alumina ceramic composite at elevated temperature in air and in steam

Composites Science and Technology 68 (2008) 2260–2266

Contents lists available at ScienceDirect

Composites Science and Technology journal homepage: www.elsevier.com/locate/compscitech

Creep behavior in interlaminar shear of NextelTM720/alumina ceramic composite at elevated temperature in air and in steamq M.B. Ruggles-Wrenn *, P.D. Laffey Department of Aeronautics and Astronautics, Air Force Institute of Technology, Wright-Patterson Air Force Base, OH 45433-7765, United States

a r t i c l e

i n f o

Article history: Received 14 March 2008 Received in revised form 4 April 2008 Accepted 9 April 2008 Available online 15 April 2008 Keywords: A. Ceramic–matrix composites (CMCs) A. Oxides B. Creep B. High-temperature properties D. Fractography

a b s t r a c t The creep behavior in interlaminar shear of an oxide–oxide ceramic matrix composite (CMC) was evaluated at 1200 °C in laboratory air and in steam using double-notch shear test specimens. The composite consists of a porous alumina matrix reinforced with laminated, woven mullite/alumina (NextelTM720) fibers, has no interface between the fiber and matrix, and relies on the porous matrix for flaw tolerance. The interlaminar shear properties were measured. The creep behavior was examined for interlaminar shear stresses in the 4–6.5 MPa range. Primary and secondary creep regimes were observed in all tests conducted in air. In steam, the composite exhibited primary, secondary and tertiary creep. In air, creep run-out defined as 100 h at creep stress was achieved in all tests. In the presence of steam, creep performance deteriorated rapidly and run-out was achieved only at 4 MPa (50% of the interlaminar shear strength at 1200 °C). The retained properties of all specimens that achieved run-out were characterized. Composite microstructure, as well as damage and failure mechanisms were investigated. Matrix degradation appears to be the cause of reduced creep lifetimes in steam. Published by Elsevier Ltd.

1. Introduction Advances in power generation systems for aircraft engines, land-based turbines, rockets, and, most recently, hypersonic missiles and flight vehicles have raised the demand for structural materials that have superior long-term mechanical properties and retained properties under high temperature, high pressure, and varying environmental factors, such as moisture [1]. Typical components include combustors, nozzles and thermal insulation. Ceramic–matrix composites (CMCs), capable of maintaining excellent strength and fracture toughness at high temperatures are prime candidate materials for such applications. Additionally, lower densities of CMCs and their higher use temperatures, together with a reduced need for cooling air, allow for improved high-temperature performance when compared to conventional nickelbased superalloys [2]. Concurrent efforts in optimization of the CMCs and in design of the combustion chamber are expected to accelerate the insertion of the CMCs into aerospace turbine engine applications, such as combustor walls [3–5]. Because these applications require exposure to oxidizing environments, the thermodynamic stability and oxidation resistance of CMCs are vital issues. The need for environmentally stable composites motivated the

q The views expressed are those of the authors and do not reflect the official policy or position of the United States Air Force, Department of Defense or the US Government. * Corresponding author. Tel.: +1 937 255 3636x4641; fax: +1 937 656 7053. E-mail address: marina.ruggles-wrenn@afit.edu (M.B. Ruggles-Wrenn).

0266-3538/$ - see front matter Published by Elsevier Ltd. doi:10.1016/j.compscitech.2008.04.009

