December 1997 Control of Interfacial Properties through Fiber Coatings: Monazite Coatings 2995 The coefficient of friction, u, in the YAG fiber system(0. 18) the fiber/coating interface, is-45o, which is similar to the mode was lower than that of the Al,O3 fiber system(0. 24). These mixity of the interface crack in the UCSB specimen. 37, 42As the differences are expected, because the sliding interface is dif- residual stress state changes, the mode mixity of the kink ferent in each case: YAG on LaPO4, for the YAG fiber com- crack also changes; it becomes more Mode ll with residual ite,and Al2O3 against LaPO4, for the Al,O3 fiber compos- compression and more Mode I with residual tension. The key point is that the pure Mode Il fracture resistance that is mea- The interfacial fracture energy Ti in the YAG fiber system ured using the pushout test gives an extremely conservative was constant as the coating thickness changed and was-165(and inaccurate)estimate for the He and Hutchinson debonding J/m2(Table II) Debonding and sliding were observed to occur criterion It the YAG/LapO4 interface. This system did not develop any eaction product between the coating and the fiber at the V. Conclusions debonding interface(Fig. 9(b))and, therefore, had a T value that remained constant as the coating thickness changed. a This study quantified the effect of the thickness of a mona- -Al2O, reaction layer was found at the coating/matrix side of zite(LaPO4) coating on the interfacial properties of alumina he interface: however, it had no effect on debonding and slid- (AL,O3)and YAG fibers that were embedded in an Al,O, ma- trix. The different thermal and elastic properties of the con- stituents resulted in different distributions of residual stresses ( Summary for each fiber system. An increase in the coating thickness It is known that coatings will modify the thermomechanical decreased the load that was required for debonding in the and thermochemical properties of ceramic composites. 2 Test Al2O3 fiber system, whereas a slight increase in the required res ults in a carbon-coated silicon carbide(SiC) fiber system load was observed for the YAG fiber system. These showed that the interfacial shear strength and the coefficient of mentally observed changes were consistent with predictions friction u decreased as the coating thickness increased 26,2 that have been extracted from the Liang and Hutchinson(LH) The carbon interlayer was hypothesized to form a compliant pushout model20 and illustrated how the interface properties layer that mitigated the effect of thermal stress: thicker carbon can be altered through changes in the coating thickness oatings reduced the thermal stresses more than thinner coat- An experimental methodology was presented for extracting ings. 28 The same effect was e entally observed for the the key interfacial parameters from pushout tests on fiber tems that have been evaluated in this study: i.e., the coating reinforced composites. The extracted parameters included the ts as a compliant layer that alters the residual stress state of interfacial fracture energy(), the coefficient of interfacial Morgan and Marshall 3 measured a mixed-mode T value which incorporated interfacial roughness). The proposed meth using the UCSB four-point flexure tests, which resulted in a odology used the Lh pushout model to curve-fit results of mixed-mode t value of 5 J/m2 The relation between the in pushout tests that had been conducted on specimens of varying terfacial fracture resistance, T, and the phase angle, y, has sample thickness. This method mechanistically rationalized the been established by evans et al s and has been applied to effect of the coating thickness on the mechanical properties d- the fiber/coating interface, which led to a detailed understa Mode ll (Y /2)i value can be a factor of >4 larger than ing of the mechanics of fiber debonding and sliding. As such, 45)36, 37Therefore, the relatively large Mode II Ti value that can be used to establish target levels of pushout stresses; more has been obtained in this study is reasonable, when compared importantly, it can be used to determine the required coat with the mixed-mode Ti value that was measured by Morgan hickness for obtaining a target interfacial sliding stress and Marshall. 