter in fatigue life, for the purpose of validating the failure models presented here.
5. Summary The notch sensitivity of the tensile strength of the SylramicTM/SiC composite at ambient temperature is consistent with the trends observed in other CFCCs. Typically a reduction of 25% in net-section strength is obtained for open holes of diameter 6 mm. This modest amount of notch sensitivity can be rationalized on the basis of the stress redistribution that occurs as a result of local inelastic straining and the size-scale dependence of the conditions at the onset of fracture through the point stress fracture criterion. The notch sensitivity at elevated temperature appears to be reduced slightly, as manifest in a reduction in the inferred characteristic distance in the point stress criterion. The unnotched LCF threshold for the SylramicTM/SiC composite at 815 C is dictated by the matrix cracking stress. Above the cracking stress, fracture occurs rapidly, typically in 4100 h. Moreover, the LCF life curve for stresses slightly above the threshold is extremely shallow, indicating a very strong sensitivity of fracture time to applied stress. For the lifetime requirements of most gas turbine engine components (103–104 h), the allowable stress levels would have to remain below this threshold. A further implication is that models to predict fracture
Appendix. Elastic stress concentrations in notched tensile specimens The elastic stress concentration factors, ke ; were calculated assuming that the composite is elastically isotropic. This assumption can be justified on the basis that the Young’s moduli measured in the 0 /90 and 45 orientations were within 2% of one another (250 GPa). This behavior is attributable to the fact that both the matrix and the fibers are essentially pure SiC. For a specimen of width 2W and with a center hole of diameter 2a, ke is given by [37]: ke
o a 3 ¼2þ 1 W net
ðA1Þ
where o is the maximum longitudinal stress (at the hole edge) and net is the applied net-section stress. For the center hole geometry used in the present experiments, a=W ¼ 0:2 and thus ke =2.51. The elastic stress concentration factor for a specimen of width 2W and with a center notch of length 2a and root radius (with a= 1) is given by [39]:
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ke
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o a ¼ 2 1 W net
rffiffiffiffiffi a
a1=2 sec W
ðA2Þ
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Notch sensitivity of fatigue life in a Sylramic composite at elevated temperature
TM
/SiC
J.C. McNulty, M.Y. He, F.W. Zok* Materials Department, University of California, Santa Barbara, CA 93106, USA Received 16 March 2000; received in revised form 20 February 2001; accepted 7 March 2001
Abstract The effects of holes and notches on the fatigue life of an advanced SylramicTM/SiC composite at 815 C have been examined. At this temperature, fracture occurs by an oxidative embrittlement mechanism, common to most SiC-based composites. In unnotched specimens, embrittlement is manifested at stresses above the matrix cracking limit, mc , leading to fracture following relatively short exposure times ( 100 h). As a consequence, a fatigue threshold is obtained at a stress, th mc . This threshold is due to the absence of an easy path for oxygen ingress when matrix cracks are not present. In center-hole specimens, an analogous threshold is obtained, at a stress, th mc =ke , where ke is the elastic stress concentration factor (= 2.5). That is, once the cracking limit is exceeded at the hole edge, embrittlement and fracture ensue. The threshold stress for center-notch specimens with stress concentration ke ¼ 7 is numerically similar to that of the center-hole specimens with ke ¼ 2:5, indicating some tolerance to local stress levels above the global matrix cracking limit in sharply notched geometries. Non-linear finite-element calculations of the stresses in the center-hole and center-notch specimens are used to infer the local conditions associated with the threshold. A key result is that the damage tolerance and notch insensitivity normally associated with inelastic straining cannot be exploited at temperatures at which the embrittlement mechanism operates. The implication is that composite structures with holes and notches must be designed extremely conservatively to ensure long lifetimes (> 100 h). # 2001 Elsevier Science Ltd. All rights reserved. Keywords: Notch sensitivity; Fatigue; SiC composite; Embrittlement
1. Introduction Continuous-fiber ceramic composites (CFCCs) based on SiC constituents have been under development for the past two decades, mainly for use in advanced gas turbine engines. The motivation for this activity is the desire to increase operating temperatures and hence improve engine efficiency and performance. Additional benefits associated with the elimination of film cooling of combustor liners and the ensuing reductions in NOx emissions have also been identified [1,2]. The selection of SiC as the main constituent in highperformance CFCCs is based on a number of attractive physical and mechanical characteristics. Notably, their low diffusivity results in good creep resistance at high temperatures. Furthermore, their high thermal conductivity and low thermal expansion coefficient lead to high thermal shock resistance. In addition, the SiC* Corresponding author. Tel.: +1-805-893-8699; fax: +1-805-8938486. E-mail address: [email protected] (F.W. Zok).