development of CMCs based on environmentally stable oxide constituents [6–11]. The main advantage of CMCs over monolithic ceramics is their superior toughness, tolerance to the presence of cracks and defects, and non-catastrophic mode of failure. It is widely accepted that in order to avoid brittle fracture behavior in CMCs and improve the damage tolerance, a weak fiber/matrix interface is needed, which serves to deflect matrix cracks and to allow subsequent fiber pullout [12–14]. It has been demonstrated that similar crack-deflecting behavior can also be achieved by means of a finely distributed porosity in the matrix instead of a separate interface between matrix and fibers [15]. This microstructural design philosophy implicitly accepts the strong fiber/matrix interface. The concept has been successfully demonstrated for oxide–oxide composites [6,9,11,16,17]. Resulting oxide/oxide CMCs exhibit damage tolerance combined with inherent oxidation resistance. An extensive review of the mechanisms and mechanical properties of porous-matrix CMCs is given in [18,19]. While CMCs exhibit improved damage tolerance compared with monolithic ceramics, two-dimensional laminated CMCs are more susceptible to failure in the matrix-rich interlaminar regions. The interlaminar failure or delamination may ultimately lead to loss of stiffness and accelerate structural failure [20]. A number of studies assessed the behavior of CMCs in shear [20–23]. Choi et al. [24–27] assessed the high-temperature life limiting behavior in interlaminar shear of several non-oxide CMCs, including three SiC fiber-reinforced CMCs and one carbon-fiber-reinforced CMC. Choi and co-workers determined the interlaminar shear strength

M.B. Ruggles-Wrenn, P.D. Laffey / Composites Science and Technology 68 (2008) 2260–2266

as a function of test rate using double-notch shear specimens and demonstrated that the interlaminar shear strength (ILSS) degraded with decreasing test rate. A phenomenological, power-law based crack growth model was proposed to account for the degradation of the ILSS of the composite at elevated temperatures. High-temperature stress-rupture tests in interlaminar shear were employed to validate the proposed model. These studies focused on the nonoxide CMCS with matrix-dense interlaminar regions, where the interlaminar failure is controlled by the fiber-matrix interface. In the case of the porous-matrix oxide–oxide CMCs, interlaminar shear failure is controlled by the exceptionally weak porous matrix. The objective of this effort is to evaluate the interlaminar shear strength and to investigate the creep behavior in interlaminar shear of an oxide–oxide CMC consisting of a porous alumina matrix reinforced with the NextelTM720 fibers. Several previous studies examined the in-plane mechanical behavior of this composite [28–31] at elevated temperature. This study investigates creep behavior of the NextelTM720/alumina (N720/A) composite in interlaminar shear at 1200 °C in air and in steam environments. The composite microstructure, as well as damage and failure mechanisms are discussed. 2. Material and experimental arrangements The material studied was NextelTM720/alumina (N720/A), an oxide–oxide CMC (manufactured by COI Ceramics, San Diego, CA)

2261

consisting of a porous alumina matrix reinforced with NextelTM720 fibers. There is no fiber coating. The composite was supplied in a form of 5.2-mm thick plates comprised of 24 0°/90° woven layers, with a density of 2.83 g/cm3, a fiber volume of 46.6%, and matrix porosity of 22%. The fiber fabric was infiltrated with the matrix in a sol–gel process. The laminate was dried with a ‘‘vacuum bag” technique under low pressure and low temperature, and then pressureless sintered [32]. Representative micrograph of the untested material is presented in Fig. 1a, which shows 0° and 90° fiber tows as well as numerous matrix cracks. In the case of the as-processed material, most are shrinkage cracks formed during processing. Porous nature of the matrix is seen in Fig. 1b. The double-notch shear (DNS) test specimens measuring 150 mm  20 mm were cut from a single N720/A panel. The thickness of the specimens was the same as the nominal thickness of the composite panel, i.e. 5.2 mm. The notches of 0.5 mm width were extended to the middle of each test specimen within ±0.05 mm so that shear failure occurred on the plane between the notch tips. The distance between the notches was 13 mm. Schematic of the DNS specimen and notch details are shown in Fig. 2a and b, respectively. Dimensions of the DNS specimens used in this effort were different from those recommended in the ASTM Standard C1425. The 13-mm distance between the notches was chosen specifically to enable the measurement of compressive strain between the notch tips with an MTS high-temperature extensometer of 12.5-mm gage length. The overall specimen

Fig. 1. As-processed material: (a) overview and (b) porous nature of the matrix is evident.