13 The resulting Ti values from the fiber pushout tests represen Acknowledgments: The authors would like to acknowledge Dr R hay of wright-Patterson Air Force Base(Wright-Patterson AFB, OH) for supply values must be related to the he and hutchinson interfacial the YAg fibers. Prof John Hutch bonding criterion s to ascertain the value of these coatings for critical comments, and Dr. David Marshall for his insight and interest. for use in structural composites. He and Hutchinsons found References that a propagating crack will deflect up the fiber/coating inter M. D. Thouless, O. Sbaizero, L. S. Sigl, and A. G. Evans, Inter- ace if the ratio of the interfacial debond toughness to the fiber toughness is <0.25. Fracture-energy values for sapphire and YAG fibers, along with the measured Mode ll interfacial frac- R. w. Rice, J.R. Spann, D. Lewis, and w. Coblenz, " The Fiber Coatings on the Room Temperature Mechanical Behavior of Ceramic- ture toughness, are given in Table Ill. Clearly, neither the Fiber Composites, " Ceram Eng. Sci. Proc. 5(7-8)614-24(1984) Al2O3 nor the YAG fiber systems satisfy the He and Hutchin- R.J.Kerans, R. S Hay, N. J. Pagano, and T. A Parthasarathy, The Role of son debonding crite if the pure Mode Il toughness is used. However, the fracture testing that was reported by Mor- 1 e 429e-42 (1989 tertace n Ceramic Composites," Am. Ceram. So:, Bul, 68 an and Marshalls clearly showed interfacial debonding and 'A G. Evans and D. B. Marshall. The CB和M,3710 ical behavior of ceramic (1989) sliding for Al,O, fibers that were coated with LaPOa. Thus, the ole of Interfaces in Fiber. issue of the mode mixity of the kinked crack is relevant to Reinforced Brittle Matrix Composites, Compos. Sci. TechnoL., 42, 3-24 deciding the validity of a given coating system. In a residual stress-free system, the mode mixity of a crack, as it kinks up T Mah, M. G. Mendiratta, A P. Katz, and K.S. Mazdiyasni, "Recent De- Cera.Soc.Ba,662]304308(1987 7R F. Cooper and K. Chyung, "Structure and Chemistry of Fiber-Matrix Interfaces in Silicon Carbide Fiber Reinforced Glass-Ceramic Composites: An Table Ill. Comparisons of Fiber Fracture Energy(Td and Electron Microscopy Study, J. Mater. Sci., 22 [913148-60(1987). Mode Il Interfacial Fracture Energy (i ) between AlO3 and structural YAG Fiber Systems with a LaPO4 coating SiC-Fiber-Reinforced Lithium Aluminum Silicate gl ic,J.Am. Ceram. Soc., 72 5]741-45( Mode ll interfacial E L. Courtright, ""Engineering Property Limitations of Structural Ceramics Fiber fracture Ceramic Composites Above 1600C, ""Ceram. Eng. Sct. Proc., 12 [9-101 Fiber LaPO coating system T(/m2) G S. Corman, ""High-Temperature Creep of Some Single Crystal Oxides, 12-20t 10-12 fFrom davis et al 39 *From Reimanis et al-o and blumenthal P E D Morgan, D B. Marshall, and R. M. Housley, " "High TemperatureThe coefficient of friction, m, in the YAG fiber system (0.18) was lower than that of the Al2O3 fiber system (0.24). These differences are expected, because the sliding interface is dif￾ferent in each case: YAG on LaPO4, for the YAG fiber com￾posite, and Al2O3 against LaPO4, for the Al2O3 fiber compos￾ite. The interfacial fracture energy Gi in the YAG fiber system was constant as the coating thickness changed and was ∼16.5 J/m2 (Table II). Debonding and sliding were observed to occur at the YAG/LaPO4 interface. This system did not develop any reaction product between the coating and the fiber at the debonding interface (Fig. 9(b)) and, therefore, had