CERAMIC COMPOSITE INTERFACES Properties and design KT Faber Department of Materials Science and Engineering, Robert R. McCormick School of Engineering and Applied Science, Northwestern University, Evanston, Illinois 60208 KEY WORDS: ceramic-matrix composites, interfaces, interphases, residual stress, roughness ABSTRACT Optimal design of the fiber-matrix interface in ceramic-matrix composites is the key to achieving desired composite performance. In this paper the interface controlling parameters are described. Techniques for measuring interfacial prop- erties are reported. Examples of interface design of both oxide and non-oxide types are illustrated INTRODUCTION It is well established that the fiber-matrix interface is the dominant design pa- ameter in ceramic-matrix composites. The characteristic that sets brittle ma- trix composites apart from ductile composites, either metal or polymer, is the reliance on a relatively weak fiber-matrix interface for enhanced mechanical be- havior. Recognition of this phenomenon came as early as the early 1970s when Sambell et al (1)noted enhanced work of fracture in carbon fiber-reinforced glass and glass-ceramics where no chemical bond existed between fiber and matrix. In contrast, carbon fibers in magnesia and zirconia fibers in magnesia and glass were found to be chemically bonded and demonstrated little, if an The chemistry of interfaces was the sole design parameter over much of the ext 25 years. More recently, physical parameters, such as the thermal sion mismatch and fiber surface roughness(both relieved through cor oatings), were found to be equally important in interfacial design. The contained herein describes the interface controlling parameters, measurement 084-6600/97/0801-049990800
P1: ARK/MBL/rkc P2: MBL/vks QC: MBL/agr T1: MBL May 16, 1997 13:47 Annual Reviews AR034-16 Annu. Rev. Mater. Sci. 1997. 27:499–524 Copyright c 1997 by Annual Reviews Inc. All rights reserved CERAMIC COMPOSITE INTERFACES: Properties and Design K. T. Faber Department of Materials Science and Engineering, Robert R. McCormick School of Engineering and Applied Science, Northwestern University, Evanston, Illinois 60208; email: k-faber@nwu.edu KEY WORDS: ceramic-matrix composites, interfaces, interphases, residual stress, roughness ABSTRACT Optimal design of the fiber-matrix interface in ceramic-matrix composites is the key to achieving desired composite performance. In this paper the interfacecontrolling parameters are described. Techniques for measuring interfacial properties are reported. Examples of interface design of both oxide and non-oxide types are illustrated. INTRODUCTION It is well established that the fiber-matrix interface is the dominant design parameter in ceramic-matrix composites. The characteristic that sets brittle matrix composites apart from ductile composites, either metal or polymer, is the reliance on a relatively weak fiber-matrix interface for enhanced mechanical behavior. Recognition of this phenomenon came as early as the early 1970s when Sambell et al (1) noted enhanced work of fracture in carbon fiber–reinforced glass and glass-ceramics where no chemical bond existed between fiber and matrix. In contrast, carbon fibers in magnesia and zirconia fibers in magnesia and glass were found to be chemically bonded and demonstrated little, if any, toughening. The chemistry of interfaces was the sole design parameter over much of the next 25 years. More recently, physical parameters, such as the thermal expansion mismatch and fiber surface roughness (both relieved through compliant coatings), were found to be equally important in interfacial design. The review contained herein describes the interface controlling parameters, measurement 499 0084-6600/97/0801-0499$08.00
500 FABER techniques for the mechanical characterization of fiber-matrix interfaces, and key examples of interface design in brittle matrix composite INTERFACE DESIGN PARAMETERS Brittle matrix composites for structural applications are deemed successful when they exhibit fiber debonding and frictional sliding as cracks propagate hrough the matrix. Such behavior is generally accompanied by notch or flaw lerance as demonstrated in a highly nonlinear stress-strain curve reminiscent of a plastically deforming metal (2). The nonlinearity, however, derives not from plasticity but from initiation and propagation of a series of through-the thickness matrix cracks that are bridged by the brittle fibers. The low toughnes interface is the first requirement to