Fig. 2. Double-notch shear specimen: (a) configuration and dimensions and (b) notch details.

M.B. Ruggles-Wrenn, P.D. Laffey / Composites Science and Technology 68 (2008) 2260–2266

length of 150 mm ensures that the local stress fields at the notch tips are not influenced by the external loading at the specimen ends. A servocontrolled MTS mechanical testing machine equipped with hydraulic water-cooled collet grips, a compact two-zone resistance-heated furnace, and two temperature controllers was used in all tests. An MTS TestStar II digital controller was employed for input signal generation and data acquisition. For elevated temperature testing, thermocouples were bonded to the specimens using alumina cement (Zircar) to calibrate the furnace on a periodic basis. Tests in steam environment employed an alumina susceptor (tube with end caps), which fits inside the furnace. The specimen gage section is located inside the susceptor, with the ends of the specimen passing through slots in the susceptor. Steam is introduced into the susceptor (through a feeding tube) in a continuous stream with a slightly positive pressure, expelling the dry air and creating a near 100% steam environment inside the susceptor. Strain measurement was accomplished with an MTS high-temperature air-cooled uniaxial extensometer of 12.5-mm gage length. For strain measurement, the extensometer rods were placed as close to the notch tips of the specimen as possible. Fracture surfaces of failed specimens were examined using an SEM (FEI Quanta 200 HV) as well as an optical microscope (Zeiss Discovery V12). The SEM specimens were carbon coated. All tests were conducted at 1200 °C. In all tests, a specimen was heated to the test temperature at a rate of 1 °C/s, and held at temperature for additional 30 min prior to testing. To investigate the interlaminar shear properties and stress-rupture behavior, the DNS specimens were loaded in compression along the specimen axis, as shown in Fig. 2a. Monotonic tests were performed in displacement control at the displacement rate of 0.05 mm/s in laboratory air and in steam environment. In creep-rupture tests specimens were loaded to the creep stress level at the stress rate of 5 MPa/s. Creep run-out was defined as 100 h at a given creep stress. In each test, stress and strain data were recorded during the loading to the creep stress level and the actual creep period. Thus both total strain and creep strain could be calculated and examined. To determine the retained interlaminar strength, specimens that achieved creep run-out were subjected to monotonic tests to failure at 1200 °C. The shear stress values reported here represent the average stress, s, between the notches along the prospective shear plane calculated as s¼

Pa WLn

Fig. 3. Micrograph of a typical specimen after interlaminar shear failure (side view).

12

T = 1200 ºC 10

Shear Stress (MPa)

2262

Air

8 6 Steam

4 2 0 0.00

0.05

0.10

0.15

0.20

Compressive Strain (%) Fig. 4. Interlaminar shear stress vs compressive strain for N720/A composite at 1200 °C in air and in steam.

ð1Þ

where Pa is the applied force, W is the specimen width and Ln is the distance between the notches. Although the shear stress distribution between the notches is not uniform, the average stresses provided by Eq. (1) are useful when comparing interlaminar shear strength values and stress-rupture responses of specimens subjected to identical mechanical tests in different environments. It is worthy of note that in all tests reported below, the DNS specimens failed in shear mode along the shear plane. Fig. 3 shows a typical shear failure of a DNS specimen. 3. Results and discussion 3.1. Interlaminar shear strength The monotonic stress–strain response is presented in Fig. 4. Three specimens were tested in air and three specimens were tested in steam. In air, the average interlaminar shear strength (ILSS) was 8.25 MPa. Such relatively low value of the ILSS (compared to the ILSS values for the dense-matrix CMCs) can be attributed to the high matrix porosity. Note that Levi et al. [15] obtained

