T Nozawa et aL/Journal of Nuclear Materials 384(2009)195-211 2. 2. Neutron irradiation on the specimen holder above a narrow groove. The fiber was then randomly selected and monotonically loaded up to the maximum Two types of neutron irradiation campaigns were performed in system allowable load capacity(650 mN). A Berkovich indenter the High Flux Isotope Reactor(HFIR)at Oak Ridge National Labora- tip was used in this experiment. An applied load rate was 0.05 N/ tory. The HFIR-14J fixed-core capsule irradiation was performed in N. S. Details of the fiber push-out test technique have been de- unshielded removable beryllium position of the HFIR. Irradia- scribed elsewhere [18. Microstructures of the F/M interphase tion dose was 7.7 dpa assuming 1.0 x 1025 n/ m2(E>0.1 Mev) and the pushed-out fiber surface were observed by scanning elec corresponds to one displacement per atom(1 dpa), while the irra- tron microscopy(SEM)for both as-received and neutron irradiated diation temperature was 800C. In contrast, small-capsule rabbit materials. irradiations were performed in the hydraulic tube of the hFiR un- Of many stress parameters, two interfacial shear properties: der the Fun and neri SiC/ SiC series of irradiations. The neutron an interfacial debond shear strength (ts)as a critical shear stress to doses were 0.7-4.2 dpa and irradiation temperatures were in the induce an interfacial crack along the bonded F/M interface, and (2) range of 380-1080C Irradiation temperature was measured and an interfacial friction stress(tr) at the debonded interface are de- ontrolled in-situ by thermocouples and capsule sweep-gas mix- ed. Both ts and tr are calculated using experimental push-out tures for the HFIR-14 experiment. The uncertainty of irradiation test results of (1)a debond initiation stress(od)and (2)a complete emperatures was reasonably low(+20C). In contrast, irradiation debonding and sliding stress(omax) as defined in Fig. 2. In the fol- mperatures in rabbit irradiation were estimated by the isochro- lowing discussion, a compressive stress is denoted with a negative lI annealing of CVD-SiC temperature monitors. In principle, the sign For evaluation of this type of data, various analytical models temperature monitor gives the temperature near the end of the have been developed [18-24. Of these models, a non-linear shear- irradiation period with a similar uncertainty to the instrument lag model proposed by Hsueh [ 21] becomes a basis to evaluate ts experiment(±20°C)[15,16 considering the precise stress interaction at the F/M interface: (1) the radial dependence of the axial stresses in both the fiber 23. Tensile test and the matrix, (2) the shear stress distribution in the matrix and(3)the exact equilibrium equation in relating the tangential Following a guideline of ASTM C1275-00, cyclic unloading/ stress to the radial stress at the interface Hsueh further developed reloading tensile tests were conducted at ambient-temperature a double shear-lag model by separately idering the stress using an electromechanical testing machine. Rectangular minia- interactions among the fiber, matrix and F/M interphase [22]. or 20 x 2Wx 15mm were prepared. Note that the total speci- cability to isotropic constituents. In the modified double shear-lag men length of 50 mm was fixed for all specimens. Specimen size model by the authors(see Appendix), anisotropy of the constitu- effect is a potential concern when testing these specimens. Consid- ents is considered. Besides, residual stresses induced by thermal ering the structural minimum unit width(or thickness)of the uni- expansion mismatch and by differential swelling among the fiber. directional composites is larger than the width(or thickness)of the F/M interphase, and matrix can be discussed together in the model, mono fiber bundle(<l mm), a very minor effect of specimen size although the effect of dynamic phenomena such as irradiation on tensile properties, which depend on the axial fiber volume frac- creep of Sic and Py c was ignored for simplicity. In contrast, Shetty tion, is expected in the size range of concern [17 Indeed, no signif- [24] originally discussed the interfacial friction issue for the deb icant size effect was found in this study. The tensile specimen was onded interface. In the