prevent fiber fracture during matrix crack growth. Debonding of fibers alone, however, is not sufficient to provide the notch tolerance. As noted by Thouless Evans(3)and by Cao et al (4), the interfacial sliding resistance t should be small enough to allow for a substantial pullout contribution through frictional dissipation by encouraging fiber fracture at significant distances from the matrix crack plane. It is generally thought that t should be between 2 and 40 MPa(3). A third requirement is that thermal mismatch stresses between fiber and matrix are not substantial enough to cause either fiber or matrix crackin To examine how these parameters are explicitly incorporated into composite performance, we first investigate the analysis of Hutchinson Jensen (5), as modified by Marshall(6), that describes the displacement u of a single crack-bridging fiber in a brittle matrix. The analysis assumes stable, interfacial debonding with interfacial sliding governed by a Coulomb friction law and can be written as +r'-sa where u characterizes the frictional properties of the interface and is inversely proportional to the coefficient of friction u, which is related to the frictional sliding resistance t. A is a dimensionless constant that includes the elastic con- stants of the composite constituents, the volume fraction of fibers, the surface roughness, and the anisotropy in thermal misfit strains, Sro is a direct measure of the radial residual stress, Sa is the applied stress normalized by the peak pplied stress. r" is related to the debond energy and is directly proportional to the interfacial toughness Gic and includes any residual stress. The term SR includes the thermal mismatch strain a and the surface roughness-induced strain 8, which relates to the amplitude of the surface roughness A. Conse quently, a complex relationship between the interface toughness, the frictional
P1: ARK/MBL/rkc P2: MBL/vks QC: MBL/agr T1: MBL May 16, 1997 13:47 Annual Reviews AR034-16 500 FABER techniques for the mechanical characterization of fiber-matrix interfaces, and key examples of interface design in brittle matrix composites. INTERFACE DESIGN PARAMETERS Brittle matrix composites for structural applications are deemed successful when they exhibit fiber debonding and frictional sliding as cracks propagate through the matrix. Such behavior is generally accompanied by notch or flaw tolerance as demonstrated in a highly nonlinear stress-strain curve reminiscent of a plastically deforming metal (2). The nonlinearity, however, derives not from plasticity but from initiation and propagation of a series of through-thethickness matrix cracks that are bridged by the brittle fibers. The low toughness interface is the first requirement to prevent fiber fracture during matrix crack growth. Debonding of fibers alone, however, is not sufficient to provide the notch tolerance. As noted by Thouless & Evans (3) and by Cao et al (4), the interfacial sliding resistance τ should be small enough to allow for a substantial pullout contribution through frictional dissipation by encouraging fiber fracture at significant distances from the matrix crack plane. It is generally thought that τ should be between 2 and 40 MPa (3). A third requirement is that thermal mismatch stresses between fiber and matrix are not substantial enough to cause either fiber or matrix cracking. To examine how these parameters are explicitly incorporated into composite performance, we first investigate the analysis of Hutchinson & Jensen (5), as modified by Marshall (6), that describes the displacement u of a single, crack-bridging fiber in a brittle matrix. The analysis assumes stable, interfacial debonding with interfacial sliding governed by a Coulomb friction law and can be written as u u∗ = −ASRo ln µ SRo − Sa SRo − 00 ¶ + 00 − Sa, 1. where u∗ characterizes the frictional properties of the interface and is inversely proportional to the coefficient of friction µ, which is related to the frictional sliding resistance τ . A is a dimensionless constant that includes the elastic constants of the composite constituents, the volume fraction of fibers, the surface roughness, and the anisotropy in thermal misfit strains, SRo is a direct measure of the radial residual stress, Sa is the applied stress normalized by the peak applied stress. 00 is related to the debond energy and is directly proportional to the interfacial toughness Gic and includes any residual stress. The term SRo includes the thermal mismatch strain εT, and the surface roughness-induced strain εsr, which relates to the amplitude of the surface roughness Asr. Consequently, a complex relationship between the interface toughness, the frictional