similar ILSS values (8 MPa) in short-beam shear tests conducted at room temperature on the N720/mullite-alumina specimens subjected to prior heat treatment at 1200 °C. The DNS specimen tested in steam was aged for 24 h at 1200 °C in steam prior to testing. Prior aging in steam had negligible effect on ILSS of N720/A composite. The specimens aged at 1200 °C in steam produced the average strength value of 8.44 MPa. However, exposure to steam noticeably influenced the stress–strain behavior of the composite. While the stress–strain curve obtained in air departs from linearity at the shear stress of approximately 4 MPa, the stress–strain curve obtained in steam becomes markedly nonlinear at a much lower shear stress of 1.5 MPa. As the shear stress continues to increase in the test conducted in steam, appreciable inelastic strain develops. Failure strain measured in steam is some 40% greater than that obtained in air. 3.2. Creep-rupture Results of the creep-rupture tests are summarized in Table 1, where rupture time and creep strain accumulation are shown for each test environment and applied shear stress level. Creep strain

2263

M.B. Ruggles-Wrenn, P.D. Laffey / Composites Science and Technology 68 (2008) 2260–2266 Table 1 Results of creep-rupture tests in interlaminar shear for the N720/A ceramic composite at 1200 °C in laboratory air and in steam environment

Laboratory air 6.5 6.5 Steam 4.0 5.0 6.5 a

Creep strain (%)

Time to rupture (s)

0.14 0.24

360,000a 360,000a

1.65 3.21 2.13

360,000a 29,040 20,400

Air

Creep Strain Rate (s-1)

Creep stress (MPa)

1.E-03

Run-out.

1.E-04

Steam

1.E-05

1.E-06

1.E-07 T = 1200 ºC

4.0

T = 1200 ºC

Creep Strain (%)

5 MPa 3.0

2.0

6.5 MPa 20000

40000

60000

Fig. 6. Minimum creep rate as a function of applied stress for N720/A ceramic composite at 1200 °C in air and in steam.

10

T = 1200 ºC 8

ILSS, Air

6

4

2

Air Steam

0 1.E+03

1.E+04

1.E+05

1.E+06

Time (s) Fig. 7. Interlaminar shear stress vs time to rupture for N720/A composite at 1200 °C.

Retained interlaminar shear strength values of the specimens that achieved a run-out at 6.5 MPa in air and at 4 MPa in steam are given in Table 2. The stress–strain curves obtained for the N720/A specimens subjected to prior creep in interlaminar shear in air and in steam are presented in Fig. 8a and b, respectively. The ILSS of the specimen pre-crept at 6.5 MPa in air has increased by nearly 37% compared to the ILSS of the as-processed specimen. Conversely, prior creep in steam has degraded the interlaminar shear strength of N720/A by about 27%. The specimen pre-crept at 4 MPa in steam retained less than 75% of its ILSS. Prior creep in either environment had little qualitative effect on stress–strain behavior.

Table 2 Retained interlaminar shear properties of the N720/A specimens subjected to prior creep in interlaminar shear at 1200 °C

4 MPa

0

10

Creep Stress (MPa)

A typical fracture surface of the DNS specimen tested in compression to failure at 1200 °C in air is shown in Fig. 9. Delamination

6.5 MPa

0.0

1

3.3. Composite microstructure

Air Steam

1.0

1.E-08

Shear Stress (MPa)