a Gi value that remained constant as the coating thickness changed. A b-Al2O3 reaction layer was found at the coating/matrix side of the interface; however, it had no effect on debonding and slid￾ing. (3) Summary It is known that coatings will modify the thermomechanical and thermochemical properties of ceramic composites.2 Test results in a carbon-coated silicon carbide (SiC) fiber system showed that the interfacial shear strength and the coefficient of friction m decreased as the coating thickness increased.26,27 The carbon interlayer was hypothesized to form a compliant layer that mitigated the effect of thermal stress: thicker carbon coatings reduced the thermal stresses more than thinner coat￾ings.28 The same effect was experimentally observed for the systems that have been evaluated in this study: i.e., the coating acts as a compliant layer that alters the residual stress state of the system. Morgan and Marshall13 measured a mixed-mode Gi value using the UCSB four-point flexure tests, which resulted in a mixed-mode Gi value of 5 J/m2 . The relation between the in￾terfacial fracture resistance, Gi , and the phase angle, C, has been established by Evans et al.36 and has been applied to experimental results.37 Theoretically and experimentally, the Mode II (C 4 p/2) Gi value can be a factor of >4 larger than the mixed-mode value that is measured by flexure testing (C ≈ 45°).36,37 Therefore, the relatively large Mode II Gi value that has been obtained in this study is reasonable, when compared with the mixed-mode Gi value that was measured by Morgan and Marshall.13 The resulting Gi values from the fiber pushout tests represent pure Mode II measures of the fracture energy. However, these values must be related to the He and Hutchinson interfacial debonding criterion38 to ascertain the value of these coatings for use in structural composites. He and Hutchinson38 found that a propagating crack will deflect up the fiber/coating inter￾face if the ratio of the interfacial debond toughness to the fiber toughness is <0.25. Fracture-energy values for sapphire and YAG fibers, along with the measured Mode II interfacial frac￾ture toughness, are given in Table III. Clearly, neither the Al2O3 nor the YAG fiber systems satisfy the He and Hutchin￾son debonding criterion38 if the pure Mode II toughness is used. However, the fracture testing that was reported by Mor￾gan and Marshall13 clearly showed interfacial debonding and sliding for Al2O3 fibers that were coated with LaPO4. Thus, the issue of the mode mixity of the kinked crack is relevant to deciding the validity of a given coating system. In a residual￾stress-free system, the mode mixity of a crack, as it kinks up the fiber/coating interface, is ∼45°, which is similar to the mode mixity of the interface crack in the UCSB specimen.37,42 As the residual stress state changes, the mode mixity of the kinked crack also changes; it becomes more Mode II with residual compression and more Mode I with residual tension. The key point is that the pure Mode II fracture resistance that is mea￾sured using the pushout test gives an extremely conservative (and inaccurate) estimate for the He and Hutchinson debonding criterion.38 V. Conclusions This study quantified the effect of the thickness of a mona￾zite (LaPO4) coating on the interfacial properties of alumina (Al2O3) and YAG fibers that were embedded in an Al2O3 ma￾trix. The different thermal and elastic properties of the con￾stituents resulted in different distributions of residual stresses for each fiber system. An increase in the coating thickness decreased the load that was required for debonding in the Al2O3 fiber system, whereas a slight increase in the required load was observed for the YAG fiber system. These experi￾mentally observed changes were consistent with predictions that have been extracted from the Liang and Hutchinson (LH) pushout