《复合材料 Composites》课程教学资源（学习资料）第五章 陶瓷基复合材料_creeep of nextel-LaPO4-ALUMINA

Availableonlineatwww.sciencedirect.com DIRECT. COMPOSITES CIENCE AND TECHNOLOGY ELSEVIER Composites Science and Technology 66(2006)2089-2099 www.elsevier.com/locate/compscitech Creep behavior of NextelM610/Monazite/ Alumina composite at elevated temperatures M B. Ruggles-Wrenn ,S.s. Musil, s Mall, K.A. Keller Department of Aeronautics and Astronautics, Air Force Institute of Technology, Wright-Patterson Air Force Base, OH 45433-7765, USA Received 7 June 2005: received in revised form 8 December 2005: accepted 22 December 2005 Available online 6 March 2006 Abstract The creep behavior of a N610/LaPO/Al,O3 composite was evaluated in this work. The composite consists of a porous alumina matrix reinforced with Nextel 610 fibers coated with monazite in a symmetric cross-ply(0%/90/0 90)s orientation. The te stress-strain behavior was investigated and the tensile properties measured at room temperature and in the 900-1200oC range the a tion of monazite coating resulted in 50% improvement in ultimate tensile strength (UTS) at temperatures <1100C, and in improvement in UTS at 1200C. Tensile creep behavior was examined at temperatures in the 900-1100C range for creep stresses rang- ing from 40 to 150 MPa. Primary and secondary creep s were observed in creep tests at 900C. At temperatures above 900C, the composite exhibited primary, secondary and tertiary creep. Minimum creep rate was reached in all tests. Creep rates accelerated with increasing temperature and creep stress. At 900C creep run-out, defined as 100 h at creep stress, was achieved for stress levels <120 MPa. The residual strength and modulus of all specimens that achieved run-out were characterized. Comparison with results obtained for N610/Al2O3(control) specimens revealed that the use of the monazite coating resulted in improved creep resistance at 900C Creep performance deteriorated rapidly as temperatures increased above 900 C. Composite microstructure, as well as damage and failure mechanisms were investigated e 2006 Elsevier Ltd. All rights reserved Keywords: A. Ceramic-matrix composites(CMCs): A Oxides; A. Fibres; A. Coatings: B Creep; B. High-temperature properties; B. Mechanical prop- erties: D. Fractography 1. Introduction and fracture toughness at high temperatures and they therefore continue to attract attention as candidate materi- The severe conditions encountered by aerospace compo als for such applications. Additionally, the lower densities nents require structural materials that have superior long- of CMCs and their higher use temperatures, along with a term mechanical properties and retained properties under reduced need for cooling air, allow for improved high-tem high temperature, high pressure, and varying environmen- perature performance when compared to conventional tal factors, such as moisture [1]. Ceramic-matrix compos- nickel-based superalloys [2]. Alternately, it is recognized ites(CMCs) are capable of maintaining excellent strength that the thermodynamic stability and oxidation resistance of CMCs are vital issues The views expressed are those of the authors and do not reflect the Non-oxide fiber/non-oxide matrix CMCs such as SiC/ official policy or position of the United States Air Force, Department of SiC generally show poor oxidation resistance [3, 4, partic Defense or the US government larly at intermediate temperatures(800C). The degrad Corresponding author. Tel: +1937 255 3636x4641: fax: +1937 656 tion involves the oxidation of the fibers fiber coatings and 4032 aruggles -wrenn(@afit. edu (M. B. Ruggles. matrices and is typically accelerated by the presence of moisture [5-11]. Using a non-oxide fiber/oxide matrix or I Under USAF Contract F33615-01-C-5214 an oxide fiber/non-oxide matrix composite generally does 02663538/S. see front matter 2006 Elsevier Ltd. All rights reserved doi:10.1016j.compscitech.2005.12.026

Creep behavior of NextelTM610/Monazite/Alumina composite at elevated temperatures q M.B. Ruggles-Wrenn a,*, S.S. Musil a , S. Mall a , K.A. Keller b,1 a Department of Aeronautics and Astronautics, Air Force Institute of Technology, Wright-Patterson Air Force Base, OH 45433-7765, USA b UES, Inc., 4401 Dayton Xenia Road, Dayton, OH 45433-7817, USA Received 7 June 2005; received in revised form 8 December 2005; accepted 22 December 2005 Available online 6 March 2006 Abstract The creep behavior of a N610/LaPO4/Al2O3 composite was evaluated in this work. The composite consists of a porous alumina matrix reinforced with Nextel 610 fibers coated with monazite in a symmetric cross-ply (0/90/0/90)S orientation. The tensile stress–strain behavior was investigated and the tensile properties measured at room temperature and in the 900–1200C range. The addi￾tion of monazite coating resulted in 50% improvement in ultimate tensile strength (UTS) at temperatures 61100 C, and in 37% improvement in UTS at 1200 C. Tensile creep behavior was examined at temperatures in the 900–1100 C range for creep stresses rang￾ing from 40 to 150 MPa. Primary and secondary creep regimes were observed in creep tests at 900 C. At temperatures above 900 C, the composite exhibited primary, secondary and tertiary creep. Minimum creep rate was reached in all tests. Creep rates accelerated with increasing temperature and creep stress. At 900 C creep run-out, defined as 100 h at creep stress, was achieved for stress levels 6120 MPa. The residual strength and modulus of all specimens that achieved run-out were characterized. Comparison with results obtained for N610/Al2O3 (control) specimens revealed that the use of the monazite coating resulted in improved creep resistance at 900 C. Creep performance deteriorated rapidly as temperatures increased above 900 C. Composite microstructure, as well as damage and failure mechanisms were investigated. 