Availableonlineatwww.sciencedirect.com BCIENCEODIRECT@ COMPOSITES SCIENCE AND TECHNOLOGY ELSEVIER Composites Science and Technology 64(2004)1311-1319 Stress-dependent matrix cracking in 2D woven SiC-fiber reinforced melt-infiltrated SiC matrix composites Ohio Aerospace Institute, Brookpark, OH, US.A Received 24 February 2003: received in revised form 23 October 2003: accepted 23 October 2003 Available online 23 December 2003 Abstract The matrix cracking of a variety of SiC/SiC composites has been characterized for a wide range of constituent variation. These composites were fabricated by the two-dimensional lay-up of 0/90 five-harness satin fabric consisting of Sylramic fiber tows that were then chemical vapor infiltrated(CVn) with BN, CVI with Sic, slurry infiltrated with Sic particles followed by molten infil- tration of Si. The composites varied in number of plies, the number of tows per length, thickness, and the effective-size of the tows. his resulted in composites with a fiber volume fraction in the load-bearing direction that ranged from 0. 12 to 0. 20. Matrix cracking was monitored with modal acoustic emission in order to estimate the stress-dependent distribution of matrix cracks. It was found that the general matrix crack properties of this system could be fairly well characterized by assuming that no matrix cracks orig- inated in the load-bearing fiber, interphase, chemical vapor infiltrated SiC tow-minicomposites, i.e., all matrix cracks originate in the 90 tow regions or the large unreinforced SiC-Si matrix regions. Also, it was determined that the higher fiber-count tow composites had a much narrower stress range for matrix cracking compared to the standard tow size composites. c 2003 Elsevier Ltd. All Keywords: A Ceramic matrix composites: Stress-strain behavior: B Matrix cracking: D. Acoustic 1. Introduction since matrix cracking results in the desired stress-strain non-linearity, composite toughness, and strength prop- Sic-fiber reinforced, melt-infiltrated(MI) SiC matrix erties [3]. For non-oxide composites, such as the Sic/Sic omposites are leading candidate materials for aircraft system, the presence of matrix cracks enables oxidizing and land-based turbine engine applications such as a environments to diffuse into the interior of the composite combustor liner [1, 2]. However, for such materials to be and cause strength-degradation, especially at intermedi sed, the stress-strain behavior of these materials needs ate temperatures [4, 5]. In addition, for some BN inter to be well characterized for composites with a wide range phase composites, the degree of strength-degradation at of physical characteristics, e.g., thickness, fiber archi- intermediate temperatures is related to the number of tecture, fiber volume fraction, etc. since composite matrix cracks [6]. Therefore, it is essential that a good structures are not necessarily simple shapes but consist of understanding of the stress-dependent matrix crack thickness changes, curvature, and attachment schemes properties of a viable composite system be well-charac depending on the need of the component. In order to terized. For example, this has been done to a large extent nodel the stress-response of these materials a good un- for the Nicalon,C infil- derstanding of their matrix crack properties are required trated(CVI) Sic matrix system [7-101 a considerable amount of has occurred for the Sic fiber-reinforced. bn inter- NASA Glenn Research Center, Ohio Space phase, MI matrix system [11-13]. This development has MS-106-5. Cleveland, OH 44135 USA. Tel: included studying different fiber-types, interphases, and +1216-433-5544. gory. n. morscher(@grc. nasa. gov (G.N. Mor- Nippon Carbon Co., Tokyo, Japa 0266-3538/S- see front matter 2003 Elsevier Ltd. All rights reserved. doi: 10. 1016/j. compscitech 2003 10.02
Stress-dependent matrix cracking in 2D woven SiC-fiber reinforced melt-infiltrated SiC matrix composites Gregory N. Morscher * Ohio Aerospace Institute, Brookpark, OH, USA Received 24 February 2003; received in revised form 23 October 2003; accepted 23 October 2003 Available online 23 December 2003 Abstract The matrix cracking of a variety of SiC/SiC composites has been characterized for a wide range of constituent variation. These composites were fabricated by the two-dimensional lay-up of 0/90 five-harness satin fabric consisting of Sylramic fiber tows that were then chemical vapor infiltrated (CVI) with BN, CVI with SiC, slurry infiltrated with SiC particles followed by molten infiltration of Si. The composites varied in number of plies, the number of