based CFCCs exhibit inelastic deformation associated with matrix cracking and interface debonding, providing damage tolerance and strength retention in the presence of holes and notches [3–5]. The most significant problem hindering SiC–CFCCs is oxidation embrittlement. The embrittlement most commonly occurs by oxygen ingress through matrix cracks, followed by reaction of the oxygen with the fibers and the fiber coatings [6–22]. The problem is particularly severe at intermediate temperatures (500–900 C) [6,15,23–34]. It is expected to be exacerbated by cyclic loading, since under such conditions the reaction gases contained within the crack are expelled during unloading and the oxidizing atmosphere drawn into the composite through matrix cracks during reloading. Cyclic loading may also accelerate fiber fracture in regions where the fibers have bonded to the surrounding matrix as a consequence of the oxidation process, thereby limiting the extent of further sliding along the fiber/matrix interface. The present paper examines the embrittlement phenomenon in an advanced SiC–CFCC which is currently of interest to both the aerospace and the land-base
0266-3538/01/$ - see front matter # 2001 Elsevier Science Ltd. All rights reserved. PII: S0266-3538(01)00032-X
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power-generation industries. Experiments are performed to assess the fatigue performance at a moderately high temperature (815 C). This temperature falls in the regime in which embrittlement has been reported for other SiCbased CFCCs, yet is also relevant to the service environments in which CFCC components are expected to operate in turbine engine applications. The sensitivity of fatigue life to stress concentrations is probed through tests on specimens with circular holes and notches as well as straight (unnotched) specimens. The study of the effects of stress concentrations is motivated by the need, in some instances, to use through-holes for either attachment or cooling in engine components. In a broader context, it might also be used to infer the extent of degradation due to foreign impact damage, assuming at the simplest level that the damaged region is unable to sustain load and hence acts similarly to a hole or notch. Some additional insights into the local conditions required for embrittlement and fracture in the notched geometries are obtained through calculations of the stresses in the notched geometries, by the use of a non-linear constitutive law appropriate to CFCCs.
2. Materials and experimental measurements All experiments were performed on a composite comprising SylramicTM SiC1 fibers in a plain-weave architecture and a SiC matrix produced by a proprietary hybrid process involving chemical vapor infiltration (CVI) and reactive melt infiltration (MI). The microstructure of the composite is shown in Fig. 1. The notable features include the presence of a uniform BN coating on the fibers, an ‘overcoat’ of SiC around the fiber tows (produced by CVI) and a SiC-based two-phase matrix (produced by reactive MI). The matrix is essentially fully dense, without the large pores that are invariably present in CFCCs made by CVI alone. The absence of pores in the matrix has two consequences. First, it increases the stress at the onset of matrix cracking relative to that of the CVI CFCCs containing pores. Secondly, it is expected to improve the thermal conductivity, especially in the direction transverse to the fibers. The fiber volume fraction is 35%. To assess the effects of stress concentrators on low cycle fatigue (LCF), three types of specimens were used: (i) standard dog-bone (unnotched) specimens; (ii) straight tensile specimens of width 2W=31.8 mm, with a center hole of diameter 2a=6.35 mm ða=W ¼ 0:2Þ; and (iii) straight tensile specimens of width 2W=31.8 mm, with a center notch of length 2a=6.35 mm ða=W ¼ 0:2Þ and notch root radius 0:2 mm. The specimens were prepared by electrodischarge machining. Four specimens of each geometry were tested at a temperature of 1
Produced by Dow Corning.