2006 Elsevier Ltd. All rights reserved. Keywords: A. Ceramic-matrix composites (CMCs); A. Oxides; A. Fibres; A. Coatings; B. Creep; B. High-temperature properties; B. Mechanical prop￾erties; D. Fractography 1. Introduction The severe conditions encountered by aerospace compo￾nents require structural materials that have superior long￾term mechanical properties and retained properties under high temperature, high pressure, and varying environmen￾tal factors, such as moisture [1]. Ceramic–matrix compos￾ites (CMCs) are capable of maintaining excellent strength and fracture toughness at high temperatures and they therefore continue to attract attention as candidate materi￾als for such applications. Additionally, the lower densities of CMCs and their higher use temperatures, along with a reduced need for cooling air, allow for improved high-tem￾perature performance when compared to conventional nickel-based superalloys [2]. Alternately, it is recognized that the thermodynamic stability and oxidation resistance of CMCs are vital issues. Non-oxide fiber/non-oxide matrix CMCs such as SiC/ SiC generally show poor oxidation resistance [3,4], particu￾larly at intermediate temperatures (800 C). The degrada￾tion involves the oxidation of the fibers, fiber coatings, and matrices and is typically accelerated by the presence of moisture [5–11]. Using a non-oxide fiber/oxide matrix or an oxide fiber/non-oxide matrix composite generally does 0266-3538/$ - see front matter 2006 Elsevier Ltd. All rights reserved. doi:10.1016/j.compscitech.2005.12.026 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 4032. E-mail address: marina.ruggles-wrenn@afit.edu (M.B. Ruggles￾Wrenn). 1 Under USAF Contract # F33615-01-C-5214. www.elsevier.com/locate/compscitech Composites Science and Technology 66 (2006) 2089–2099 COMPOSITES SCIENCE AND TECHNOLOGY

M B. Ruggles-Wrenn et al /Composites Science and Technology 66(2006 )2089-2099 not substantially improve the high temperature oxidation objective of this study was to investigate the effectiveness resistance[12]. The need for environmentally stable com- of using monazite coatings in a crossply NextelM610/alu posites led to the development of CMCs based on oxide mina composite to improve the creep resistance of the constituents [13-21 CMC. Keller et al. [39] investigated the effectiveness of It is widely accepted that in order to avoid brittle frac- monazite coatings in Nextel 610/alumina composites after ture behavior in CMCs and improve the damage tolerance, long-term exposure at 1100 and 1200C Coated fiber sam- a weak fiber/matrix interface is needed, which serves to ples exhibited better tensile strength retention after 1000 h deflect matrix cracks and to allow subsequent fiber pull- at 1200C when compared to the control(uncoated fiber out [22-25]. A matrix consisting of finely distributed poros- material. This effort investigates the creep-rupture behavior o y can provide a similar crack-deflecting behavior instead of the Nextel 610/Monazite/Alumina(N610/M/A)com- of a separate interface between matrix and fibers [26]. In posite developed at the Air Force Research Laboratory this case, there is strong bonding between the fiber and(AFRL/MLLN), Materials and Manufacturing Director matrix; consequently, there is a minimum matrix porosity ate. Creep-rupture tests were conducted at temperatures needed for this concept to work [27]. For a dense (90%) in the 900-1 100C range for creep stresses ranging from matrix composite, however, an interfacial coating is needed 40 to 150 MPa Resulting creep performance imposes lim oxide coatings for oxide and non-oxide composites has applications. The composite microstructure, as wlle for crack deflection. An extensive review of the research on itations on the use of this material in high-temperature been given by Kerans et al. [28]. damage and failure mechanisms are discussed The development of oxide-oxide composites that rely on a weak fiber/matrix interface for crack deflection prompte 2. Material and experimental arrangements research into oxidation resistant fiber coatings that are hemically stable with the composite constituents. Various Nextel 610 fiber tows were desized in air at 1100C and oxidation-resistant coating materials have been investi- then coated with a monazite precursor solution at 1100oC gated, including monazite(LaPO4) and scheelite Of these, As described elsewhere [35] this monazite sol was washed a significant amount of work has been completed on com- multiple times to remove any residual ions, which have been posites containing monazite coatings. The initial works of associated with fiber strength loss. After coating, the fibers Morgan and Marshall [20, 29, 30] showed the potential of were filament wound using an alumina slurry