model20 and illustrated how the interface properties can be altered through changes in the coating thickness. An experimental methodology was presented for extracting the key interfacial parameters from pushout tests on fiber￾reinforced composites. The extracted parameters included the interfacial fracture energy (Gi ), the coefficient of interfacial friction (m), and the interfacial clamping pressure (sclamping, which incorporated interfacial roughness). The proposed meth￾odology used the LH pushout model to curve-fit results of pushout tests that had been conducted on specimens of varying sample thickness. This method mechanistically rationalized the effect of the coating thickness on the mechanical properties of the fiber/coating interface, which led to a detailed understand￾ing of the mechanics of fiber debonding and sliding. As such, the LH model, in conjunction with the proposed methodology, can be used to establish target levels of pushout stresses; more importantly, it can be used to determine the required coating thickness for obtaining a target interfacial sliding stress. Acknowledgments: The authors would like to acknowledge Dr. R. Hay of Wright-Patterson Air Force Base (Wright-Patterson AFB, OH) for supplying the YAG fibers, Prof. John Hutchinson at Harvard University (Cambridge, MA) for critical comments, and Dr. David Marshall for his insight and interest. References 1 M. D. Thouless, O. Sbaizero, L. S. Sigl, and A. G. Evans, ‘‘Effect of Inter￾face Mechanical Properties on Pullout in a SiC-Fiber-Reinforced Lithium Alu￾minum Silicate Glass-Ceramic,’’ J. Am. Ceram. Soc., 72 [4] 525–32 (1989). 2 R. W. Rice, J. R. Spann, D. Lewis, and W. Coblenz, ‘‘The Effect of Ceramic Fiber Coatings on the Room Temperature Mechanical Behavior of Ceramic￾Fiber Composites,’’ Ceram. Eng. Sci. Proc., 5 [7–8] 614–24 (1984). 3 R. J. Kerans, R. S. Hay, N. J. Pagano, and T. A. Parthasarathy, ‘‘The Role of the Fiber–Matrix Interface in Ceramic Composites,’’ Am. Ceram. Soc. Bull., 68 [2] 429–42 (1989). 4 A. G. Evans and D. B. Marshall, ‘‘The Mechanical Behavior of Ceramic Matrix Composites,’’ Acta Metall., 37 [10] 2567–83 (1989). 5 A. G. Evans, F. W. Zok, and J. Davis, ‘‘The Role of Interfaces in Fiber￾Reinforced Brittle Matrix Composites,’’ Compos. Sci. Technol., 42, 3–24 (1991). 6 T. Mah, M. G. Mendiratta, A. P. Katz, and K. S. Mazdiyasni, ‘‘Recent De￾velopments in Fiber-Reinforced High-Temperature Ceramic Composites,’’ Am. Ceram. Soc. Bull., 66 [2] 304–308 (1987). 7 R. F. Cooper and K. Chyung, ‘‘Structure and Chemistry of Fiber–Matrix Interfaces in Silicon Carbide Fiber Reinforced Glass-Ceramic Composites: An Electron Microscopy Study,’’ J. Mater. Sci., 22 [9] 3148–60 (1987). 8 E. Bischoff, M. Ru¨hle, O. Sbaizero, and A. G. Evans, ‘‘Microstructural Studies of the Interfacial Zone of a SiC-Fiber-Reinforced Lithium Aluminum Silicate Glass-Ceramic,’’J. Am. Ceram. Soc., 72 [5] 741–45 (1989). 9 E. L. Courtright, ‘‘Engineering Property Limitations of Structural Ceramics and Ceramic Composites Above 1600°C,’’ Ceram. Eng. Sci. Proc., 12 [9–10] 1725–44 (1991). 10G. S. Corman, ‘‘High-Temperature Creep of Some Single Crystal Oxides,’’ Ceram. Eng. Sci. Proc., 12 [9–10] 1745–66 (1991). 11P. E. D. Morgan and D. B. Marshall, ‘‘Functional Interfaces for Ox￾ide/Oxide Composites,’’ Mater. Sci. Eng., A, A162, 15–25 (1993). 12P. E. D. Morgan, D. B. Marshall, and R. M. Housley, ‘‘High Temperature Table III. Comparisons of Fiber Fracture Energy (Gf ) and Mode II Interfacial Fracture Energy (Gi ) between Al2O3 and YAG Fiber Systems with a LaPO4 Coating Fiber system Fiber fracture energy, Gf (J/m2 ) Mode II interfacial fracture energy with LaPO4 coating, Gi (J/m2 ) Gi /Gf Al2O3 12–20† 12 >0.25 YAG 10–12‡ 16 >0.25 † From Davis et al.39 ‡From Reimanis et al.40 and Blumenthal.41 December 1997 Control of Interfacial Properties through Fiber Coatings: Monazite Coatings 2995