vs time curves obtained at 1200 °C in air and in steam are shown in Fig. 5. The creep curves obtained at 6.5 MPa in air exhibit primary and secondary creep regimes. Transition from primary to secondary creep occurs fairly early in creep life. Secondary creep is likely to persist for the duration of the creep lifetime. In air, creep run-out of 100 h is achieved at the shear stress of 6.5 MPa (78.5% ILSS). The strains accumulated during 100 h at 6.5 MPa are comparable to those obtained in the monotonic test. Creep curve produced in steam at the shear stress of 4 MPa also exhibits only primary and secondary creep regimes. In contrast, creep curves obtained in steam at 5 and 6.5 MPa show primary, secondary and tertiary creep. Transition from primary to secondary creep occurs almost immediately. Secondary creep persists for 70% of the creep life before transitioning to tertiary creep. Creep strain accumulation first increases as the applied shear stress increases from 4 to 5 MPa, then decreases as the applied stress increases to 6.5 MPa. It is noteworthy that in steam, all accumulated creep strains are at least an order of magnitude higher than the failure strain obtained in the monotonic test. In steam creep run-out was achieved only at 4 MPa (50% ILSS). Minimum creep rate was measured in all tests. Creep strain rate as a function of applied stress is shown in Fig. 6. In steam, the minimum creep rate increases by a factor of 10 when applied stress increases from 4 to 5 MPa. At 6.5 MPa, creep rate in steam is at least two orders of magnitude higher than that in air. Stress-rupture behavior is summarized in Fig. 7, where applied shear stress is plotted vs time to rupture at 1200 °C in air and in steam. In air, creep life (up to 100 h) appears to be relatively independent of applied stress up to 78% ILSS. All creep tests conducted at 6.5 MPa in air achieved a run-out. For applied shear stress P5 MPa the presence of steam dramatically reduced creep lifetimes. At 6.5 MPa, the reduction in creep life due to steam was 94% as compared to the run-out of 100 h.

80000

100000

Time (s) Fig. 5. Creep strain vs time curves for N720/A composite obtained at applied interlaminar shear stresses in the 4–6.5 MPa range at 1200 °C in air and in steam.

Creep stress (MPa)

Retained interlaminar shear strength (MPa)

Failure strain (%)

Laboratory air 6.5

11.2

0.22

Steam 4.0

6.13

0.19

2264

M.B. Ruggles-Wrenn, P.D. Laffey / Composites Science and Technology 68 (2008) 2260–2266

12

Shear Stress (MPa)

10

As-processed

8 100 h at 6.5 MPa in air

6 4 2 0 0.00

T = 1200 ºC, Air 0.05

0.10

0.15

0.20

0.25

Compressive Strain (%) 12

T = 1200 ºC, Steam

Shear Stress (MPa)

10

24 h at 0 MPa in steam 8 6 4 2

100 h at 4 MPa in steam 0 0.00

0.05

0.10

0.15

0.20

0.25

Compressive Strain (%) Fig. 8. Effects of prior creep in interlaminar shear on interlaminar shear stress– compressive strain behavior of N720/A ceramic composite at 1200 °C: (a) in air and (b) in steam.

of the woven 0°/90° fiber layers from the matrix-rich regions appears to be the primary mechanism of interlaminar shear failure. While most of the fracture surface shown in Fig. 9a is fairly smooth

and clean, indicating that only a single fiber layer is associated with delamination, some rough areas exposing debris and fiber breakage are also visible. In the course of delamination the departing fibers leave distinct troughs in the remaining matrix (see higher magnification views in Fig. 9b and c). Small amounts of the matrix material remain bonded to the fibers exposed during delamination. By contrast, the fracture surface of the DNS specimen tested in compression following 100 h of prior creep at 6.5 MPa at 1200 °C in air (see Fig. 10a and b) reveals that the failure mechanism in this case includes considerable fiber fracture. The rougher fracture surface shows increased damage in fiber tows, frequently exposing multiple 0°/90° fiber layers. Only some limited areas show clean delamination of a single fiber layer from the matrix-rich regions (Fig. 10a). As seen in Fig. 10c, the failure also involves extensive damage to the matrix. Recent studies [33,34] demonstrated that for a composite consisting of NextelTM720 fibers in a porous alumina matrix, a porosity reduction of 6% was observed after a 10-min exposure at 1200 °C, which was caused by additional sintering of the matrix. It is likely that additional sintering of the matrix occurred during the 100 h creep test at 6.5 MPa. The resultant strengthening of the matrix is manifested in the retained ILSS of the composite. Results in Table 2 show that the ILSS increased after prior creep at 6.5 MPa. The strengthening is also manifested in the change of the failure mechanism. The failure of the pre-crept composite involves extensive fiber fracture, while the as-processed material fails predominantly through matrix damage and interply delamination. The fracture surfaces obtained in steam (Figs. 11 and 12) are considerably more violent and rough than those obtained in air. The fracture surface produced in creep at 6.5 MPa (Fig. 11a and b) as well as the fracture surface obtained in compression test on a specimen subjected to 100 h of prior creep at 4 MPa (Fig. 12a and b) reveal extensive fracture of fiber tows. As seen in Figs. 11c and 12c, the amount of matrix material remaining bonded to the exposed fibers is greater than that in the specimens tested in air. In fact, it appears that during tests of over 100 h duration conducted in steam, the fiber tows become bonded together by the matrix material and then fail in coordinated fashion. It is possible that the sintering of the matrix is accelerated in the presence of