consisting of a monazite as a weak interface material, particularly for alu- polymeric alumina sol and alumina powder (AKP-53 mina-based composites, due to the chemical compatibility Sumitomo Corp. ) In addition, uncoated fibers were wound of monazite with alumina at high temperature. Since then, for control samples. As described by Keller et al. [39], dur- numerous studies have examined the production of mona- ing the winding process, the number of wheel revolutions zite coatings and its use with different fiber/matrix combi- was recorded to determine the number of fiber tows within nations [31-35]. Fiber strength degradation caused by the a given tape width. This value was then used to calculate the coating and long-term, high-temperature heat treatments volume percentage of fibers in the composite. The resultant was identified as the key problem with monazite coatings unidirectional tapes were cut from the spool, stacked into a Boakye et al. [35] investigated the effects of different metal mold in an 8-layer symmetric cross-ply orientation of liquid precursors on coating characteristics and tensile [(0%/90%)]s, and warm vacuum-molded to form green cross- strength of coated fibers, and developed monazite coating ply composites. After drying, the billets were sintered at that did not cause significant loss of fiber strength 1200C for 5 h in air. During the heating cycle, a one hour Porous-matrix oxide/oxide CMCs exhibit several behav- hold at low temperature was performed to remove any or trends that are distinctly different from those exhibited residual organics in the matrix. Control billets, N610/Alu by traditional CMCs with a fiber-matrix interface. Most mina(N610/A), containing uncoated fibers were produced Sic-fiber-containing CMCs exhibit longer life under static with the same procedure. Billet properties, namely fiber loading and shorter life under cyclic loading [36]. For these volume fraction and density, are summarized in Table I materials, fatigue is significantly more damaging than Micrographs of the as-processed material shown in Fig. I creep. Zawada et al. [37] examined the high-temperature reveal shrinkage and sintering cracks that occurred during mechanical behavior of a porous matrix Nextel 610/alumi- the cooling stage of the composite processing. Extensive nosilicate composite Results revealed excellent fatigue per- surface microcracking is seen in Fig. 1(a), while Fig. I(b) formance at 1000C, the material exhibited high fatigue shows interlaminar matrix cracks. limit, long fatigue life and near 100% strength retention The N610/M/A specimens were cut from seven different Conversely, creep lives were short, indicating low creep billets, and N610/A specimens from three different billets. resistance and limiting the use of that CMC to tempera- Specimen numbers contain reference to the billet number. tures below 1000 For example, number Bl-I refers to the specimen I from Because creep is shown to be considerably more damag- billet 1. Billets were cut into flat rectangular coupons, which ing to porous-matrix oxide/oxide CMCs [37,38) creep test- were machined into dog bone-shaped specimens shown in ing is well suited for assessing the long term durability an nd Fig. 2. Diamond-grit grinding was used for billets Bl-B8 gh-temperature erformance of these materials. The and the abrasive water-jet machine, for billets B9-Bll

not substantially improve the high temperature oxidation resistance [12]. The need for environmentally stable com￾posites led to the development of CMCs based on oxide constituents [13–21]. It is widely accepted that in order to avoid brittle frac￾ture 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 pull￾out [22–25]. A matrix consisting of finely distributed poros￾ity can provide a similar crack-deflecting behavior instead of a separate interface between matrix and fibers [26]. In this case, there is strong bonding between the fiber and matrix; consequently, there is a minimum matrix porosity needed for this concept to work [27]. For a dense (>90%) matrix composite, however, an interfacial coating is needed for crack deflection. An extensive review of the research on oxide coatings for oxide and non-oxide composites has been given by Kerans et al. [28]. The development of oxide–oxide composites that rely on a weak fiber/matrix interface for crack deflection prompted research into oxidation resistant fiber coatings that are chemically stable with the composite constituents. Various oxidation-resistant coating materials have been investi￾gated, including monazite (LaPO4) and scheelite. Of these, a significant amount of work has been completed on com￾posites containing monazite coatings. The initial works of Morgan and Marshall [20,29,30] showed the potential of monazite as a weak interface material, particularly for alu￾mina-based composites, due to the chemical compatibility of monazite with alumina at high temperature. Since then, numerous studies have examined the production of mona￾zite coatings and its use with different fiber/matrix combi￾nations [31–35]. Fiber strength degradation caused by the coating and long-term, high-temperature heat treatments was identified as the key problem with monazite coatings [34]. Boakye et al. [35] investigated the effects of different liquid precursors on