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steel interconnect joint under thermo-mechanical cycling
Renewable Energy 138 (2019) 1205e1213
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Mechanical durability of solid oxide fuel cell glass-ceramic sealant/ steel interconnect joint under thermo-mechanical cycling Chih-Kuang Lin a, *, Kun-Yi Chen a, Si-Han Wu b, Wei-Hong Shiu b, Chien-Kuo Liu b, Ruey-Yi Lee b a b
Department of Mechanical Engineering, National Central University, Jhong-Li District, Tao-Yuan City 32001, Taiwan Division of Nuclear Fuels and Materials, Institute of Nuclear Energy Research, Lung-Tan District, Tao-Yuan City 32546, Taiwan
a r t i c l e i n f o
a b s t r a c t
Article history: Received 4 September 2018 Received in revised form 7 December 2018 Accepted 9 February 2019 Available online 13 February 2019
A testing method is developed to quantitatively determine the thermo-mechanical cycling life in oxidizing atmosphere for the joint between a solid oxide fuel cell glass-ceramic sealant and a ferritic stainless steel interconnect. Thermo-mechanical cycling tests are performed under cyclic shear or tensile loading in conjunction with cyclic temperature variance between 40 C and 800 C. Results reveal thermo-mechanical cycling life under both shear and tensile loadings increases with a decrease in end stress at 800 C, for a certain end stress at 40 C. Nevertheless, for a certain end stress at 800 C, the tensile thermo-mechanical cycling life increases with a decrease in end stress at 40 C, while the shear thermo-mechanical cycling life is independent of end stress at 40 C. A difference in fracture pattern is also observed between shear and tensile loadings. For shear loading, fracture mainly takes place along the interface between glass-ceramic sealant and an oxide layer, such as BaCrO4 or Cr2O3. However, for tensile loading, fracture mainly occurs within the glass-ceramic layer, following crack initiation at the interface of Cr2O3/sealant or Cr2O3/BaCrO4. © 2019 Elsevier Ltd. All rights reserved.
Keywords: Solid oxide fuel cell Glass-ceramic sealant Interconnect Joint Thermo-mechanical cycling
1. Introduction Solid oxide fuel cell (SOFC) is an energy conversion device which directly converts chemical energy into electricity by a series of electrochemical reactions. Compared to other fuel cells developed, SOFC generally has the highest efficiency of energy conversion and a better flexibility of fuel utilization (e.g. pure hydrogen and hydrocarbons) thanks to a high operating temperature [1]. Planar SOFCs (pSOFCs) have recently received more attention than tubular ones due to an easier fabricating process and a higher power density. For an anode-supported cell design with a thinner electrolyte, intermediate-temperature pSOFCs (IT-pSOFCs) can operate at a temperature between 600 C and 800 C due to a lower ohmic loss. A pSOFC stack is generally assembled with multiple unit cells which are connected in series by interconnects and bonded by sealants for a higher voltage and power in practical applications. Hermetic seals are required in joining SOFC components and must be stable at high temperature in both oxidizing and reducing
* Corresponding author. E-mail address: [email protected] (C.-K. Lin). https://doi.org/10.1016/j.renene.2019.02.041 0960-1481/© 2019 Elsevier Ltd. All rights reserved.