Fig. 9. Fracture surface of the DNS specimen tested in compression to failure at 1200 °C in air.

Fig. 10. Fracture surface of the DNS specimen tested in compression to failure following 100 h at 6.5 MPa at 1200 °C in air.

M.B. Ruggles-Wrenn, P.D. Laffey / Composites Science and Technology 68 (2008) 2260–2266

2265

Fig. 11. Fracture surface of the DNS specimen tested in creep at 6.5 MPa at 1200 °C in steam.

Fig. 12. Fracture surface of the DNS specimen tested in compression to failure following 100 h at 4.0 MPa at 1200 ° in steam.

steam. However, while prior creep in air resulted in strengthening of the matrix, retained properties in Table 2 suggest that prior creep in steam causes weakening of the matrix. The retained ILSS of the specimen pre-crept in steam is noticeably lower than the ILSS of the as-processed composite. Recent studies [35,36] show that water attacks grain boundaries and degrades the strength of the polycrystalline alumina. Kronenberg et al. [36] found two types of hydrogen defects incorporated in alumina specimens subjected to heat treatment at temperatures ranging from 850 °C to 1025 °C under 1500 MPa hydrostatic pressure in the presence of water, followed by compression tests at temperatures in the 630–850 °C range. The infrared measurements revealed interstitial hydrogens in the bulk and molecular water clusters near surfaces, grain boundaries and cracks of the hydrothermally treated alumina specimens. Furthermore, the presence of hydrogen defects reduced the yield stress of fine-grained alumina by a factor of six. It is possible that hydrogen defects that are introduced into the alumina matrix of the N720/A specimens during tests conducted at 1200 °C in steam are behind the degradation of creep performance in steam.

creep run-out was achieved only at 4 MPa. The run-out specimen retained only 75% of its ILSS. For applied stress P5 MPa, the presence of steam drastically reduced creep lifetimes. The hydrothermal weakening of the alumina matrix due to incorporation of hydrogen defects during testing at 1200 °C in steam may be behind the degraded creep performance of the N720/A in steam. In tests of short duration conducted in air, the failure occurs primarily through delamination of the woven 0°/90° fiber layers from the matrix-rich regions, with minimal fiber fracture. Generally, only one fiber layer is associated with delamination. For test durations >100 h, the failure mechanism includes considerable fiber fracture. It is possible that the matrix undergoes additional sintering during the long-term tests conducted in air. In tests conducted in steam, the failure mechanism is dominated by fiber fracture.