Fig. 1. Scanning electron micrographs of a cross-section through the composite: (a) at low magnification, showing the longitudinal and transverse fiber tows; (b) at higher magnifications, showing the meltinfiltrated matrix surrounding a fiber tow, and (c) the interior of the tow. The micrographs were taken in backscatter electron imaging (BEI) mode, thereby revealing atomic number contrast. The coatings on the fibers are BN. The SiC within the tows was produced by a chemical vapor infiltration process, prior to melt-infiltration of the remainder of the matrix.
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815 C in ambient air. The specimens were heated using an induction heating system and a steel susceptor. Temperatures were measured using thermocouples at three locations on the specimen surface: at the center and at points located 20 mm from the center. The temperatures at these three locations were maintained within 5 C of the targeted temperature (815 C) for the duration of each test. Steel tabs were affixed to the specimen ends using a high temperature ceramic adhesive and the tabbed sections then inserted into hydraulic wedge grips. Additionally, several tests were performed at room temperature on straight (unnotched) specimens as well as center hole specimens with hole diameters of either 2a=3.18 or 6.35 mm and with a=W ¼ 0:2. The latter tests were used to assess the extent of notch sensitivity under ambient conditions. These results formed the basis on which the notch sensitivity of LCF life was subsequently assessed. The elevated temperature tests included one monotonic tension test on each specimen type at a stressing rate of d=dt ¼ 100 MPa/min, typically leading to fracture in 43 min. The remaining three specimens were subject to LCF tests with a stress ratio of R=0.01. The loading spectrum comprised relatively rapid loading and unloading (each taking 1 min), and a hold time of 2 h at the peak stress. This spectrum was chosen to simulate the loading conditions that a gas turbine engine component might encounter in service. The peak stress levels were selected on the basis of the monotonic tensile test results. In cases in which fracture did not occur, the tests were terminated after 1 week ( 160–170 h) of loading. The room temperature tests were also performed at a stressing rate of 100 MPa/min. The monotonic tensile test results on the unnotched specimens are plotted in Fig. 2. At both 20 and 815 C, the tensile response is essentially bilinear, with the transition in slope occurring at the onset of matrix cracking
Fig. 2. Tensile response of unnotched specimens at both 20 and 815 C. Note the strong similarity in the stress-strain response at the two test temperatures, the key difference being the strains at which fracture occurs. The solid lines represent a bilinear fit to the data, used for calibrating the nonlinear constitutive law.
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at a stress of 150–165 MPa. There was no apparent effect of temperature on this response, except that the fracture strain at 815 C was slightly lower than that at 20 C. Optical microscopy of polished cross-sections through the fractured test specimens revealed periodic matrix cracks, as commonly observed in this class of CFCC. The effects of holes on the net-section strength at room temperature are shown in Fig. 3. The strength decreases with increasing hole size, to 70% of the unnotched strength at 2a=6.3 mm. Similar reductions have been reported previously for SiC/SiC, SiC/MAS [5] and all-oxide ceramic composites [35]. These effects can be rationalized on the basis of the stress redistribution due to inelastic straining around the hole as well as the size-scale dependence of the local conditions needed to precipitate fracture, as described in Section 4. The LCF behavior of 815 C is shown in Fig. 4. Several notable features emerge. (i) The monotonic tension tests on the center-hole and center-notch geometries yield strengths similar to one another, c 160–165 MPa,
Fig. 3. Effects of center holes on room-temperature tensile strength. The solid line represents the prediction of the point stress fracture model for a characteristic distance of d=0.75 mm.
Fig. 4. Summary of LCF life at 815 C. Arrows pointing to the right indicate runout; arrows pointing to the left indicate test results obtained from monotonic tension tests.
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despite large differences in the elastic stress concentration factors: ke=2.5 vs. 7.1 (see Appendix). These strengths are 60% of the unnotched monotonic strength, u 270 MPa, measured at 815 C. Evidently the degree of notch sensitivity at this temperature is only slightly greater than that at 25 C. (ii) In the unnotched specimens, a fatigue threshold (defined by tf 5 160h) was obtained at a stress O level, th 165 MPa, corresponding closely to the matrix cracking limit: mc 150–160 MPa (Fig. 2). Similar correlations between the threshold stress and the matrix cracking stress had been established previously for static loading of a CVI SiC/SiC at a temperature of 900 C [31]. However, the cracking stress of the latter material was considerably lower (60–70 MPa), a consequence of the matrix porosity. (iii) In the center-hole and center-notch specimens, the threshold was considerably lower, th 60–85 MPa, representing 40–50% of the unnotched threshold. Essentially the same behavior was obtained with notches and holes. (iv) For all three specimen geometries, the stress rupture curves were extremely shallow, with the exception of the regime at very short fracture times (tf << 1 h). Evidently fracture occurs very rapidly at stresses above the threshold level. The implication is that life prediction at stresses above the threshold is likely to be exceedingly difficult.