previous work by the authors [25 ]. Shetty passively gripped via aluminum grip-end tabs using a pneumatic model was updated for anisotropic composites and this method wedge-type gripping device Tensile strain was measured by a pair was applied in this study. f strain gauges with a gauge length of 5.0 mm, which were adhe- The detailed analogy of the modified Hsueh model was dis- sively bonded on specimen surfaces of the middle gauge section. a cussed elsewhere [ 22, 25]. The resultant form can provide an inter- onstant crosshead displacement rate was 0.5 mm/min th(ts) -it)-z(od+Oth+Orr ) exp(it)+exp(-it)-2 exp (it)-exp(-it) Youngs modulus was defined as an initial tangent modulus in where the constants i and Z are determined by dimensions and the tensile stress-strain curve. Proportional limit tensile stress elastic constants of the constituents, and th and irr are stress (PLS)was determined as a stress of 5% deviation in stress from ini- parameters defined by the thermal expansion mismatch and the tial linearity following ASTM C1275-00. Ultimate tensile strength differential swelling among the fiber, matrix and F/M interphase. (UTS)and total elongation were defined as a stress and a strain respectively(see Appendix). Assuming no contribution from the thermal residual stress(or irradiation-induced stress), oth (or girr becomes 2. 4. Single fiber push-out test Assuming a coulomb friction, an interfacial friction stress (tr) an be determined as [25]: Single fiber push-out tests were conducted at room- ture using a nano-indentation test system. A thin-strip 可=-(++m) vith a thickness of 30-220 um was cut from the tensile n where, u is a coefficient of friction and oth. rough and ormad are radial with both surfaces polished by the standard metallographic tech- clamping stresses induced by thermal expansion mismatch, fiber niques to a surface finish of +0.5 um. The specimen was bonded surface roughness and differential swelling. respectively.These2.2. Neutron irradiation Two types of neutron irradiation campaigns were performed in the High Flux Isotope Reactor (HFIR) at Oak Ridge National Labora￾tory. The HFIR-14 J fixed-core capsule irradiation was performed in an unshielded removable beryllium position of the HFIR. Irradia￾tion dose was 7.7 dpa assuming 1.0 1025 n/m2 (E > 0.1 MeV) corresponds to one displacement per atom (1 dpa), while the irra￾diation temperature was 800 C. In contrast, small-capsule rabbit irradiations were performed in the hydraulic tube of the HFIR un￾der the FUN and NERI SiC/SiC series of irradiations. The neutron doses were 0.7–4.2 dpa and irradiation temperatures were in the range of 380–1080 C. Irradiation temperature was measured and controlled in-situ by thermocouples and capsule sweep–gas mix￾tures for the HFIR-14 experiment. The uncertainty of irradiation temperatures was reasonably low (±20 C). In contrast, irradiation temperatures in rabbit irradiation were estimated by the isochro￾nal annealing of CVD-SiC temperature monitors. In principle, the temperature monitor gives the temperature near the end of the irradiation period with a similar uncertainty to the instrument experiment (±20 C) [15,16]. 2.3. Tensile test Following a guideline of ASTM C1275-00, cyclic unloading/ reloading tensile tests were conducted at ambient-temperature using an electromechanical testing machine. Rectangular minia￾ture tensile specimens with a gauge size of either 20L 4W 1.5T or 20L 2W 1.5T mm3 were prepared. Note that the total speci￾men length of 50 mm was fixed for all specimens. Specimen size effect is a potential concern when testing these specimens. Consid￾ering the structural minimum unit width (or thickness) of the uni￾directional composites is larger than the width (or thickness) of the mono fiber bundle (<1 mm), a very minor effect of specimen size on tensile properties, which depend on the axial fiber volume frac￾tion, is expected in the size range of concern [17]. Indeed, no signif￾icant size effect was found in this study. The tensile specimen was passively gripped via aluminum grip-end tabs using a pneumatic wedge-type gripping device. Tensile strain was measured by a pair of strain gauges with a gauge length of 5.0 mm, which were adhe￾sively bonded on specimen surfaces of the middle gauge section. A constant crosshead displacement rate was 0.5 mm/min. Young’s modulus was defined as an initial tangent modulus in the tensile stress–strain curve. Proportional limit tensile stress (PLS) was determined as a stress of 5% deviation in stress from ini￾tial linearity following ASTM C1275-00. Ultimate tensile strength (UTS) and total elongation were defined as a stress and a strain at composite fracture, respectively. 