coating characteristics and tensile strength of coated fibers, and developed monazite coating that did not cause significant loss of fiber strength. Porous-matrix oxide/oxide CMCs exhibit several behav￾ior trends that are distinctly different from those exhibited by traditional CMCs with a fiber–matrix interface. Most SiC-fiber-containing CMCs exhibit longer life under static loading and shorter life under cyclic loading [36]. For these materials, fatigue is significantly more damaging than creep. Zawada et al. [37] examined the high-temperature mechanical behavior of a porous matrix Nextel 610/alumi￾nosilicate composite. Results revealed excellent fatigue per￾formance at 1000 C, the material exhibited high fatigue limit, long fatigue life and near 100% strength retention. Conversely, creep lives were short, indicating low creep resistance and limiting the use of that CMC to tempera￾tures below 1000 C. Because creep is shown to be considerably more damag￾ing to porous-matrix oxide/oxide CMCs [37,38], creep test￾ing is well suited for assessing the long term durability and high-temperature performance of these materials. The objective of this study was to investigate the effectiveness of using monazite coatings in a crossply NextelTM610/alu￾mina composite to improve the creep resistance of the CMC. Keller et al. [39] investigated the effectiveness of monazite coatings in Nextel 610/alumina composites after long-term exposure at 1100 and 1200 C. Coated fiber sam￾ples exhibited better tensile strength retention after 1000 h at 1200 C when compared to the control (uncoated fiber) material. This effort investigates the creep-rupture behavior of the Nextel 610/Monazite/Alumina (N610/M/A) com￾posite developed at the Air Force Research Laboratory (AFRL/MLLN), Materials and Manufacturing Director￾ate. Creep-rupture tests were conducted at temperatures in the 900–1100 C range for creep stresses ranging from 40 to 150 MPa. Resulting creep performance imposes lim￾itations on the use of this material in high-temperature applications. The composite microstructure, as well as damage and failure mechanisms are discussed. 2. Material and experimental arrangements Nextel 610 fiber tows were desized in air at 1100 C and then coated with a monazite precursor solution at 1100 C. As described elsewhere [35], this monazite sol was washed multiple times to remove any residual ions, which have been associated with fiber strength loss. After coating, the fibers were filament wound using an alumina slurry consisting of a polymeric alumina sol and alumina powder (AKP-53, Sumitomo Corp.). In addition, uncoated fibers were wound for control samples. As described by Keller et al. [39], dur￾ing the winding process, the number of wheel revolutions was recorded to determine the number of fiber tows within a given tape width. This value was then used to calculate the volume percentage of fibers in the composite. The resultant unidirectional tapes were cut from the spool, stacked into a metal mold in an 8-layer symmetric cross-ply orientation of [(0/90)]2S, and warm vacuum-molded to form green cross￾ply composites. After drying, the billets were sintered at 1200 C for 5 h in air. During the heating cycle, a one hour hold at low temperature was performed to remove any residual organics in the matrix. Control billets, N610/Alu￾mina (N610/A), containing uncoated fibers were produced with the same procedure. Billet properties, namely fiber volume fraction and density, are summarized in Table 1. Micrographs of the as-processed material shown in Fig. 1 reveal shrinkage and sintering cracks that occurred during the cooling stage of the composite processing. Extensive surface microcracking is seen in Fig. 1(a), while Fig. 1(b) shows interlaminar matrix cracks. The N610/M/A specimens were cut from seven different billets, and N610/A specimens from three different billets. Specimen numbers contain reference to the billet number. For example, number B1-1 refers to the specimen 1 from billet 1. Billets were cut into flat rectangular coupons, which were machined into dog bone-shaped specimens shown in Fig. 2. Diamond-grit grinding was used for billets B1–B8, and the abrasive water-jet machine, for billets B9–B11. 2090 M.B. Ruggles-Wrenn et al. / Composites Science and Technology 66 (2006) 2089–2099

M B. Ruggles-Wrenn et al/ Composites Science and Technology 66(2006)2089-2099 Table I 10.0 ummary of billet properties for the N610/M/A and the N610/A 36.0 fraction (a) Density (g/cc) MonazitelAlumina co …--1 B2 63.0 44. 9088 Fig. 2. Test specimen, dimensions in mm. B10 ment rate of 0.05 mm/s at 23. 