CERAMIC COMPOSITE INTERFACES sliding resistance, and residual stress state determines the loac cement response. Each of these components is expounded upon below Interface Toughn Conditions for fiber pullout in a fiber-reinforced material first require that a propagating crack deflects along the interface rather than penetrates the fiber uch conditions have been aptly described by He Hutchison(7, 8). They evaluated the relative energy release rate for a crack deflecting along an interface and for the crack penetrating the interface. The competition between crack growth along the interface and penetration through a fiber was found to depend upon the ratio of the fracture energy of the interface Gic and that of the adjoining material Gc, in this case the fiber. In a system where the modulus and fiber have identical moduli, the interface toughness Gic must be less than one quarte of the matrix toughness G. This finding now serves as the basis of interface design. Shown in Figure I is a contour separating regions of crack deflection 0.5 -0.5 0.5 Dundurs' Parameter a Figure I Debonding map showing crack penetration and crack deflection regimes as a function of Dundurs' parameter(after 7, 8)
P1: ARK/MBL/rkc P2: MBL/vks QC: MBL/agr T1: MBL May 16, 1997 13:47 Annual Reviews AR034-16 CERAMIC COMPOSITE INTERFACES 501 sliding resistance, and residual stress state determines the load displacement response. Each of these components is expounded upon below. Interface Toughness Conditions for fiber pullout in a fiber-reinforced material first require that a propagating crack deflects along the interface rather than penetrates the fiber. Such conditions have been aptly described by He & Hutchison (7, 8). They evaluated the relative energy release rate for a crack deflecting along an interface and for the crack penetrating the interface. The competition between crack growth along the interface and penetration through a fiber was found to depend upon the ratio of the fracture energy of the interface Gic and that of the adjoining material Gc, in this case the fiber. In a system where the modulus and fiber have identical moduli, the interface toughness Gic must be less than one quarter of the matrix toughness Gc. This finding now serves as the basis of interface design. Shown in Figure 1 is a contour separating regions of crack deflection Figure 1 Debonding map showing crack penetration and crack deflection regimes as a function of Dundurs’ parameter (after 7, 8)
502 FABER and penetration as a function of Dundurs parameter a where Ef-em and the subscripts f and m refer to the fiber and the matrix, and Ex=Ex(I-v2) he plane strain modulus for the phase x. E and v are the elastic modu- lus and Poissons ratio for the respective phases. A positive value of a re- flects conditions of a fiber stiffer than the matrix, a likely occurrence even in brittle matrix composites. The magnitude of a rarely exceeds 0.5 in these materials Recently, Lee et al (9)analyzed conditions for kinking back into the re- nforcement after deflection-a problem more common to laminates than to fiber-reinforced materials, but important nonetheless liding Resistance The coefficient of friction u at the sliding interface plays a dominant role in ceramic composite toughening by determining the sliding resistance t. For a crack-bridging fiber, the bridging tractions are determined by the relative fiber-matrix displacements as controlled through the sliding resistance. At first glance. one would then wish to maximize the friction coefficient. However. as u and hence t increase, fiber fracture is expected closer to the crack plane(3) The contribution to toughening from pullout is then diminished. Systematic changes in t in a single system can be arrived at by altering the residual stress profile or fiber surface roughness, both described bel Roughness Fiber surface roughness was first acknowledged to influence interfacial prop- erties by Jero Kerans(10)who noticed that fibers would"reseat" with a oncomitant decrease in load when they were pushed back into the matrix to heir original position. The load drop was attributed to the residual sliding resis- tance resulting from the relaxation from roughness-induced misfit strain upon reseating. Additional evidence of the role of fiber surface roughness including stress birefringence of an asperity misfit (I1)and direct evidence of the role of fiber-surface roughness of as-processed fibers(11-16) have been reported Kerans Parthasarathy(17) have included the role of roughness in Equation I by treating it as an additional component to the effective interfacial clamping k(sth +ear) where g. is the radial thermal mismatch strain. gs is the roughness-induced radial misfit strain, and k is a constant accounting for the elastic properties of