2.4. Single fiber push-out test Single fiber push-out tests were conducted at room-tempera￾ture using a nano-indentation test system. A thin-strip specimen with a thickness of 30–220 lm was cut from the tensile specimen with both surfaces polished by the standard metallographic tech￾niques to a surface finish of 0.5 lm. The specimen was bonded on the specimen holder above a narrow groove. The fiber was then randomly selected and monotonically loaded up to the maximum system allowable load capacity (650 mN). A Berkovich indenter tip was used in this experiment. An applied load rate was 0.05 N/ N  s. Details of the fiber push-out test technique have been de￾scribed elsewhere [18]. Microstructures of the F/M interphase and the pushed-out fiber surface were observed by scanning elec￾tron microscopy (SEM) for both as-received and neutron irradiated materials. Of many stress parameters, two interfacial shear properties: (1) an interfacial debond shear strength (ss) as a critical shear stress to induce an interfacial crack along the bonded F/M interface, and (2) an interfacial friction stress (sf) at the debonded interface are de- fined. Both ss and sf are calculated using experimental push-out test results of (1) a debond initiation stress (rd) and (2) a complete debonding and sliding stress (rmax) as defined in Fig. 2. In the fol￾lowing discussion, a compressive stress is denoted with a negative sign. For evaluation of this type of data, various analytical models have been developed [18–24]. Of these models, a non-linear shear￾lag model proposed by Hsueh [21] becomes a basis to evaluate ss considering the precise stress interaction at the F/M interface: (1) the radial dependence of the axial stresses in both the fiber and the matrix, (2) the shear stress distribution in the matrix and (3) the exact equilibrium equation in relating the tangential stress to the radial stress at the interface. Hsueh further developed a double shear-lag model by separately considering the stress interactions among the fiber, matrix and F/M interphase [22]. One major drawback of the original Hsueh models is limited appli￾cability to isotropic constituents. In the modified double shear-lag model by the authors (see Appendix), anisotropy of the constitu￾ents is considered. Besides, residual stresses induced by thermal expansion mismatch and by differential swelling among the fiber, F/M interphase, and matrix can be discussed together in the model, although the effect of dynamic phenomena such as irradiation creep of SiC and PyC was ignored for simplicity. In contrast, Shetty [24] originally discussed the interfacial friction issue for the deb￾onded interface. In the previous work by the authors [25], Shetty’s model was updated for anisotropic composites and this method was applied in this study. The detailed analogy of the modified Hsueh model was dis￾cussed elsewhere [22,25]. The resultant form can provide an inter￾facial debond shear strength (ss) as where the constants k and Z are determined by dimensions and elastic constants of the constituents, and rth and rirr are stress parameters defined by the thermal expansion mismatch and the differential swelling among the fiber, matrix and F/M interphase, respectively (see Appendix). Assuming no contribution from the thermal residual stress (or irradiation-induced stress), rth (or rirr) becomes zero. Assuming a coulomb friction, an interfacial friction stress (sf) can be determined as [25]: sf ¼ l rth r þ rrough r þ rirrad r ; ð2Þ where, l is a coefficient of friction and rth r , rrough r and rirrad r are radial clamping stresses induced by thermal expansion mismatch, fiber surface roughness and differential swelling, respectively. These ss ¼  rfk 2 rd½expðktÞ þ expðktÞ  Zðrd þ rth þ rirrÞ½expðktÞ þ expðktÞ  2 expðktÞ  expðktÞ  ; ð1Þ T. Nozawa et al. / Journal of Nuclear Materials 384 (2009) 195–211 197