4. Concluding remarks

References

The creep behavior of the N720/A ceramic composite in interlaminar shear was assessed and the interlaminar shear properties were measured at 1200 °C in laboratory air and in steam using double-notch shear test specimens. The interlaminar shear strength (ILSS) was 8.25 MPa. Prior aging for 24 h at 1200 °C in steam had no effect on ILSS. The N720/A composite exhibits primary and secondary creep regimes in air. In steam, primary, secondary and tertiary creep regimes are observed. Creep strains accumulated in steam are at least an order of magnitude larger than those produced in air. Creep strain rates were approximately 2.1  107 s1 in air, and ranged from 3.4  106 to 7.0  105 s1 in steam. At 6.5 MPa, creep rate in steam is at least two orders of magnitude higher than the creep rate produced in air. Creep run-out was achieved in all tests conducted in air. The run-out specimens exhibited substantial increase in ILSS. In steam,

Acknowledgements The financial support of Dr. R. Sikorski and Dr. J. Zelina, Propulsion Directorate, Air Force Research Laboratory is highly appreciated.

[1] Ohnabe H, Masaki S, Onozuka M, Miyahara K, Sasa T. Potential application of ceramic matrix composites to aero-engine components. Composites: Part A 1999;30:489–96. [2] Zawada LP, Staehler J, Steel S. Consequence of intermittent exposure to moisture and salt fog on the high-temperature fatigue durability of several ceramic–matrix composites. J Am Ceram Soc 2003;86(8):1282–91. [3] Parlier M, Ritti MH. State of the art and perspectives for oxide/oxide composites. Aerospace Sci Technol 2003;7:211–21. [4] Mattoni MA, Yang JY, Levi CG, Zok FW, Zawada LP. Effects of combustor rig exposure on a porous-matrix oxide composite. Int J Appl Ceram Technol 2005;2(2):133–40. [5] Parthasarathy TA, Zawada LP, John R, Cinibulk MK, Zelina J. Evaluation of oxide–oxide composites in a novel combustor wall application. Int J App Ceram Technol 2005;2(2):122–32. [6] Szweda A, Millard ML, Harrison MG. Fiber-reinforced ceramic–matrix composite member and method for making. US Patent No. 5 601 674, 1997. [7] Sim SM, Kerans RJ. Slurry infiltration and 3-D woven composites. Ceram Eng Sci Proc 1992;13(9–10):632–41. [8] Moore EH, Mah T, Keller KA. 3D composite fabrication through matrix slurry pressure infiltration. Ceram Eng Sci Proc 1994;15(4):113–20.

2266

M.B. Ruggles-Wrenn, P.D. Laffey / Composites Science and Technology 68 (2008) 2260–2266

[9] Lange FF, Tu WC, Evans AG. Processing of damage-tolerant, oxidation-resistant ceramic matrix composites by a precursor infiltration and pyrolysis method. Mater Sci Eng A 1995;A195:145–50. [10] Mouchon E, Colomban P. Oxide ceramic matrix/oxide fiber woven fabric composites exhibiting dissipative fracture behavior. Composites 1995;26:175–82. [11] Tu WC, Lange FF, Evans AG. Concept for a damage-tolerant ceramic composite with strong interfaces. J Am Ceram Soc 1996;79(2):417–24. [12] Evans AG, Zok FW. Review the physics and mechanics of fiber-reinforced brittle matrix composites. J Mater Sci 1994;29:3857–96. [13] Kerans RJ, Parthasarathy TA. Crack deflection in ceramic composites and fiber coating design criteria. Composites: Part A 1999;30:521–4. [14] Kerans RJ, Hay RS, Parthasarathy TA, Cinibulk MK. Interface design for oxidation-resistant ceramic composites. J Am Ceram Soc 2002;85(11):2599–632. [15] Levi CG, Yang JY, Dalgleish BJ, Zok FW, Evans AG. Processing and performance of an all-oxide ceramic composite. J Am Ceram Soc 1998;81:2077–86. [16] Hegedus AG. Ceramic bodies of controlled porosity and process for making same. US Patent No. 5 0177 522, May 21, 1991. [17] Lu T. Crack branching in all-oxide ceramic composites. J Am Ceram Soc 1996;79:266–74. [18] Zok FW, Levi CG. Mechanical properties of porous-matrix ceramic composites. Adv Eng Mater 2001;3(1-2):15–23. [19] Zok F. Developments in oxide fiber composites. J Am Ceram Soc 2006;89(11):3309–24. [20] Fang NJ, Chou TW. Characterization of interlaminar shear strength of ceramic matrix composites. J Am Ceram Soc 1993;76(10):2539–48. [21] Brondsted P, Heredia FE, Evans AG. In-plane shear properties of 2-D ceramic composites. J Am Ceram Soc 1994;77(10):2569–74. [22] Lara-Curzio E, Ferber MK. Shear strength of continuous fiber ceramic composites. In: Jenkins M, Gonczy S, Lara-Curzio E, Ashbaugh N, Zawada L, editors. Thermal and mechanical test methods and behavior of continuous fiber ceramic composites. ASTM STP 1309; 2000. [23] Unal O, Bansal NP. In-plane and interlaminar shear strength of a unidirectional Hi-Nicalon fiber-reinforced Celsian matrix composite. Ceram Int 2002;28:527–40. [24] Choi SR, Bansal NP. Interlaminar tension/shear properties and stress rupture in shear of various continuous fiber-reinforced ceramic matrix composites. Ceram Trans 2006;175:119–34.