tows per length, thickness, and the effective-size of the tows. This resulted in composites with a fiber volume fraction in the load-bearing direction that ranged from 0.12 to 0.20. Matrix cracking was monitored with modal acoustic emission in order to estimate the stress-dependent distribution of matrix cracks. It was found that the general matrix crack properties of this system could be fairly well characterized by assuming that no matrix cracks originated in the load-bearing fiber, interphase, chemical vapor infiltrated SiC tow-minicomposites, i.e., all matrix cracks originate in the 90 tow regions or the large unreinforced SiC–Si matrix regions. Also, it was determined that the higher fiber-count tow composites had a much narrower stress range for matrix cracking compared to the standard tow size composites. 2003 Elsevier Ltd. All rights reserved. Keywords: A. Ceramic matrix composites; Stress–strain behavior; B. Matrix cracking; D. Acoustic emission 1. Introduction SiC-fiber reinforced, melt-infiltrated (MI) SiC matrix composites are leading candidate materials for aircraft and land-based turbine engine applications such as a combustor liner [1,2]. However, for such materials to be used, the stress–strain behavior of these materials needs to be well characterized for composites with a wide range of physical characteristics, e.g., thickness, fiber architecture, fiber volume fraction, etc. since composite structures are not necessarily simple shapes but consist of thickness changes, curvature, and attachment schemes depending on the need of the component. In order to model the stress-response of these materials a good understanding of their matrix crack properties are required since matrix cracking results in the desired stress–strain non-linearity, composite toughness, and strength properties [3]. For non-oxide composites, such as the SiC/SiC system, the presence of matrix cracks enables oxidizing environments to diffuse into the interior of the composite and cause strength-degradation, especially at intermediate temperatures [4,5]. In addition, for some BN interphase composites, the degree of strength-degradation at intermediate temperatures is related to the number of matrix cracks [6]. Therefore, it is essential that a good understanding of the stress-dependent matrix crack properties of a viable composite system be well-characterized. For example, this has been done to a large extent for the NicalonTM, 1 C interphase, chemical vapor infiltrated (CVI) SiC matrix system [7–10]. A considerable amount of composite development has occurred for the SiC fiber-reinforced, BN interphase, MI matrix system [11–13]. This development has included studying different fiber-types, interphases, and * Present address: NASA Glenn Research Center, Ohio Space Institute, Lewis Field, MS-106-5, Cleveland, OH 44135, USA. Tel.: +1-216-433-5512; fax: +1-216-433-5544. E-mail address: gregory.n.morscher@grc.nasa.gov (G.N. Morscher). 1 Nippon Carbon Co., Tokyo, Japan. 0266-3538/$ - see front matter 2003 Elsevier Ltd. All rights reserved. doi:10.1016/j.compscitech.2003.10.022 Composites Science and Technology 64 (2004) 1311–1319 www.elsevier.com/locate/compscitech COMPOSITES SCIENCE AND TECHNOLOGY
G N. Morscher Composites Science and Technology 64 (2004)1311-1319 matrix compositions in order to maximize composite properties. Most of the development composite panels E荡 ave been processed with the sylramic- or the newer Sylramic-iBN fiber-type. The latter offers the strongest most creep-resistant, and most reliable composites to date. A variety of composite specimens were tested in this study from those developmental panels with a wide 语35 variation of the 2D woven, five-harness satin(5HS)ar- 寸R≥9g85g chitecture, e.g., changes in composite thickness, number of plies, number of tows per length, and the number of fibers per woven tow. An earlier work [14] compared the effects of these changes on the stress-strain curve in general. This study will concentrate on characterizing 甚ssss the effect these changes have on matrix cracking toward the development of a general relationship that can be used for purposes of component design and perfor- nance modeling 2. Experimental 6,55864,日 Unload-reload tensile hysteresis tests were performed R商送因后后 on ten different composite specimens, each from a dif ferent SiC/SiC composite panel, that varied in fiber tow ends per unit length, number of plies, and composite thickness. Two different fiber-types were used, Sylramic 上过过 Syl) and in situ-BN Sylramic(Syl-iBN) which is a modified Sylramic fiber that has been