P1: ARK/MBL/rkc P2: MBL/vks QC: MBL/agr T1: MBL May 16, 1997 13:47 Annual Reviews AR034-16 502 FABER and penetration as a function of Dundurs’ parameter α where α = E¯f − E¯m E¯f + E¯m , 2. and the subscripts f and m refer to the fiber and the matrix, and E¯x = Ex(1-ν2 x ), the plane strain modulus for the phase x. E and ν are the elastic modulus and Poisson’s ratio for the respective phases. A positive value of α re- flects conditions of a fiber stiffer than the matrix, a likely occurrence even in brittle matrix composites. The magnitude of α rarely exceeds 0.5 in these materials. Recently, Lee et al (9) analyzed conditions for kinking back into the reinforcement after deflection—a problem more common to laminates than to fiber-reinforced materials, but important nonetheless. Sliding Resistance The coefficient of friction µ at the sliding interface plays a dominant role in ceramic composite toughening by determining the sliding resistance τ . For a crack-bridging fiber, the bridging tractions are determined by the relative fiber-matrix displacements as controlled through the sliding resistance. At first glance, one would then wish to maximize the friction coefficient. However, as µ and hence τ increase, fiber fracture is expected closer to the crack plane (3). The contribution to toughening from pullout is then diminished. Systematic changes in τ in a single system can be arrived at by altering the residual stress profile or fiber surface roughness, both described below. Roughness Fiber surface roughness was first acknowledged to influence interfacial properties by Jero & Kerans (10) who noticed that fibers would “reseat” with a concomitant decrease in load when they were pushed back into the matrix to their original position. The load drop was attributed to the residual sliding resistance resulting from the relaxation from roughness-induced misfit strain upon reseating. Additional evidence of the role of fiber surface roughness including stress birefringence of an asperity misfit (11) and direct evidence of the role of fiber-surface roughness of as-processed fibers (11–16) have been reported. Kerans & Parthasarathy (17) have included the role of roughness in Equation 1 by treating it as an additional component to the effective interfacial clamping stress σn, as shown here: σn = k ¡ εth r + εsr¢ , 3. where εth r is the radial thermal mismatch strain, εsr is the roughness-induced radial misfit strain, and k is a constant accounting for the elastic properties of
CERAMIC COMPOSITE INTERFACES 503 the fiber and matrix. To first order E= A/R, where a is the amplitude of the surface roughness and R is the fiber radius. The amplitude of the face roughness then appears explicitly in Equation 1, because Sro is directl proportional to the clamping stress o Fiber surface roughness can be altered with compliant, low-fracture resis- tance coatings. A rationale for a coating scheme is best seen in Figure 2, where increasing coating thickness allows for systematic modification of the asperity asperity interactions of fiber and matrix. An increase in the coating thickness reduces the roughness asperity interactions between the fiber and matrix dur- ing interfacial sliding. This results in a smaller roughness-induced strain and clamping stress, which reduces the frictional resistance to sliding. A coating with a thickness greater than the amplitude of the fiber surface roughness can completely negate the contribution to the frictional sliding stress. This is best illustrated by the following example Fiber-sliding measurements were made on a model composite system by Mumm Faber(16). SiC monofilaments coated with four different thick- nesses of carbon(relative to the amplitude of the asperity roughness)in a soda-borosilicate glass matrix were used for model fiber pullout experiments In the extreme, the coating is meant to completely eliminate asperity contact during debonding and sliding. The fiber force-displacement measurements for the series are shown in Figure 3. First, the controlled thickness coatings induce systematic changes