3. Fractography Representative fractured test specimens were examined in a scanning electron microscope (SEM). Typical features are shown in Figs. 5 and 6. Specimens subjected to monotonic tensile loading at elevated temperature exhibited little fiber pullout (Fig. 5), suggesting a relatively high interfacial strength. However, there was no indication of oxidation on the fracture surfaces following the short exposure times ( 3 min) of these test specimens. Similarly low levels of pullout were observed on the fracture surfaces of specimens tested at room temperature. By contrast, the LCF specimens exposed to elevated temperature for many hours exhibited significant levels of oxidation on the fracture plane. Fig. 6 shows micrographs of a center-hole specimen, tested at a peak stress, p ¼85 MPa. The regions of the fracture surface near the hole edge exhibited particularly severe oxidation, as manifest by the presence of a glassy layer on the axial fiber tows. Similar features were also observed on the center-notch specimens. The inference from the presence of the glassy layer on the fiber fracture surfaces is that the fibers are failing progressively, starting near the tip of the notch or hole (where the stress concentration is at a maximum) and proceeding through the remaining (unnotched) section of the material. This sequence is consistent with that observed in other SiC-based CFCCs, as detailed in [15].
Fig. 5. Fracture surface of a center hole specimen tested under monotonic tension at 815 C.
4. Modeling of notch sensitivity In light of the extremely shallow stress rupture curves, the modelling activity focused predominantly on the effects of notches and holes on the threshold stress levels. A secondary priority was the notch sensitivity of the monotonic tensile strength at ambient temperature. 4.1. Notch sensitivity of monotonic tensile strength The notch sensitivity of strength under ambient (nonoxidizing) conditions is dictated by two factors: (i) the extent to which inelastic straining mitigates the stress concentration at the notch tip, and (ii) the dependence of the fracture condition on the stress gradients that exist ahead of the notch. The former effects can be calculated by finite element methods using a non-linear constitutive law appropriate to CFCCs. In the present study, the law developed by Genin and Hutchinson [36] has been used for such calculations. This constitutive law is based on a phenomenological description of the development of inelastic strain in cross-ply CFCCs under biaxial stressing. The pertinent functions are
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calibrated using the results of tensile tests performed both along one of the fiber axes (in the 0 /90 orientation) and at 45 to the fiber axes. The constitutive law is then implemented in a finite-element code for the purpose of calculating stress and strain fields in non-uniform specimen or component geometries. The capability of the model to predict the latter response has been validated through comparisons with strain measurements on specimens containing notches and circular holes [5,35]. The model tacitly assumes that the material response is scale-independent. This presents some limitations on the size-scale of features that can be accurately simulated, as discussed in Section 4.2. Additional details of the implementation and calibration procedures are described in Refs. [36] and [5]. The effects of stress gradients on fracture are less well understood. Nevertheless, a simple fracture criterion based on the attainment of a critical stress (taken as the unnotched strength) over a characteristic distance d along the incipient fracture plane has been found to adequately describe notched strength of CFCCs, provided the effects of inelastic straining on the stress distribution are taken
into account [5,35]. Furthermore, previous studies have revealed that the characteristic distance inferred from the model and the experiment is rather insensitive to the details of the composite microstructure, falling in the narrow range of 0.5–0.75 mm for SiC/SiC [5], SiC/ MAS [5] and an all-oxide CFCC [35]. This approach has been used to model the strength of the center-hole specimens of the SylramicTM/SiC composite tested at room temperature, using a characteristic distance, d=0.75 mm. The effects of inelastic straining on the stress distribution ahead of a hole are illustrated by the results plotted in Fig. 7(a). The strength predictions are plotted in Fig. 3, for comparison with the experimental measurements. Excellent agreement between experiment and theory is obtained. Furthermore, the inferred characteristic distance is remarkably similar to the values obtained in other CFCCs, indicating a commonalty in the factors controlling fracture of the various composite systems. The same approach has been used to rationalize the effects of both notches and holes on the strength at 815 C. In this case, the characteristic distance was inferred
Fig. 6. Fracture surface of a center hole specimen tested in fatigue at p ¼ 85 MPa at 815 C.