heat-treated [13] 兰 to form an in situ BN layer (iBN)on the fiber surface prior to composite fabrication. However, the differences 338 in the fiber-types are considered to not affect matrix cracking behavior. Both have 800 fibers per standard 当限喜 tow with an average fiber diameter of 10 um. For one composite panel, 041, two standard 800 fiber-count tows were woven together at 3.9 tow ends per cm(epcm)[15] In effect this was like having a woven tow with twice as many fibers, i.e., 1600, as the standard tow. However, 041 had the same fraction of fibers in the loading 号期可 direction as a single-tow weave of 7.9 epcm(011 in Table 1). Optical micrographs of polished longitudinal sections of the oll and the 041 composites are shown in ∞∞∞g必∞|s Fig. 1. It is evident that the double-tow woven com- posite essentially doubles the width dimension of the effective woven tow. The composites were processed by 5≥ the former Honeywell Advanced Composites(Newark, DE), currently known as General Electric Power Sys- [2]. Table 1 lists the consti ations for the composites tested. Note that there was considerable variation in fiber volume fraction and cvi 百 SiC volume fraction between the panels tested. Composite processing entails first stacking of bal anced 0/90 five-harness fabric woven from SYL or SYL iBN tows, a Bn interphase layer deposition (0.5 um) via CVi, a Sic interphase over-coating via CVI, Sic 自 3旨象 2 Dow Corning Corporation, Midland, MI
matrix compositions in order to maximize composite properties. Most of the development composite panels have been processed with the Sylramic 2 or the newer Sylramic-iBN fiber-type. The latter offers the strongest, most creep-resistant, and most reliable composites to date. A variety of composite specimens were tested in this study from those developmental panels with a wide variation of the 2D woven, five-harness satin (5HS) architecture, e.g., changes in composite thickness, number of plies, number of tows per length, and the number of fibers per woven tow. An earlier work [14] compared the effects of these changes on the stress–strain curve in general. This study will concentrate on characterizing the effect these changes have on matrix cracking towards the development of a general relationship that can be used for purposes of component design and performance modeling. 2. Experimental Unload–reload tensile hysteresis tests were performed on ten different composite specimens, each from a different SiC/SiC composite panel, that varied in fiber tow ends per unit length, number of plies, and composite thickness. Two different fiber-types were used, Sylramic (Syl) and in situ-BN Sylramic (Syl-iBN) which is a modified Sylramic fiber that has been heat-treated [13] to form an in situ BN layer (iBN) on the fiber surface prior to composite fabrication. However, the differences in the fiber-types are considered to not affect matrixcracking behavior. Both have 800 fibers per standard tow with an average fiber diameter of 10 lm. For one composite panel, 041, two standard 800 fiber-count tows were woven together at 3.9 tow ends per cm (epcm) [15]. In effect this was like having a woven tow with twice as many fibers, i.e., 1600, as the standard tow. However, 041 had the same fraction of fibers in the loading direction as a single-tow weave of 7.9 epcm (011 in Table 1). Optical micrographs of polished longitudinal sections of the 011 and the 041 composites are shown in Fig. 1. It is evident that the double-tow woven composite essentially doubles the width dimension of the effective woven tow. The composites were processed by the former Honeywell Advanced Composites (Newark, DE), currently known as General Electric Power Systems Composites [2]. Table 1 lists the constituent variations for the composites tested. Note that there was considerable variation in fiber volume fraction and CVI SiC volume fraction between the panels tested. Composite processing entails first stacking of balanced 0/90 five-harness fabric woven from SYL or SYLiBN tows, a BN interphase layer deposition (0.5 lm) via CVI, a SiC interphase over-coating via CVI, SiC Table 1 Properties of tested SiC/SiC specimens Specimen (type of debondinga) Sylramic fiber typeb Tow ends per cm No. of plies Specimen thickness (mm) Crack density after specimen failure, #/mm Fiber fractionc E (GPa) rth (MPa) BN fractionc CVI SiC fractionc fmini Emini Single-tow woven composites 007 (OD) AP 4.9 8 1.63 8.0 0.15 219 )35 0.04 0.11 0.30 357 016 (ID) iBN 4.9 8 2.04 11.6 0.12 279 )35 0.03 0.23 0.38 386 017 (ID) iBN 4.9 8 1.46 12.0 0.17 224 )53 0.04 0.14 0.35 366 012 (ID) iBN 7.1 