in the load fluctuations; their amplitude decreases ith increasing thickness and is essentially eliminated for coatings thicker thaI he roughness amplitude. The increased slope of the load-deflection curves likely due to an increase in the coefficient of friction with increasing coating thickness owing to changes in the real area of contact(18) Residual stress Few ceramic composites are free of residual stresses. The stresses derive from hermal expansion mismatch between fiber and matrix. The role of such stresses is obvious: As clamping stresses increase, the interfacial frictional stress in- creases. At the extreme, such stresses result in spontaneous cracking of the matrix for a fiber under residual compressive stresses(19). Conversely, sponta neous fiber debonding results when tensile stresses in the fiber exceed a critical value. In addition to interfacial sliding, residual stresses influence the condi- tions for deflection and/or penetration shown in Figure 1. Compressive residual stresses in the reinforcement enhance conditions for deflection, shifting the con- tour upward, while tensile stresses enhance penetration(20) Residual stresses can be controlled through the appropriate choice of the fiber-matrix pair. Moreover, the residual stress profile can be altered with fiber loading. Singh et al (21) found a decrease in debonding and frictional
P1: ARK/MBL/rkc P2: MBL/vks QC: MBL/agr T1: MBL May 16, 1997 13:47 Annual Reviews AR034-16 CERAMIC COMPOSITE INTERFACES 503 the fiber and matrix. To first order εsr = Asr/Rf, where Asr is the amplitude of the surface roughness and Rf is the fiber radius. The amplitude of the surface roughness then appears explicitly in Equation 1, because SRo is directly proportional to the clamping stress σn. Fiber surface roughness can be altered with compliant, low-fracture resistance coatings. A rationale for a coating scheme is best seen in Figure 2, where increasing coating thickness allows for systematic modification of the asperityasperity interactions of fiber and matrix. An increase in the coating thickness reduces the roughness asperity interactions between the fiber and matrix during interfacial sliding. This results in a smaller roughness-induced strain and clamping stress, which reduces the frictional resistance to sliding. A coating with a thickness greater than the amplitude of the fiber surface roughness can completely negate the contribution to the frictional sliding stress. This is best illustrated by the following example. Fiber-sliding measurements were made on a model composite system by Mumm & Faber (16). SiC monofilaments coated with four different thicknesses of carbon (relative to the amplitude of the asperity roughness) in a soda-borosilicate glass matrix were used for model fiber pullout experiments. In the extreme, the coating is meant to completely eliminate asperity contact during debonding and sliding. The fiber force-displacement measurements for the series are shown in Figure 3. First, the controlled thickness coatings induce systematic changes in the load fluctuations; their amplitude decreases with increasing thickness and is essentially eliminated for coatings thicker than the roughness amplitude. The increased slope of the load-deflection curves is likely due to an increase in the coefficient of friction with increasing coating thickness owing to changes in the real area of contact (18). Residual Stress Few ceramic composites are free of residual stresses. The stresses derive from thermal expansion mismatch between fiber and matrix. The role of such stresses is obvious: As clamping stresses increase, the interfacial frictional stress increases. At the extreme, such stresses result in spontaneous cracking of the matrix for a fiber under residual compressive stresses (19). Conversely, spontaneous fiber debonding results when tensile stresses in the fiber exceed a critical value. In addition to interfacial sliding, residual stresses influence the conditions for deflection and/or penetration shown in Figure 1. Compressive residual stresses in the reinforcement enhance conditions for deflection, shifting the contour upward, while tensile stresses enhance penetration (20). Residual stresses can be controlled through the appropriate choice of the fiber-matrix pair. Moreover, the residual stress profile can be altered with fiber loading. Singh et al (21) found a decrease in debonding and frictional