Fig. 7. Typical stress distributions for (a) center-hole and (b) centernotched specimens of the SylramicTM/SiC composite, calculated using the nonlinear constitutive law of Genin and Hutchinson [36]. Also shown for comparison are the corresponding elastic predictions. The results demonstrate the role of inelastic straining in mitigating stress concentrations.
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Fig. 8. Comparisons of the measured and predicted monotonic tensile strengths of specimens with holes and notches at 815 C. The strength is underestimated using the elastic stress concentration factor in combination with a critical stress fracture criterion, i.e. n ¼ u =ke . In contrast, the model based on the inelastic stress distribution coupled with the point-stress criterion provides an adequate representation, for d 0.3–0.5 mm.
from comparisons of the measured and predicted notched strengths, assuming various values for d in the model. The comparisons are presented in Fig. 8. The inferred characteristic distance is d 0.3–0.5 mm, considerably smaller than the value obtained from the room temperature tests (d 0.75 mm). This reduction in d reflects a degradation in notch-sensitivity. It is unknown whether this degradation is intrinsic to the fracture mechanism at elevated temperature, it reflects the onset of oxidative embrittlement, or it is associated with scatter in the experimental data. 4.2. Notch sensitivity of the threshold stress An analogous stress-based fracture model has been used to rationalize the threshold stress levels in the center-notch and center-hole specimens. The threshold stress was taken to be that for which the longitudinal stress, , along the incipient fracture plane exceeds the o unnotched threshold, th 165 MPa, over a characteristic distance, dth, ahead of the hole or notch. As before, the characteristic distance was obtained by comparing the calculated results with the experimental measurements. Such comparisons are plotted in Fig. 9. For the center-hole specimens, the inferred distance spans the range dth 00:2 mm: the variation arising from the difference between the lowest stress at which fracture had occurred, 85 MPa, and the highest stress at which run-out was obtained, 60 MPa. Such low values of dth are indicative of extremely strong notch sensitivity. Indeed, in the limit of dth=0, the threshold corresponds to the case where the maximum stress (at the hole edge) is equal to the unnotched threshold stress. Since the unnotched threshold is dictated by the matrix cracking stress and the material response up to the cracking limit is elastic, a conservative estimate of the threshold can be obtained using the elastic stress concentration factor ke
Fig. 9. Comparisons of measured and predicted thresholds at 815 C for (a) center-hole and (b) center-notch specimens, based on the nonlinear finite element model and the point-stress fracture criterion.
in combination with the matrix cracking stress, whereupon the predicted threshold becomes o th ¼ th =ke ¼ mc =ke
ð1Þ
This represents a lower-bound estimate of the threshold stress. For the center-hole specimen with a=W ¼ 0:2, ke=2.5 [37] (Appendix) and thus th 66 MPa [also plotted on Fig. 9(a)]. This prediction is in reasonable agreement with the range obtained from the experiments (60–85 MPa), reaffirming the poor notch sensitivity associated with LCF fracture at the test temperature. Furthermore, the extreme sensitivity of the fracture time to the presence of matrix cracks is consistent with the oxidation embrittlement characteristics of other SiC-based CFCCs at intermediate temperatures [12,15,17,22,23,27,38]. For the center-notch specimens, the inferred characteristic distance is somewhat larger, dth 0:30:4 mm, indicating improved damage tolerance and lower notch sensitivity. Furthermore, the lower bound estimate of the threshold stress, th 23 MPa, based on
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Eq. (1) and the elastic stress concentration factor, ke=7.1 [39], is significantly below the measured value, 70–85 MPa. This difference is believed to be associated with the steep stress gradients that exist in the notch tip region coupled with the intrinsic volume dependence of the matrix strength. That is, the high stress levels exist over only small volumes ahead of the notch (over distances ), reducing significantly the probability of encountering a matrix flaw to initiate cracking. Finite element calculations for the center-notched geometry used in the present study (with 2a=6.35 mm and =0.2 mm) reveal that the elastic stress diminishes rapidly with distance from the notch tip, by 50% at a distance of 0.1 mm [Fig. 7(b)]. (By contrast, the reduction in elastic stress in the center-hole specimen with a hole diameter of 6.35 mm is only 6% at the same distance from the hole edge.) It is surmised that even a modest volume dependence of strength might lead to a significant elevation in the maximum local stress at the onset of cracking. This elevation in the local cracking stress may account for the rather high threshold of the notched specimens. If so, the beneficial effects of inelastic straining on the stress concentration may be overestimated by the non-linear constitutive law used in this study, because, in essence, the volume dependence of the cracking stress is neglected. A further implication is that the inferred characteristic distance (dth 0.3–0.4 mm) may be anomalously high. It remains to be established which of these effects is dominant.