atmospheres to prevent leakage and fuel-oxidant mixing. Glass and glass-ceramic sealants have been extensively studied and practically used to seal SOFC components as they have a favorable combination of properties to satisfy the requirements [1]. Interconnects not only play a crucial role in electrical connection of repeated cells and separation of fuel and oxidant gas but also provide the necessary structural support. Reducing the operating temperature to 600 Ce800 C in IT-pSOFCs makes it possible to use metallic alloys, especially ferritic stainless steels, as interconnects for lower cost, better workability, higher electrical conductivity, and better mechanical properties [2,3]. Joining glassceramic sealant with metallic interconnect/frame is commonly seen in a rigid type of sealing for IT-pSOFC stack [4]. Although bonding characteristics such as chemical stability and interfacial compatibility are usually considered first for glassceramic sealant/metallic interconnect joint in IT-pSOFC stacks, structural integrity/durability is also an important issue because of thermal stresses generated in the stack. The major sources of thermal stress in a pSOFC stack during thermal cycles of operation include thermal gradient and mismatch of coefficient of thermal expansion (CTE) between adjacent components [5,6]. Therefore, study of mechanical strength and durability of sealant/interconnect
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joint is necessary for improvement in reliability and durability of pSOFC systems. Mechanical properties of SOFC glass-ceramic sealant/metallic interconnect joint have been studied in prior work [7e20]. However, most of those studies were mainly focused on the initial bonding strength and limited work was conducted on the long-term high-temperature creep behavior at steady operation stage [17,20]. In addition to consideration of creep mechanism acting on the SOFC sealant/interconnect joint at a hightemperature operating environment, effect of thermal cycling on the mechanical durability should not be ignored, either. Depending on application type, a pSOFC stack needs to survive possibly several to thousands of thermal cycles (including start-up, steady operation, and shutdown stages) during service life [21e24]. During periodic operation, thermal stresses change simultaneously and cyclically with temperature in each component of a pSOFC stack [5,6]. When an engineering component is subjected to a combined cyclic thermal and mechanical loading, it could suffer from thermomechanical fatigue (TMF) damage [25]. As such cyclic thermomechanical loading damage may accelerate degradation of the sealant/interconnect joint robustness, it is necessary to assess such effect on the structural integrity and cycle life of pSOFC stack. Effects of thermal cycling on the joining characteristics between glass-ceramic sealant and metallic interconnect have been investigated in a few studies [22e24,26,27], but only pure thermal cycling was applied in those studies without considering the synergistic effect of cyclic mechanical loading. For example, Smeacetto et al. [23,24] studied the effect of thermal cycling from room temperature (RT) to 800 C on interfacial stability and adherence between glass-ceramic sealant and metallic interconnect. In another study, Weil et al. [27] compared the rupture strength of glass-ceramic/metallic interconnect joints before and after thermal cycling from <70 C to 800 C. Their results [27] show that the decrease in seal strength of the joint specimens after a number of thermal cycles is attributed to the difference in CTE between crystalline phases and glassy phase and to the interfacial reaction products between glass-ceramic sealant and metallic interconnect. However, study on the mechanical durability of SOFC sealant/ interconnect joint under cyclic thermo-mechanical loading is still lacking. The aim of this study is thus to quantitatively assess the thermo-mechanical cycling (TMC) life in oxidizing atmosphere for the joint between a glass-ceramic sealant (GC-9) and an interconnect steel (Crofer 22 H). It is hoped that results of this study and previous studies [16e20] can provide a comprehensive understanding of the mechanical behavior of such a joint for better design of a reliable and durable IT-pSOFC stack.
stainless steel, Crofer 22 H (ThyssenKrupp VDM GmbH, Werdohl, Germany), which is developed for IT-pSOFC interconnect. Chemical composition and mechanical properties of the given interconnect steel have been studied previously [28,29]. Metallic coupons of the joint specimens are prepared and machined in a way similar to that described in Ref. [16]. A borosilicate glass (designated as GC-9) for pSOFC sealing is applied to joining the two steel coupons in each joint specimen. The major composition in mol% of this glass-ceramic sealant includes 34 BaO, 34 SiO2, 12 CaO, 9.5 B2O3, 5 La2O3, 4.5 Al2O3, and 1 ZrO2 [30]. GC-9 has proved to be suitable for application in pSOFCs thanks to a good combination of thermal properties, chemical compatibility and stability, and hermetic properties [31,32]. Mechanical properties of GC-9 sealant have been investigated in prior work [33e36]. As the sandwich joint specimens are prepared and fabricated following the same way given in a previous study, detailed specimen preparation procedures are described in that study [16]. 2.2. Thermo-mechanical cycling test Previous analyses of thermal stress in a pSOFC stack indicate that the thermal stress generated during periodic operation varies with temperature and location in a component [5,6]. The maximum stress at a certain position at RT is greater than that at hightemperature operation stage [5,6]. It is due to a larger temperature difference between RT and the initial stress-free assembling temperature (800 C) of the given pSOFC stack [5,6]. In a pSOFC stack, glass-ceramic sealants and metallic interconnects are joined together for tight-sealing at high temperature. During joining at high temperature, glass-ceramic is soft enough to create an initial “stress free” condition in the joints. After cooling, thermal stresses are generated within the joints as a result of CTE mismatch and temperature difference/gradient. Such a case in which thermal stress varies cyclically with temperature at a certain position in a pSOFC stack is similar to TMC. To simulate such TMC experienced by the sealant/interconnect joints in a pSOFC stack, a TMC pattern (Fig. 1) is applied to the given sandwich joint specimens. In each TMC test performed in air, cyclic mechanical loading in conjunction with cyclic temperature variance is applied using a commercial closed-loop servo-hydraulic material test machine attached with a furnace. This TMC testing technique has been developed and employed in studying the TMF behavior of Crofer 22 H steel such that details of the testing equipment and procedures can be found
2. Experimental procedures 2.1. Materials and specimens As compressive loads normal to the interface of a joint are not expected to cause interfacial cracking or debonding, only shear and tensile loadings are applied for cyclic thermo-mechanical test in the present study. In this regard, two types of thin-plate sandwich joint specimen (metal/glass-ceramic sealant/metal) with a thickness of 2.5 mm are designed and made, one for shear loading test and the other for tensile loading test. Pin holes are drilled at both ends of each specimen for applying pin loading to minimize bending and twisting effects during mechanical testing. Tensile or shear stress can be generated at the interface of metal/glass-ceramic joint by applying a uni-axial tensile loading. Such specimen designs have been developed for assessing the bonding strength of a similar metal/glass-ceramic joint in a previous study [16] in which detailed specimen geometry and dimensions can be found. The metallic parts of the joint specimens are made of a commercial ferritic
Fig. 1. Schematic diagram of thermo-mechanical loading applied in TMC test.