900, 1000, 1100 and 1200oC /610/Almina composi Creep-rupture tests were conducted in load control in 595 accordance with Astm standard c 1337 at 900 and 1100C. Specimens were loaded to the creep stress level at the rate of 20 MPa/s. In each test, stress-strain data were recorded during the loading to the creep stress level and the A servocontrolled MTS mechanical testing actual creep period. Thus both total strain and creep strain equipped with hydraulic water-cooled collet grips could be calculated and examined. Creep run-out was pact single-zone resistance-heated furnace, and a defined as 100 h at a given creep stress. To determine the ture controller were used in all tests. An MTS TestStar II retained strength and modulus, specimens that achieved digital controller was employed for input signal generation run-out were subjected to tensile tests to failure at the tem- and data acquisition. Strain measurement was accom- perature of the creep test. One specimen was tested per test plished with an MTS high-temperature air-cooled uniaxial condition. The authors recognize that this is a limited set of extensometer. For elevated temperature testing, thermo- data. However, this scoping research serves to identify the couples were bonded to the specimens using alumina temperature range where the use of monazite coating cement( Zircar)to calibrate the furnace on a periodic basis. results in improved creep resistance. Furthermore, results The furnace controller(using a non-contacting thermocou- of this exploratory effort can be used to determine whether ple exposed to the ambient environment near the test spec- a more rigorous investigation of the effectiveness of mona imen) was adjusted to determine the power setting needed zite coating in this CMC or in a different material system to achieve the desired temperature of the test specimen. should be undertaken The determined power setting was then used in actual tests Fracture surfaces of failed specimens were gold coated 4. Results and discussion and examined using SEM(Model 360FE, Leica)as well as optical microscopy. 4. Monotonic tension 3. Test procedures The N610/M/A specimens were tested at 900, 1000,1100 and 1200C, and the N610/A specimens, at 23, 1100 and All tests were conducted in laboratory air environment. 1200C. Tensile test results are summarized in Table 2 In elevated temperature tests, a specimen was heated to the where elastic modulus, ultimate tensile strength (UTS) test temperature at a rate of lC/s, and held at tempera- and failure strain are presented for each test temperature. ture for additional 15 min prior to testing. Tensile tests The tensile stress-strain curves for both N610/A and were performed in stroke control with a constant displace- N610/M/A composites are shown in Fig 3 10 mm Fig. I. Micrographs of the as-processed material showing shrinkage cracks: (a) extensive surface microcracking and (b) interlaminar matrix cracks

A servocontrolled MTS mechanical testing machine equipped with hydraulic water-cooled collet grips, a com￾pact single-zone resistance-heated furnace, and a tempera￾ture controller were used in all tests. An MTS TestStar II digital controller was employed for input signal generation and data acquisition. Strain measurement was accom￾plished with an MTS high-temperature air-cooled uniaxial extensometer. For elevated temperature testing, thermo￾couples were bonded to the specimens using alumina cement (Zircar) to calibrate the furnace on a periodic basis. The furnace controller (using a non-contacting thermocou￾ple exposed to the ambient environment near the test spec￾imen) was adjusted to determine the power setting needed to achieve the desired temperature of the test specimen. The determined power setting was then used in actual tests. Fracture surfaces of failed specimens were gold coated and examined using SEM (Model 360FE, Leica) as well as optical microscopy. 3. Test procedures All tests were conducted in laboratory air environment. In elevated temperature tests, a specimen was heated to the test temperature at a rate of 1 C/s, and held at tempera￾ture for additional 15 min prior to testing. Tensile tests were performed in stroke control with a constant displace￾ment rate of 0.05 mm/s at 23, 900, 1000, 1100 and 1200 C. Creep-rupture tests were conducted in load control in accordance with ASTM standard C 1337 at 900 and 1100 C. Specimens were loaded to the creep stress level at the rate of 20 MPa/s. In each test, stress–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. Creep run-out was defined as 100 h at a given creep stress. To determine the retained strength and modulus, specimens that achieved run-out were subjected to tensile tests to failure at the tem￾perature of the creep test. One specimen was tested per test condition. The authors recognize that this is a limited set of data. However, this scoping research serves to identify the temperature range where the use of monazite coating results in improved creep resistance. Furthermore, results of this exploratory effort can be used to determine whether a more rigorous investigation of the effectiveness of mona￾zite coating in this CMC or in a different material system should be undertaken. 4. Results and discussion 4.1. Monotonic tension The N610/M/A specimens were tested at 900, 1000, 1100 and 1200 C, and the N610/A specimens, at 23, 1100 and 1200 C. Tensile test results are summarized in Table 2, where elastic modulus, ultimate tensile strength (UTS), and failure strain are presented for each test temperature. The tensile stress–strain curves for both N610/A and N610/M/A composites are shown in Fig. 3. Fig. 1. Micrographs of the as-processed material showing shrinkage cracks: (a) extensive surface microcracking and (b) interlaminar matrix cracks. Table 1 Summary of billet properties for the N610/M/A and the N610/A composites Billet Fiber volume fraction (%) Density (g/cc) N610/Monazite/Alumina composite B1 44.3 2.55 B2 39.7 2.48 B4 40.2 2.49 B5 43.0 2.50 B9 44.7 2.58 B10 42.9 2.58 B11 41.0 2.53 N610/Alumina composite B3 51.6 2.95 B6 54.8 2.99 B8 46.7 2.85 R=50 36.0 63.0 8.0 10.0 5.0 Fig. 2. Test specimen, dimensions in mm. M.B. Ruggles-Wrenn et al. / Composites Science and Technology 66 (2006) 2089–2099 2091