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times in the post-cracking regime are not likely to be useful, because of the precipitous drop in fracture time to unacceptably low levels. The LCF threshold for center-hole specimens is similarly dictated by the matrix cracking limit. Once the cracking stress is exceeded locally (even over a very small distance ahead of the hole), fracture ensues. A reasonably accurate estimate of the threshold is obtained through the elastic stress concentration factor in combination with the matrix cracking stress [Eqn. (1)]. This highly notch-sensitive behavior suggests the need for an extremely conservative design approach for CFCC structures containing holes. Furthermore, it precludes the exploitation of the inelastic deformation mechanisms that operate in these materials and impart the requisite damage tolerance in non-oxidizing environments. The LCF threshold for center-notch specimens is comparable to that of the center hole specimens. Recognizing that the stress concentration factor is higher in the notched geometry, the composite appears to exhibit superior damage tolerance in the notched geometry. However, this apparent damage tolerance may be a result of the volumedependence of the matrix cracking stress, not a consequence of local inelastic straining. Further experiments are needed to determine the statistical parameters associated with the matrix strength and to assess their role in the conditions for fracture ahead of a notch. Additional experiments are also needed to assess the extent of scatter in fatigue life, for the purpose of validating the failure models presented here.
5. Summary The notch sensitivity of the tensile strength of the SylramicTM/SiC composite at ambient temperature is consistent with the trends observed in other CFCCs. Typically a reduction of 25% in net-section strength is obtained for open holes of diameter 6 mm. This modest amount of notch sensitivity can be rationalized on the basis of the stress redistribution that occurs as a result of local inelastic straining and the size-scale dependence of the conditions at the onset of fracture through the point stress fracture criterion. The notch sensitivity at elevated temperature appears to be reduced slightly, as manifest in a reduction in the inferred characteristic distance in the point stress criterion. The unnotched LCF threshold for the SylramicTM/SiC composite at 815 C is dictated by the matrix cracking stress. Above the cracking stress, fracture occurs rapidly, typically in 4100 h. Moreover, the LCF life curve for stresses slightly above the threshold is extremely shallow, indicating a very strong sensitivity of fracture time to applied stress. For the lifetime requirements of most gas turbine engine components (103–104 h), the allowable stress levels would have to remain below this threshold. A further implication is that models to predict fracture
Appendix. Elastic stress concentrations in notched tensile specimens The elastic stress concentration factors, ke ; were calculated assuming that the composite is elastically isotropic. This assumption can be justified on the basis that the Young’s moduli measured in the 0 /90 and 45 orientations were within 2% of one another (250 GPa). This behavior is attributable to the fact that both the matrix and the fibers are essentially pure SiC. For a specimen of width 2W and with a center hole of diameter 2a, ke is given by [37]: ke
o a 3 ¼2þ 1 W net
ðA1Þ
where o is the maximum longitudinal stress (at the hole edge) and net is the applied net-section stress. For the center hole geometry used in the present experiments, a=W ¼ 0:2 and thus ke =2.51. The elastic stress concentration factor for a specimen of width 2W and with a center notch of length 2a and root radius (with a= 1) is given by [39]:
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ke
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o a ¼ 2 1 W net
rffiffiffiffiffi a
a1=2 sec W
ðA2Þ
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