C.-K. Lin et al. / Renewable Energy 138 (2019) 1205e1213
in Ref. [29]. As shown in Fig. 1, a cyclic temperature range of 40 Ce800 C with a heating and cooling rate of 5 C min1 is applied to simulating the temperature range between shutdown and steady operation stages in a typical pSOFC system. As restricted by the cooling ability of the furnace, it takes 60 min to cool down from 100 C to 40 C in furnace cooling. Therefore, the total time for one period of TMC is 352 min (5.87 h). Together with the temperature change, the cyclic mechanical loading in Fig. 1 is applied under force control. Thermal stress also varies with location of the sealant/ interconnect joint in periodic operation of SOFC [5,6]. Therefore, several stress levels of specified joint strength ratios (JSRs) at 40 C and 800 C are selected in performing the TMC tests to simulate various combinations of thermal stresses experienced in the joints of a pSOFC stack in periodic operation. JSR is defined as the applied peak or valley stress divided by the corresponding bonding strength at that specific peak or valley temperature. In other words, JSR is a normalized stress level. JSRs of 0.2, 0.4, and 0.6 are selected as the peak or valley stress of mechanical loading applied at 40 C and 800 C in each cycle. In the following, (x, y) is used to represent the applied end stresses in a TMC test. x is the JSR corresponding to the end stress applied at 40 C while y is the JSR corresponding to the end stress applied at 800 C. For example, a (0.4, 0.2) shear TMC loading indicates a shear stress of 2.64 MPa (JSR ¼ 0.4) is applied at 40 C and then changes bi-linearly with temperature to reach a level of 0.94 MPa (JSR ¼ 0.2) at 800 C in a half cycle and vice versa for the other half cycle, as shown in Fig. 1. Note that the average shear bonding strength of the given joint is of 6.6 MPa at RT and 4.7 MPa at 800 C, while the tensile counterpart has a value of 23 MPa at RT and 12.7 MPa at 800 C [16]. As shown in Fig. 1, each TMC test starts from 800 C. Note that in some TMC tests the end stress applied at 40 C is greater than that applied at 800 C and vice versa for the others. Note that each line segment in Fig. 1 represents a constant stress rate or thermal rate through appropriate force and temperature control such that both temperature and stress reach the end levels at the same time. 2.3. Fractography analysis Fracture surfaces of broken specimens after TMC test were observed using optical microscopy (OM) and scanning electron microscopy (SEM). For composition analysis of elemental distribution on the fracture surfaces, an energy dispersive spectrometer (EDS) module attached in the SEM was employed. X-ray diffraction (XRD) was applied to determining the crystalline phases. 3. Results and discussion 3.1. Thermo-mechanical cycling under shear loading Table 1 lists the number of cycles to failure for joint specimens subjected to various shear TMC loadings. Note that the TMC test was terminated if the specimen was not ruptured at 50 or 60 cycles.
Table 1 Number of cycles to failure for various shear TMC loadings. Stress (JSR) applied at 40 C
Number of cycles to failure Stress (JSR) applied at 800 C
1.32 MPa (0.2) 2.64 MPa (0.4) 3.96 MPa (0.6) a
Runout test.