M.B. Ruggles-Wrenn et al /Composites Science and Technology 66(2006)2089-2099 failure strain at 1200C is approximately 4.5 times that at ummary of tensile properties for the N610/M/A and the N6IO/A 1100 %( composites It is important to note that in all tension tests, as well as Specimen Temperature Elastic modulus UTS Failure strain in all other tests reported herein, the failure occurred within GPa) the gage section of the extensometer N10/ Monazite/Alumina composite 0.3l 4. 2. Creep rupture B2-1 15 1200 130 142 Results of the creep-rupture tests are summarized in N610/Almina composite Table 3, where creep strain accumulation and time to rup l17 0.09 ture are shown for each test temperature and applied stress 105 level. Creep curves obtained for the coated fiber composite at 1100C are presented in Fig. 4. Creep curves obtained at 900C for N610/M/A and N610/A are shown in Figs. 5(a) The stress-strain curves obtained for the uncoated fiber and(b), respectively. The time scale in Figs. 4(a)and 5 is composite at 23 and 1100C are nearly linear to failure. reduced in order to clearly show the creep curves produced all suggest a monolithic fracture behavior with no toughen- 900 C the uncoated fiber composite did not achieve a ing from fibers. For temperatures <1100C and strains run-out even for a low stress of 73 MPa. Recognizing that <0.1%, the stress-strain behavior of the N610/M/A com- increase in temperature would result in reduced creep life, posite is also nearly linear elastic. However, as the strain the N610/A was not expected to achieve a run-out at exceeds 0. 1% the stress-strain curves depart from linearity. 1100 oC for stresses >75 MPa. and was therefore not The subsequent non-linear behavior, indicative of progres- tested at that temperature sive matrix cracking and crack deflection, continues until Creep curves produced by N610/M/A at 1100C exhibit failure. The fracture surfaces show considerable fiber pull primary, secondary and tertiary creep regimes. Transition decrease with increasing temperature, while the failure ately. For stresses <80 MPa, secondary creep persists for strain remains fairly temperature-independent. Results in n70% of the creep life before transitioning to tertiary Table 2 reveal that at 1100C, the addition of the monazite creep. However, for stresses >100 MPa, secondary creep coating results in a nearly 50% increase in UTS. However, transitions to tertiary creep during the first third of the this increase in strength is accompanied by a 34% decrease creep life. Creep strain accumulation decreases witl in modulus and a threefold increase in failure strain increasing applied stress. While at 40 MPa creep strain The stress-strain behavior of both composites changes a significant 7.66%, at 120 MPa creep strain is only 0.7% It is noteworthy that at 1100C, for stress levels in the gely nonlinear, the strength and modulus decrease signif 40-120 MPa range, the accumulated creep strains signifi- cantly,while failure strains increase. The increase in cantly exceed the failure strain obtained in the tension test temperature from 1100 to 1200C causes a 10% loss in The creep run-out, defined as survival of 100 h at a given UTS and a 57% loss in modulus for the uncoated fiber composite, and a 17 loss in UTS and a 34% loss in mod. creep stress, was not achieved ulus for the coated fiber composite. For both materials the Table 3 Summary of creep-rupture results for the N610/M/A and the N610/A composites Specimen Temperature stress Creep strain Time to 900c rupture(s Ng10/Monazite/Alumina cor 15023c 1100°c 7.66 1200°c B4-2 1100c 1.58 B4-3 l100 1200°c B9. 0.05 63.060 Bl0-1 0.04 522,3652 Bl0-2 432,1752 N610/Monazite/ Alumina B51 0.05 BIl-I l40 0.03 BIl-2 150 0.75 STRAIN (% N610/Almina composite B6-1 Fig. 3. Tensile stress-strain curves for N610/A and N610/M/A ceram B6-2 0.03 19,995 composites at various temperatures

The stress–strain curves obtained for the uncoated fiber composite at 23 and 1100 C are nearly linear to failure. The linear behavior, low UTS, and flat fracture surface all suggest a monolithic fracture behavior with no toughen￾ing from fibers. For temperatures 61100C and strains 60.1%, the stress–strain behavior of the N610/M/A com￾posite is also nearly linear elastic. However, as the strain exceeds 0.1% the stress–strain curves depart from linearity. The subsequent non-linear behavior, indicative of progres￾sive matrix cracking and crack deflection, continues until failure. The fracture surfaces show considerable fiber pull out. For both composites, the elastic modulus and UTS decrease with increasing temperature, while the failure strain remains fairly temperature-independent. Results in Table 2 reveal that at 1100 C, the addition of the monazite coating results in a nearly 50% increase in UTS. However, this increase in strength is accompanied by a 34% decrease in modulus and a threefold increase in failure strain. The stress–strain behavior of both composites changes dramatically at 1200 C. The stress–strain curves are lar￾gely nonlinear, the strength and modulus decrease signifi- cantly, while failure strains increase. The increase in temperature from 1100 to 1200 C causes a 10% loss in UTS and a 57% loss in modulus for the uncoated fiber composite, and a 17% loss in UTS and a 34% loss in mod￾ulus for the coated fiber composite. For both materials the failure strain at 1200 C is approximately 4.5 times that at 1100 C. It is important to note that in all tension tests, as well as in all other tests reported herein, the failure occurred within the gage section of the extensometer. 