504 FABER Matr coating/fiber contact area Fiber/matrix contact area Thin ce Fibe Thick Coating Fib Figure 2 Schematic illustration of the effect of fiber coatings on fiber sliding behavior. The coatings control the degree of asperity interactions by providing varying amounts of separation of the fiber and matrix phase. As the coating thickness is increased, from top to bottom, the asperity interaction(and the resulting radial misfit strain) is reduced (courtesy of D Mumn
P1: ARK/MBL/rkc P2: MBL/vks QC: MBL/agr T1: MBL May 16, 1997 13:47 Annual Reviews AR034-16 504 FABER Figure 2 Schematic illustration of the effect of fiber coatings on fiber sliding behavior. The coatings control the degree of asperity interactions by providing varying amounts of separation of the fiber and matrix phase. As the coating thickness is increased, from top to bottom, the asperity interaction (and the resulting radial misfit strain) is reduced (courtesy of D Mumm)
CERAMIC COMPOSITE INTERFACES 2 Uncoated Increasing Coatin Thickness 6070 COD (um) Figure 3 Force-displacement curve for pullout testing of four SiC-glass composites with carbon erlayers of increasing thickness(from 16) sliding and attributed it to the reduction in residual stress with fiber loading Alternatively, compliant interphase layers of varying thicknesses also can serve to accomodate large residual thermal mismatch stresses (22, 23) MEASURING INTERFACE PROPERTIES Over the past fifteen years, a variety of interfacial tests for ceramic compos ites either have evolved from evaluation methods used in the polymer-matrix composite field or have developed anew for brittle-brittle composites. Shown in Figure 4 is a sampling of tests for the mechanical evaluation of interfaces (24, 25). Of the tests shown, the bimaterial bend test( Figure 4a), the bimaterial cantilever beam test( Figure 4e ), the single-edge notch beam test(Figure 4), the Brazilian disk test( Figure 4g), the double-cleavage drilled compression tests (Figure 4h), the vickers indentation test(Figure 4)), and Hertzian indentation tests(Figure 4) are occasionally used in screening tests to ascertain whether
P1: ARK/MBL/rkc P2: MBL/vks QC: MBL/agr T1: MBL May 16, 1997 13:47 Annual Reviews AR034-16 CERAMIC COMPOSITE INTERFACES 505 Figure 3 Force-displacement curve for pullout testing of four SiC-glass composites with carbon interlayers of increasing thickness (from 16). sliding and attributed it to the reduction in residual stress with fiber loading. Alternatively, compliant interphase layers of varying thicknesses also can serve to accomodate large residual thermal mismatch stresses (22, 23). MEASURING INTERFACE PROPERTIES Over the past fifteen years, a variety of interfacial tests for ceramic composites either have evolved from evaluation methods used in the polymer-matrix composite field or have developed anew for brittle-brittle composites. Shown in Figure 4 is a sampling of tests for the mechanical evaluation of interfaces (24, 25). Of the tests shown, the bimaterial bend test (Figure 4a), the bimaterial cantilever beam test (Figure 4e), the single-edge notch beam test (Figure 4f ), the Brazilian disk test (Figure 4g), the double-cleavage drilled compression tests (Figure 4h), the Vickers indentation test (Figure 4i), and Hertzian indentation tests (Figure 4j) are occasionally used in screening tests to ascertain whether
FABER 干千7 Figure Schematic of test geometries to measure interfacial mechanical properties crocomposite test, (c) single-fiber double cantilever beam, (n)single- beam, (g) Brazilian age drilled compression test() Vickers inde O Hertzian indentation(after 24, 25) a given materials pair is compatible. In each of these tests, the reinforcement must be of a monolithic form. Consequently, they rarely contain the exact sur- face chemistry, microstructure, or residual stress profile of the true fiber-matrix pair. Therefore, the discussion here is limited to those geometries in which the fiber and matrix can be made identically to those in an actual composite rather than to those geometries that allow materials only similar to the mate- rials used in the composites. The former include push-in, push-through tests