6 1.99 7.2 0.14 289 )37 0.03 0.20 0.36 379 009 (MD) AP 7.9 8 2.18 9.0 0.18 246 )55 0.04 0.11 0.33 358 011 (MD) iBN 7.9 8 2.04 10.3 0.19 228 )57 0.04 0.12 0.34 361 018 (ID) iBN 7.9 4 1.37 5.0 0.14 235 )53 0.03 0.12 0.29 364 044 (OD) iBN 8.7 8 2.14 9.5 0.20 216 )35 0.03 0.13 0.35 369 068 (ID) IBN 8.7 8 2.21 10.4 0.20 277 )67 0.04 0.15 0.39 366 Double-tow woven composite 041 (OD) IBN 3.9(2)d 8 2.07 9.0 0.19 197 )50 0.03 0.11 0.33 363 a ID, inside debonding; OD, outside debonding; MD, mixed debonding. b AP, as-produced; iBN, in situ BN. Each fiber tow consisted of 800 fibers. The average fiber diameter was 10 lm. c Volume fraction of constituent in the load-bearing direction, the total fraction would be double this amount. dTwo tows were woven together into a 3.9 epcm fabric. In effect there were the same number of 800 count tows per length as the 7.9 epcm 011 specimen; however, the effective tow size for 041 was 1600 fibers. 2 Dow Corning Corporation, Midland, MI. 1312 G.N. Morscher / Composites Science and Technology 64 (2004) 1311–1319
G N. Morscher Composites Science and Technology 64(2004)1311-1319 at 500 w for 30 min. The etchant reacts with the free si in the matrix, removing much of it, making it impossible to observe cracks in the Mi part of the matrix. Matrix cracks can only be observed in the dense CVi SiC layer between the bn and the mi matrix 3. Results and analysis 3. 1. Standard single-tow woven composites Monotonic and unload-reload stress strain data with E activity plotted as energy are shown in Fig. 2 for two different specimens from the same panel. Several aspects of Fig. 2 are characteristic of the Sylramic/BN/MI SiC system For specimens from the same panel of material, the stress-strain properties are very consistent, i.e., little scatter from specimen to specimen and little difference 器m for monotonic and unload-reload experiments. Also the ae activity is very consistent and occurs over a Fig. I Polished longitudinal sections of standard tow woven (olD) range of stress(strain). Finally, the matrix is in residual composite and double-tow woven(041)composite. compression, which is indicative of the intersection of the intercepts of the average slope of the top portion of particulate infiltration via slurry-infiltration, and finally the hysteresis loop in the positive stress-strain quadrant liquid Si infiltration [1, 2] according to Steen and Valles [18] The tensile tests were performed on 150 mm long pecimens with a contoured gage section (dog-bone 2.5mm width in grip regid idth in gage section) using a universal-testing machine (Instron Model 8562, Instron, Ltd, Canton Mass. with an elec SYL-iBN tromechanical actuator. Glass fiber reinforced epoxy 8.epcm: 8 ply: f=0.2 12 tabs were mounted on both sides of the specimen in the grip regions and the specimens were gripped with rigidly mounted hydraulically actuated wedge grips. A clip on strain gage, with a range of 2. 5% strain over 25.4 mm gage length was used to measure the deformation of the gage section. Tensile tests were performed in load- control at 2 kN/min Modal acoustic emission(AE)was monitored during the tensile tests with two wide -band. 50 kHz to 2.0 MHz 0050.10.150202503035 high fidelity sensors placed just outside the tapered re- gion of the dog-bone specimen. Vacuum grease was used as a couplant and mechanical clips were used to mount (b) 1 the sensors to the specimen. The AE waveforms were recorded and digitized using a 4-channel, Fracture Wave Detector(FWD) produced by Digital Wave Corpora tion(Englewood, CO). The load and strain were also E0.5 recorded. After the tensile test, the ae data was filtered 3 04 using the location software from the fwd manufac- 0.3 turer in order to separate out the ae that occurred outside of the gage section. For more information on the AE procedure and analysis, see [16, 17] Since residual compressive stresses in the matrix close dense the matrix cracks, to measure crack density, sections of the tested tensile specimens in the gage section at least Fig. 2. Typical (068)monotonic and load-unload 10 mm long were polished and then plasma(CF4)etched stress-strain behavior and(b)stress-dependent AE activity