4.2. Creep rupture Results of the creep-rupture tests are summarized in Table 3, where creep strain accumulation and time to rup￾ture are shown for each test temperature and applied stress level. Creep curves obtained for the coated fiber composite at 1100 C are presented in Fig. 4. Creep curves obtained at 900 C for N610/M/A and N610/A are shown in Figs. 5(a) and (b), respectively. The time scale in Figs. 4(a) and 5 is reduced in order to clearly show the creep curves produced at higher stress levels. Results in Table 3 reveal that at 900 C the uncoated fiber composite did not achieve a run-out even for a low stress of 73 MPa. Recognizing that increase in temperature would result in reduced creep life, the N610/A was not expected to achieve a run-out at 1100 C for stresses P75 MPa, and was therefore not tested at that temperature. Creep curves produced by N610/M/A at 1100 C exhibit primary, secondary and tertiary creep regimes. Transition from primary to secondary creep occurs almost immedi￾ately. For stresses 680 MPa, secondary creep persists for 70% of the creep life before transitioning to tertiary creep. However, for stresses P100 MPa, secondary creep transitions to tertiary creep during the first third of the creep life. Creep strain accumulation decreases with increasing applied stress. While at 40 MPa creep strain is a significant 7.66%, at 120 MPa creep strain is only 0.7%. It is noteworthy that at 1100 C, for stress levels in the 40–120 MPa range, the accumulated creep strains signifi- cantly exceed the failure strain obtained in the tension test. The creep run-out, defined as survival of 100 h at a given creep stress, was not achieved. Table 2 Summary of tensile properties for the N610/M/A and the N610/A composites Specimen Temperature (C) Elastic modulus (GPa) UTS (MPa) Failure strain (%) N610/Monazite/Alumina composite B9-1 900 83 180 0.31 B5-1 1000 78 162 0.28 B2-1 1100 76 157 0.34 B1-1 1200 50 130 1.42 N610/Alumina composite B3-1 23 129 117 0.09 B3-2 1100 116 105 0.11 B8-1 1200 49 95 0.46 0 50 100 150 200 0.00 0.25 0.50 0.75 1.00 1.25 1.50 STRAIN (%) STRESS (MPa) 1200˚C 1200˚C N610/Monazite/Alumina N610/Alumina 1100˚C 1100˚C 900˚C 23˚C 1000˚C Fig. 3. Tensile stress–strain curves for N610/A and N610/M/A ceramic composites at various temperatures. Table 3 Summary of creep-rupture results for the N610/M/A and the N610/A composites Specimen Temperature (C) Creep stress (MPa) Creep strain (%) Time to rupture (s) N610/Monazite/Alumina composite B4-1 1100 40 7.66 50,432 B4-2 1100 80 3.36 1,452 B2-2 1100 100 1.58 360 B4-3 1100 120 0.70 75 B9-2 1000 80 0.05 63,060 B10-1 900 80 0.04 522,365a B10-2 900 120 0.04 432,175a B5-1 900 130 0.05 40,655 B11-1 900 140 0.03 54,075 B11-2 900 150 0.03 805 N610/Alumina composite B6-1 900 73 0.06 350,055 B6-2 900 80 0.03 19,995 a Run-out. 2092 M.B. Ruggles-Wrenn et al. / Composites Science and Technology 66 (2006) 2089–2099

M.B. Ruggles-Wrenn et al. Composites Science and Technology 66(2006)2089-2099 0.25 N610/Monazite/Alur 100c N610/Monazite/Alumina e0.15 120 MPa 0.10 140 MPa 150 MPa 0.05 120 MPa 40 MPa 100 0.00 2000040000 80000100000 Time(s) 610/Alumina (b) R5.0 e0.1 80 MPa 方0.10 0.05 80 MPa 73 MPa 100002000030000400005000060000 0.00 4000060000 Time(s Fig 4. Creep strain vs time curves for the N610/M/A ceramic composite at 1100C: (a)time scale is reduced to clearly show creep curves obtained Fig. 5. Creep strain vs time curves at 900C for:(a)N610/Monazite/ at stress levels >100 MPa and(b) time scale chosen to show creep strai Alumina and(b) N610/Alumin ccumulated at 40 MPa ble differences between the creep strains accumulated at dif Because creep rupture of this CMC is likely dominated ferent stress levels together with the inherent data scatter by creep of Nextel 610 fibers, it is useful to compare results account for an atypical order of the creep curves in presented here with those reported by Zawada et al. [37]for Fig. 5(a)(i.e 130 MPa creep curve showing slightly higher N610/AS. At 1100C and creep stress levels of 50 and strains than the 140 MPa curve). Note that all creep strains 75 MPa, N610/AS did not exhibit tertiary creep and conse- are an order of magnitude lower than the failure strain quently accumulated considerably lower creep strains. obtained in the tension test at 900C. Creep behavior of However, rupture times were comparable to those obtained the uncoated fiber composite is qualitatively similar to that for N610/M/A. Furthermore, N610/AS did not achieve a of the coated fiber CMC. Likewise, creep strains are similar run-out even for stress as low as 50 MPa. At 1000C, to those accumulated by the coated fiber composite N610/AS exhibited improved creep resistance, achieving Minimum creep rate was reached in all tests. Creep a run-out at 75 MPa. Conversely, creep