P1: ARK/MBL/rkc P2: MBL/vks QC: MBL/agr T1: MBL May 16, 1997 13:47 Annual Reviews AR034-16 506 FABER Figure 4 Schematic of test geometries to measure interfacial mechanical properties: (a) bimaterial bend test, (b) concentric cylinder tensile test or microcomposite test, (c) single-fiber pullout test, (d ) fiber pullout or push-down test, (e) bimaterial double cantilever beam, (f ) single-edge notched beam, (g) Brazilian disk, (h) double-cleavage drilled compression test (i) Vickers indentation, and ( j) Hertzian indentation (after 24, 25). a given materials pair is compatible. In each of these tests, the reinforcement must be of a monolithic form. Consequently, they rarely contain the exact surface chemistry, microstructure, or residual stress profile of the true fiber-matrix pair. Therefore, the discussion here is limited to those geometries in which the fiber and matrix can be made identically to those in an actual composite, rather than to those geometries that allow materials only similar to the materials used in the composites. The former include push-in, push-through tests
CERAMIC COMPOSITE INTERFACES 507 (Figure 4d) relying on an indenter for loading, pullout tests(Figure 4c), and microcomposite tests(Figure 4b) Indentation Push-in and Push-through Techniques Indentation tests receiving the greatest attention due to their simplicity involve a sharp(26, 27)or blunt(27)indenter that is used to push in a fiber in a composite or push through a fiber in a composite of thin cross-section. First developed by Marshall (26), a sharp indenter was used to displace a fiber into a matrix, and the residual displacement could be ascertained from impressions left in the matrix in he near vicinity of the fiber. Marshall Oliver(28)used a nanoindenter for the same purpose and instrumented the test to provide a continuous measure of the force and displacement during loading, unloading, and load cycling. Analysis of the force-displacement results allowed upper bound estimates of the debond fracture energy and frictional sliding stress, in contrast to the original push-in or push-through test, which was limited to frictional stress evaluation. A further variation of the push-through technique uses a cylindrical indenter that allows no contact with the matrix(29) Elegant analysis of the experiment has been presented by Zhou Mai( 32), who include the radial constraint imposed by neighboring fibers on their analysis of stress transfer and frictional push-out in such a test and have recently included roughness effects(33). Not surprisingly, the frictional push-out stress increases with reinforcement volume fraction. and radial constraints of sur- increase as the embedded length increases More recently, the push-out test has been used for arrays of fibers by Mackin zok(34). protruding fibers, 10 to 15 um in height, trolled etching of the matrix and form the push surface. A displacement piston n the underside of the sample measures displacement. The and the number of fibers displaced Pullout Techniques Conventional pullout tests are prepared with an end of the fiber protruding from the matrix material, which is gripped directly to the loading apparatus (35), a variation of Figure 4c. The free length of the fiber provides a processing challenge, an alignment challenge during mechanical testing, and an enhanced compliance in the system that may prove undesirable for unstable debond crack initiation. The matrix crack is replaced by a matrix surface, which may pos an artificial barrier to debond crack initiation and, hence, unstable debonding Despite this, evaluations by Bright et al (36) showed that sliding resistance measurements on a SiC fiber-reinforced borosilicate glass made via pullout and push-out tests were equivalent. However, constant friction shear stress over the
P1: ARK/MBL/rkc P2: MBL/vks QC: MBL/agr T1: MBL May 16, 1997 13:47 Annual Reviews AR034-16 CERAMIC COMPOSITE INTERFACES 507 (Figure 4d) relying on an indenter for loading, pullout tests (Figure 4c), and microcomposite tests (Figure 4b). Indentation Push-in and Push-through Techniques Indentation tests receiving the greatest attention due to their simplicity involve a sharp (26, 27) or blunt (27) indenter that is used to push in a fiber in a composite or push through a fiber in a composite of thin cross-section. First developed by Marshall (26), a sharp indenter was used to displace a fiber into a matrix, and the residual displacement could be ascertained from impressions left in the matrix in the near vicinity of the fiber. Marshall & Oliver (28) used a nanoindenter for the same purpose and instrumented the test to provide a continuous measure of the force and displacement during loading, unloading, and load cycling. Analysis of the force-displacement results allowed upper bound estimates of the debond fracture energy and frictional sliding stress, in contrast to the original push-in or push-through test, which was limited to frictional