particulate infiltration via slurry-infiltration, and finally, liquid Si infiltration [1,2]. The tensile tests were performed on 150 mm long specimens with a contoured gage section (dog-bone, 12.5 mm width in grip region and 10 mm width in gage section) using a universal-testing machine (Instron Model 8562, Instron, Ltd, Canton Mass.) with an electromechanical actuator. Glass fiber reinforced epoxy tabs were mounted on both sides of the specimen in the grip regions and the specimens were gripped with rigidly mounted hydraulically actuated wedge grips. A clip on strain gage, with a range of 2.5% strain over 25.4 mm gage length was used to measure the deformation of the gage section. Tensile tests were performed in loadcontrol at 2 kN/min. Modal acoustic emission (AE) was monitored during the tensile tests with two wide-band, 50 kHz to 2.0 MHz, high fidelity sensors placed just outside the tapered region of the dog-bone specimen. Vacuum grease was used as a couplant and mechanical clips were used to mount the sensors to the specimen. The AE waveforms were recorded and digitized using a 4-channel, Fracture Wave Detector (FWD) produced by Digital Wave Corporation (Englewood, CO). The load and strain were also recorded. After the tensile test, the AE data was filtered using the location software from the FWD manufacturer in order to separate out the AE that occurred outside of the gage section. For more information on the AE procedure and analysis, see [16,17]. Since residual compressive stresses in the matrix close the matrix cracks, to measure crack density, sections of the tested tensile specimens in the gage section at least 10 mm long were polished and then plasma (CF4) etched at 500 W for 30 min. The etchant reacts with the free Si in the matrix, removing much of it, making it impossible to observe cracks in the MI part of the matrix. Matrix cracks can only be observed in the dense CVI SiC layer between the BN and the MI matrix. 3. Results and analysis 3.1. Standard single-tow woven composites Monotonic and unload–reload stress strain data with AE activity plotted as energy are shown in Fig. 2 for two different specimens from the same panel. Several aspects of Fig. 2 are characteristic of the Sylramic/BN/MI SiC system. For specimens from the same panel of material, the stress–strain properties are very consistent, i.e., little scatter from specimen to specimen and little difference for monotonic and unload–reload experiments. Also, the AE activity is very consistent and occurs over a range of stress (strain). Finally, the matrix is in residual compression, which is indicative of the intersection of the intercepts of the average slope of the top portion of the hysteresis loop in the positive stress–strain quadrant, according to Steen and Valles [18]. Fig. 1. Polished longitudinal sections of standard tow woven (011) composite and double-tow woven (041) composite. Fig. 2. Typical (068) monotonic and load–unload–reload hysteresis (a) stress–strain behavior and (b) stress-dependent AE activity. G.N. Morscher / Composites Science and Technology 64 (2004) 1311–1319 1313
G N. Morscher Composites Science and Technology 64 (2004)1311-1319 500 0+1 3.9(epcm; f 3501060:8pcm:t= 250 .epcm: f= 6:4epcm;f=0.12 50 0.5 0.6 0.7 Strain. Fig. 3. Stress-strain curves for the specimens tested and analyzed in his study. Note, unload-reload hysteresis loops have been removed for clarity Fig 4. Example of matrix cracks observed along a polished and etched longitudinal section of the tensile specimen(017) after failure at room temperatur Fig. 3 shows stress-strain curves. unload-reload loops removed, for a variety of different architecture composites,i.e, different panels. As is expected, the T12 posites tend to exhibit higher 649epcm;向=0.12 action com ultimate strengths, higher stresses for the"knee"in the 04t;3.92)epcm;017:49 epcm, I stress-strain curve, and steeper secondary slopes [14] There was considerable scatter in elastic moduli. some of 88068:8.epcm: 011;7.9 epcm: fe019 which was due to the anomalies described below The matrix crack density was determined fo 007: 4.9opcm: fat 012:7.1 eocm: f=0.13 of matrix cracking in these composites for a portion a p/ specimen after tensile testing. Fig 4 shows an example oeE pecimen cut and polished from the gage section fol lowed by a plasma-etch. No matrix cracks are visible, because of the residual compression in the matrix and 00050.10.15020.25 0.35 the higher interfacial shear stresses of composites with Strain. nic fibers, without plasma etching The energy of aE has shown to be a good measure of the relative crack density when the more accurate 017:4.9epcm0.17 "modal"AE approach is used for these types of com- 8.epcm, fa posite systems[17]. In other words, the relative amount 2 041:3.92pcm of cumulative AE energy is nearly directly related to the g relative number of matrix cracks formed. Therefore the 007:49epmo final matrix crack density measured from the composite test specimens was multiplied by the