resistance of the strain rate as a function of applied stress is shown in N610/M/A composite remained poor at 1000C. A scop- Fig. 6, where results of the present investigation are plotted ing creep test conducted at 1000C and 80 MPa failed after together with the data from Wilson and Visser [40]for 17.5h. Short creep lives at 1000 and 1100C, and exces- Nextel 610 fibers, Zawada et al. [37] for a N610/AS com- a ely large creep strains at 1100C reveal poor creep per- posite, and Casas and Martinez-Esnaola [41] for a Nextel mance, making this material unacceptable for 610/Mox composite. To further facilitate comparison applications invol sustained loading at temperatures between the creep properties of the fibers and the compos- >1000C. Therefore, creep performance at 900C was ites, the Nextel 610 fiber data adjusted for V=0.2 and examined Vr=0. 15 (volume fractions of the on-axis fibers in the A It is seen in Fig. 5(a) that the creep curves obtained for N610/M/A and N610/AS composites, respectively), are 10/M/A at 900C exhibit primary and secondary creep also shown regimes. Transition from primary to secondary creep As expected, the creep strain rates increase with increas- ure.Increasing creep stress appears to have little effect on pplied stress as well as with increasing temperature.At occurs early in creep life. Secondary creep continues to fail- ing creep rates of N610/M/A are what may creep strain(see Table 3), which remains <0.05%. Negligi- expected from Nextel 610 fibers alone. Experimental resu

Because creep rupture of this CMC is likely dominated by creep of Nextel 610 fibers, it is useful to compare results presented here with those reported by Zawada et al. [37] for N610/AS. At 1100 C and creep stress levels of 50 and 75 MPa, N610/AS did not exhibit tertiary creep and conse￾quently accumulated considerably lower creep strains. However, rupture times were comparable to those obtained for N610/M/A. Furthermore, N610/AS did not achieve a run-out even for stress as low as 50 MPa. At 1000 C, N610/AS exhibited improved creep resistance, achieving a run-out at 75 MPa. Conversely, creep resistance of the N610/M/A composite remained poor at 1000 C. A scop￾ing creep test conducted at 1000 C and 80 MPa failed after 17.5 h. Short creep lives at 1000 and 1100 C, and exces￾sively large creep strains at 1100 C reveal poor creep per￾formance, making this material unacceptable for applications involving sustained loading at temperatures P1000 C. Therefore, creep performance at 900 C was examined. It is seen in Fig. 5(a) that the creep curves obtained for N610/M/A at 900 C exhibit primary and secondary creep regimes. Transition from primary to secondary creep occurs early in creep life. Secondary creep continues to fail￾ure. Increasing creep stress appears to have little effect on creep strain (see Table 3), which remains 60.05%. Negligi￾ble differences between the creep strains accumulated at dif￾ferent stress levels together with the inherent data scatter, account for an atypical order of the creep curves in Fig. 5(a) (i.e. 130 MPa creep curve showing slightly higher strains than the 140 MPa curve). Note that all creep strains are an order of magnitude lower than the failure strain obtained in the tension test at 900 C. Creep behavior of the uncoated fiber composite is qualitatively similar to that of the coated fiber CMC. Likewise, creep strains are similar to those accumulated by the coated fiber composite. Minimum creep rate was reached in all tests. Creep strain rate as a function of applied stress is shown in Fig. 6, where results of the present investigation are plotted together with the data from Wilson and Visser [40] for Nextel 610 fibers, Zawada et al. [37] for a N610/AS com￾posite, and Casas and Martinez-Esnaola [41] for a Nextel 610/Umox composite. To further facilitate comparison between the creep properties of the fibers and the compos￾ites, the Nextel 610 fiber data adjusted for Vf = 0.2 and Vf = 0.15 (volume fractions of the on-axis fibers in the N610/M/A and N610/AS composites, respectively), are also shown. As expected, the creep strain rates increase with increas￾ing applied stress as well as with increasing temperature. At 1100 C, creep rates of N610/M/A are what may be expected from Nextel 610 fibers alone. Experimental results 0.0 0.5 1.0 1.5 2.0 0 100 200 300 400 Time (s) Strain (%) N610/Monazite/Alumina T = 1100˚C 80 MPa 100 MPa 120 MPa 40 MPa 0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 8.0 0 10000 20000 30000 40000 50000 60000 Time (s) Strain (%) N610/Monazite/Alumina T = 1100˚C 80 MPa 40 MPa (a) (b) Fig. 4. Creep strain vs time curves for the N610/M/A ceramic composite at 1100 C: (a) time scale is reduced to clearly show creep curves obtained at stress levels P100 MPa and (b) time scale chosen to show creep strain accumulated at 40 MPa. 0.00 0.05 0.10 0.15 0.20 0.25 0 20000 40000 60000 80000 100000 Time (s) Strain (%) N610/Monazite/Alumina T = 900˚C 80 MPa 120 MPa 140 MPa 130 MPa 150 MPa 0.00 0.05 0.10 0.15 0.20 0.25 0 20000 40000 60000 80000 100000 Time (s) Strain (%) N610/Alumina T = 900˚C 80 MPa 73 MPa (a) (b) Fig. 5. Creep strain vs time curves at 900 C for: (a) N610/Monazite/ Alumina and (b) N610/Alumina. M.B. Ruggles-Wrenn et al. / Composites Science and Technology 66 (2006) 2089–2099 2093