[25] Choi SR, Bansal NP. Shear strength as a function of test rate for SiCf/BSAS ceramic matrix composite at elevated temperature. J Am Ceram Soc 2004;87(10):1912–8. [26] Choi SR, Bansal NP, Calomino AM, Verrilli MJ. Shear strength behavior of ceramic matrix composites at elevated temperatures. Ceram Trans 2005;165:131–45. [27] Choi SR, Kowalik RW, Alexander DJ, Bansal NP. Assessments of life limiting behavior in interlaminar shear for Hi–NiC SiC/SiC ceramic matrix composite at elevated temperature. In: Lara-Curzio E, Salem J, Zhu D, editors. Mechanical properties and performance of engineering ceramics and composites III. John Wiley & Sons, Inc.; 2007. p. 179–89. [28] Ruggles-Wrenn MB, Mall S, Eber CA, Harlan LB. Effects of steam environment on high-temperature mechanical behavior of NextelTM 720/alumina (N720/A) continuous fiber ceramic composite. Composites A 2006;37(11):2029–40. [29] Mehrman JM, Ruggles-Wrenn MB, Baek SS. Influence of hold times on the elevated-temperature fatigue behavior of an oxide–oxide ceramic composite in air and in steam environment. Compos Sci Technol 2007;67:1425–38. [30] Ruggles-Wrenn MB, Hetrick G, Baek SS. Effects of frequency and environment on fatigue behavior of an oxide–oxide ceramic composite at 1200 °C. Int J Fatigue 2008;30:502–16. [31] Ruggles-Wrenn MB, Siegert GT, Baek SS. Effects of environment on creep behavior of an oxide–oxide ceramic composite with ±45° fiber orientation at 1200 °C. Compos Sci Technol 2008;68:1588–95. [32] Jurf RA, Butner SC. Advances in oxide–oxide CMC. J Eng Gas Turb Power 1999;122(2):202–5. [33] Fujita H, Jefferson G, McMeeking RM, Zok FW. Mullite/alumina mixtures for use as porous matrices in oxide fiber composites. J Am Ceram Soc 2004;87(2):261–7. [34] Fujita H, Levi CG, Zok FW, Jefferson G. Controlling mechanical properties of porous mullite/alumina mixtures via precursor-derived alumina. J Am Ceram Soc 2005;88(2):367–75. [35] Tai WP, Watanabe T. High-temperature stability of alumina in argon and argon/water-vapor environments. J Am Ceram Soc 1999;82(1):245–8. [36] Kronenberg AK, Castaing J, Mitchell TE, Kirby SH. Hydrogen defects in a-Al2O3 and water weakening of sapphire and alumina ceramics between 600 °C and 1000 °C – I. Infrared characterization of defects. Acta Mater 2000;48:1481–94.