normalized cumu- lative AE energy(e.g, Fig. 2(b))for each specimen in 018:79epcm台=0.14 order to estimate the stress-dependent matrix crack distribution. the estimated crack distributions are shown in Fig. 5 for a number of specimens versus strain and stress. The strain and stress distribution for matrix 150200250300350 cracking varies from specimen to specimen considerably. ter to define the earliest formation of Fig. 5. Estimated matrix cracking(normalized AE measured during large matrix cracks is theonset"strain or stress at strain and (b)stress for standard single-tow woven composites and a which the rate of AE rapidly increases. The sudden double-tow woven composite(041). crease in AE activity is due to high-energy events that are associated with large bridged matrix cracks that with another matrix crack) of the specimen [16, 17] propagate through-the-thickness(or at least a significant Fig. 2(a) and(b) show the determination of Eonset and portion of the cross-section if a matrix crack links up onset from extrapolation of the initial high-rate AE
Fig. 3 shows stress–strain curves, unload–reload loops removed, for a variety of different architecture composites, i.e., different panels. As is expected, the higher volume fraction composites tend to exhibit higher ultimate strengths, higher stresses for the ‘‘knee’’ in the stress–strain curve, and steeper secondary slopes [14]. There was considerable scatter in elastic moduli, some of which was due to the anomalies described below. The matrix crack density was determined for each specimen after tensile testing. Fig. 4 shows an example of matrix cracking in these composites for a portion of a specimen cut and polished from the gage section followed by a plasma-etch. No matrix cracks are visible, because of the residual compression in the matrix and the higher interfacial shear stresses of composites with Sylramic fibers, without plasma etching. The energy of AE has shown to be a good measure of the relative crack density when the more accurate ‘‘modal’’ AE approach is used for these types of composite systems [17]. In other words, the relative amount of cumulative AE energy is nearly directly related to the relative number of matrix cracks formed. Therefore, the final matrix crack density measured from the composite test specimens was multiplied by the normalized cumulative AE energy (e.g., Fig. 2(b)) for each specimen in order to estimate the stress-dependent matrix crack distribution. The estimated crack distributions are shown in Fig. 5 for a number of specimens versus strain and stress. The strain and stress distribution for matrix cracking varies from specimen to specimen considerably. One useful parameter to define the earliest formation of large matrix cracks is the ‘‘onset’’ strain or stress at which the rate of AE rapidly increases. The sudden increase in AE activity is due to high-energy events that are associated with large bridged matrix cracks that propagate through-the-thickness (or at least a significant portion of the cross-section if a matrix crack links up with another matrix crack) of the specimen [16,17]. Fig. 2(a) and (b) show the determination of eonset and ronset from extrapolation of the initial high-rate AE Fig. 3. Stress–strain curves for the specimens tested and analyzed in this study. Note, unload–reload hysteresis loops have been removed for clarity. Fig. 4. Example of matrix cracks observed along a polished and etched longitudinal section of the tensile specimen (017) after failure at room temperature. Fig. 5. Estimated matrix cracking (normalized AE measured during the stress–strain test multiplied by measured crack density) versus (a) strain and (b) stress for standard single-tow woven composites and a double-tow woven composite (041). 1314 G.N. Morscher / Composites Science and Technology 64 (2004) 1311–1319
G N. Morscher Composites Science and Technology 64(2004)1311-1319 0.12 01· onset stra typically had lower elastic modulus values, M220+ we 0.1 Onset stress 口·g044+200 20 MPa, compared to the 250+ 30 MPa measured for ID composites. Note that the two composites with the higher stress-distributions for matrix crack ing were OD(044)and MD(Ol1) specimens (2)The matrix crack density varied from 5 to 12 cracks/ mm and did not appear to show any correspondence with the fiber volume fraction or interfacial shear strength(Table 1) 0.120.140.16 (3)Many of the composites did not fully saturate with matrix cracks. Specifically, those composites with Fig. 6. Eonset and onset versus fraction of fibers in the loading direction lower fiber volume fractions where