stress evaluation. A further variation of the push-through technique uses a cylindrical indenter that allows no contact with the matrix (29). Elegant analysis of the experiment has been presented by Zhou & Mai (30– 32), who include the radial constraint imposed by neighboring fibers on their analysis of stress transfer and frictional push-out in such a test and have recently included roughness effects (33). Not surprisingly, the frictional push-out stress increases with reinforcement volume fraction, and radial constraints of surrounding fibers increase as the embedded length increases. More recently, the push-out test has been used for arrays of fibers by Mackin & Zok (34). Protruding fibers, 10 to 15 µm in height, are created through a controlled etching of the matrix and form the push surface. A displacement piston on the underside of the sample measures displacement. The average interface sliding stress is determined from the applied load, the measured displacement, and the number of fibers displaced. Pullout Techniques Conventional pullout tests are prepared with an end of the fiber protruding from the matrix material, which is gripped directly to the loading apparatus (35), a variation of Figure 4c. The free length of the fiber provides a processing challenge, an alignment challenge during mechanical testing, and an enhanced compliance in the system that may prove undesirable for unstable debond crack initiation. The matrix crack is replaced by a matrix surface, which may pose an artificial barrier to debond crack initiation and, hence, unstable debonding. Despite this, evaluations by Bright et al (36) showed that sliding resistance measurements on a SiC fiber-reinforced borosilicate glass made via pullout and push-out tests were equivalent. However, constant friction shear stress over the
FABER TT Fiber Debonding and Partial Interfacial Sliding Fiber Fracture 20 Release of stored Strain Energy 10 Matrix Fracture Fiber pullout 01020304050 708090100 COD (um) re 5 Continuous load-displacement curve from the modified single-fiber pullout test(from sliding distance leads to underestimates of the friction stress in pullout tests and overestimates in push-out tests due to the poisson contraction and expansion, A modified fiber pullout test has been developed by Mumm Faber(37)that affords the measurement of the continuous force-displacement response for a single crack-bridging fiber shown in Figure 5. A rectangular parallelepiped having a single fiber aligned along the long axis of the beam comprises the test geometry. The specimen is notched at the mid-point to have a greatly reduced cross-section of matrix material through which a crack is propagated, leaving the fiber as a bridging element. The specimen is pulled in tension by bonding the specimen to the load train of a mechanical test system. Crack opening displacement measurements are collected continuously from two linear voltage transducers mounted on opposite sides of the sample, and averaged to eliminate any error due to bending n comparison with earlier interfacial properties tests, the modified fiber pull out test does not rely on an artificial matrix crack(as seen in a conventional
P1: ARK/MBL/rkc P2: MBL/vks QC: MBL/agr T1: MBL May 16, 1997 13:47 Annual Reviews AR034-16 508 FABER Figure 5 Continuous load-displacement curve from the modified single-fiber pullout test (from 36). sliding distance leads to underestimates of the friction stress in pullout tests and overestimates in push-out tests due to the Poisson contraction and expansion, respectively. A modified fiber pullout test has been developed by Mumm & Faber (37) that affords the measurement of the continuous force-displacement response for a single crack-bridging fiber shown in Figure 5. A rectangular parallelepiped having a single fiber aligned along the long axis of the beam comprises the test geometry. The specimen is notched at the mid-point to have a greatly reduced cross-section of matrix material through which a crack is propagated, leaving the fiber as a bridging element. The specimen is pulled in tension by bonding the specimen to the load train of a mechanical test system. Crack opening displacement measurements are collected continuously from two linear voltage transducers mounted on opposite sides of the sample, and averaged to eliminate any error due to bending. In comparison with earlier interfacial properties tests, the modified fiber pullout test does not rely on an artificial matrix crack (as seen in a conventional