the rate of ae activity remained high until failure(e. g, 007, 012, activity to the abscissa for strain and stress, respectively 016, and 018 in Fig. 5). Matrix crack saturation oc- The initial low energy ae prior to the AE energy in curs when the rate of cumulative Ae energy dimin crease corresponds to the formation of microcracks or ishes to near zero(slope of Fig. 2(b) plateaus at tunnel cracks that form in the 90 bundles but do not higher stress). This is evident in Fig. 5 for 00 penetrate, at least not very much, into the load-bearing 011, 068, 044, and probably 017 fibers. Eonset and onset was determined for all the speci- men data in Fig. 5. In general, Eonset and onset increase 3. 2. Normalizing matrix cracking behavior for standard with fiber volume fraction in the loading direction single-tow woren composites (Fig. 6). However, composites with the same fo but higher E have lower Eonset. Two examples are shown in It is evident that there is some relationship between Fig. 6 for specimens with a fo=0. 2(068 and 044)and fiber volume fraction and stress-distribution for matrix specimens with a fo=0.14(012 and 018). For both cracking and it would be useful to characterize matrix examples, onset was very similar but Eonset was less for cracking based on the constituent properties of the the higher E material. The stress-dependent crack den- composites and the matrix in particular. Matrix cracks sity distribution in general follows the same trend as originate within the 90 tow-minicomposites or large, Conset and occurs over a higher stress range for higher unreinforced matrix regions and then propagate fiber volume fraction through the load-bearing minicomposites [7, 8, 20). In Even though the composite panels were fabricated by other words, matrix cracks do not originate for these the same vendor over a relatively short period of time, l materials in the load-bearing fiber, interphase, CVI SIC year, there were several anomalies observed for the tow-minicomposite. If the volume fraction and elastic composites tested in this study modulus of an average""minicomposite"can be deter (1) Two types of interfacial debonding and sliding be- mined from the processing data, then the average stress haviors were present in the data set for this study in the matrix region excluding the bn and CVI SiC in and noted in Table 1. Most of the specimens exhib- the load-bearing minicomposite could be backed out ited debonding between the fiber and the BN-inter- from a rule-of-mixtures approach and a relationshi phase, this was referred to as""inside debonding between matrix cracking and matrix stress can be (ID). Two specimens exhibited debonding between established the BN-interphase and the Cvi SiC portion of the The fraction of load-bearing minicomposites, mini matrix, this was referred to"outside debonding was estimated from half of the combined fraction of fi (OD). Also, two specimens exhibited a mixture of in- ber, BN, and CVI SiC determined from the processing side and outside debonding, i. e, "mixed debonding" data sheet supplied by the composite fabricator. The (MD). This was pointed out in[14] and has been de- elastic modulus of the minicomposites, Emini, was esti scribed in greater detail in a recent paper [19]. od mated via the rule-of-mixtures from the elastic moduli of composites have much lower interfacial shear stres- each constituent of the minicomposite (Er=380 GPa, ses, 10 MPa, compared to the 70+10 MPa mea- EBN=60 GPa, and ECVLSiC 425 GPa)and the frac- sured for ID composites. Also, OD composites tion of each constituent in the loading direction. Again, a rule-of-mixtures approach can be used to"back-out the stress in the"minimatrix'"surrounding the load-bearing The interfacial shear stress. t [12, 19]:(1)fiber"push-in, a direct measure of t, and (2)best fitting the stress-strain curve for t using Eqs. (3H5)(below) and assuming the matrix crack distribution from AE energy and final matrix crack The raw data from the manufacturer is base density, an indirect measure. Both methods produced consistent each processing step. The constituent fractions determined from results processing data are tabulated in Table I for each panel
activity to the abscissa for strain and stress, respectively. The initial low energy AE prior to the AE energy increase corresponds to the formation of microcracks or tunnel cracks that form in the 90 bundles but do not penetrate, at least not very much, into the load-bearing fibers. eonset and ronset was determined for all the specimen data in Fig. 5. In general, eonset and ronset increase with fiber volume fraction in the loading direction (Fig. 6). However, composites with the same f0 but higher E have lower eonset. Two examples are shown in Fig. 6 for specimens with a f0 ¼ 0:2 (068 and 044) and specimens with a f0 ¼ 0:14 (012 and 018). For both examples, ronset was very similar but eonset was less for the higher E material. The stress-dependent crack density distribution in general follows the same trend as ronset and occurs over a higher stress range for higher fiber volume fraction composites. Even though the composite panels were fabricated by the same vendor over a relatively short period of time, 1 year, there were several anomalies observed for the composites tested in this study: (1) Two types of interfacial debonding and sliding behaviors were present in the data set for this study and noted in Table 1. Most of the specimens exhibited debonding between the fiber and the BN-interphase, this was referred to as ‘‘inside debonding’’ (ID). Two specimens exhibited debonding between the BN-interphase and the CVI SiC portion of the matrix, this was referred to ‘‘outside debonding’’ (OD). Also, two specimens exhibited a mixture of inside and outside debonding, i.e., ‘‘mixed debonding’’ (MD). This was pointed out in [14] and has been described in greater detail in a recent paper [19]. OD composites have much lower interfacial shear stresses, 10 MPa, compared to the 70 10 MPa measured for ID composites. 3 Also, OD composites typically had lower elastic modulus values, 220 20 MPa, compared to the 250 30 MPa measured for ID composites. Note that the two composites with the higher stress-distributions for matrix cracking were OD (044) and MD (011) specimens. (2) The matrix crack density varied from 5 to 12 cracks/ mm and did not appear to show any correspondence with the fiber volume fraction or interfacial shear strength (Table 1). (3) Many of the composites did not fully saturate with matrix cracks. Specifically, those composites with lower fiber volume fractions where the rate of AE activity remained high until failure (e.g., 007, 012, 016, and 018 in Fig. 5). Matrix crack saturation occurs when the rate of cumulative AE energy diminishes to near zero (slope of Fig. 2(b) plateaus at higher stress). This is evident in Fig. 5 for 009, 011, 068, 044, and probably 017. 3.2. Normalizing matrix cracking behavior for standard single-tow woven composites It is evident that there is some relationship between fiber volume fraction and stress-distribution for matrix cracking and it would be useful to characterize matrix cracking based on the constituent properties of the composites and the matrix in particular. Matrix cracks originate within the 90 tow-minicomposites or large, unreinforced matrix regions and then propagate through the load-bearing minicomposites [7,8,20]. In other words, matrix cracks do not originate for these materials in the load-bearing fiber, interphase, CVI SiC ‘‘tow-minicomposite’’. If the volume fraction and elastic modulus of an average ‘‘minicomposite’’ can be determined from the processing data, then the average stress in the matrix region excluding the BN and CVI SiC in the load-bearing minicomposite could be backed out from a rule-of-mixtures approach and a relationship between matrix cracking and matrix stress can be established. The fraction of load-bearing minicomposites, fmini, was estimated from half of the combined fraction of fi- ber, BN, and CVI SiC determined from the processing data sheet supplied by the composite fabricator. 4 The elastic modulus of the minicomposites, Emini, was estimated via the rule-of-mixtures from the elastic moduli of each constituent of the minicomposite (Ef ¼ 380 GPa, EBN ¼ 60 GPa, and ECVI–SiC ¼ 425 GPa) and the fraction of each constituent in the loading direction. Again, a rule-of-mixtures approach can be used to ‘‘back-out’’ the stress in the ‘‘minimatrix’’ surrounding the load-bearing Fig. 6. eonset and ronset versus fraction of fibers in the loading direction. 3 The interfacial shear stress, s, was determined by two methods in [12,19]: (1) fiber ‘‘push-in’’, a direct measure of s, and (2) best fitting the stress–strain curve for s using Eqs. (3)–(5) (below) and assuming the matrix crack distribution from AE energy and final matrix crack density, an indirect measure. Both methods produced consistent results. 4 The raw data from the manufacturer is based on weight gains after each processing step. The constituent fractions determined from the processing data are tabulated in Table 1 for each panel. G.N. Morscher / Composites Science and Technology 64 (2004) 1311–1319 1315