0.94 MPa (0.2)
1.88 MPa (0.4)
2.82 MPa (0.6)
>50a >60a >50a
6 8 6
<1 <1 1
1207
Accordingly, it is considered a runout test and the TMC life is expressed as > 50 or >60 cycles in Table 1. As shown in Table 1, the number of cycles to failure under shear cyclic loading is increased with a decrease in JSR applied at 800 C, with a given JSR applied at 40 C. However, the end stress applied at 40 C does not affect the shear TMC life significantly as the TMC life is comparable for a certain JSR applied at 800 C regardless of the JSR applied at 40 C. Therefore, shear TMC life is controlled by the end stress applied at 800 C and may be considered an accumulated duration of creep loading at 800 C. As a combination of high temperature environment and mechanical stress may generate creep damage, the TMC life is compared with the estimated creep rupture time through a creep stress-life relation given in a previous study [17]. To make a comparison of TMC life with accumulated duration of creep loading, the estimated creep rupture time is calculated by the creep stress-life relation determined for a similar Crofer 22 H/GC-9/Crofer 22 H joint tested at 800 C [17]. The creep stress-life relation under constant shear loading at 800 C is expressed as follows [17],
Shear creep loading : ttr0:066 ¼ 1:68
(1)
where t is the applied shear stress in unit of MPa and tr is time to rupture in unit of h. As the creep stress-life relation above is obtained under constant shear loading at 800 C, the estimated creep rupture time for each shear TMC test is calculated by substituting into Eq. (1) the end stress applied at 800 C. In consideration of deviation in temperature control during mechanical test in the present study and Ref. [17], a small temperature range of 795 Ce800 C around the peak temperature is selected and considered for generating a similar creep loading effect at 800 C. In this way, the sample is assumed to experience a similar creep loading effect of 800 C for 2 min at 795 Ce800 C in each thermo-mechanical loading cycle, due to a heating and cooling rate of 5 C min1. In other words, a TMC cycle is equivalent to 2 min of creep loading at 800 C. Table 2 shows the accumulated duration of creep loading at 795 Ce800 C for various shear TMC loadings and the corresponding estimated creep rupture time. According to Eq. (1), the shear stress applied at 800 C for a creep rupture time over 1000 h is about 1 MPa which is slightly higher than 0.94 MPa. It is thus expected that the shear TMC life for an end stress of 0.94 MPa (0.2 JSR) applied at 800 C is much longer than 60 cycles (352.2 h). For shear loading of 0.4 JSR (1.88 MPa) applied at 800 C combined with various JSRs applied at 40 C, the TMC life is of 6e8 cycles which is equivalent to a duration of 12e16 min of creep loading at 795 Ce800 C. An estimated creep rupture time by Eq. (1) for a shear stress of 1.88 MPa applied at 800 C is about 12 min which is comparable with the accumulated duration of creep loading at 795 Ce800 C for shear TMC tests with 0.4 JSR (1.88 MPa) applied at 800 C. For a shear TMC test with 0.6 JSR (2.86 MPa) applied at 800 C, the estimated creep rupture time is 0.03 min, so the TMC test is expected not to run more than 1 cycle. As shown in Table 2, the accumulated duration of creep loading in shear TMC test is in good agreement with the estimated creep rupture time. It indicates that creep mechanism at 800 C may play an important role in determining the TMC life under shear loading. Fig. 2 shows the representative failure patterns of TMC under shear loading. Note that in Fig. 2 and the following similar figures, fracture surfaces on both sides of each broken specimen are presented with mirror symmetry. Fig. 2(a) shows the fracture pattern for TMC life of 6 cycles under (0.6, 0.4) shear loading. SEM micrographs of two selected areas in the lower fracture surface of Fig. 2(a), one at the peripheral region and the other at the central region, are shown in Fig. 3. Note that on each fracture surface shown in Fig. 2, the peripheral region and the central region are
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Table 2 Accumulated duration of creep loading at 795 Ce800 C for various shear TMC loadings. Stress (JSR) applied at 40 C
Accumulated duration of creep loading at 795 Ce800 C (min) Stress (JSR) applied at 800 C
1.32 MPa (0.2) 2.64 MPa (0.4) 3.96 MPa (0.6) Estimated creep rupture time by Eq. (1) a
0.94 MPa (0.2)
1.88 MPa (0.4)
2.82 MPa (0.6)
>100a >120a >100a 404,098
12 16 12 11.9
<2 <2 2 0.03
Runout test.
(a)
(b) Fig. 2. Representative failure patterns in shear specimens of various TMC lives: (a) 6 cycles; (b) < 1 cycle.
separated by a dash-line loop. SEM micrograph of a peripheral yellow area is shown in Fig. 3(a) and identified as a BaCrO4 layer on top of GC-9 glass-ceramic substrate. BaCrO4 has been found to exhibit a yellow color in a previous study [17]. Existence of BaCrO4 in Fig. 3(a) is also confirmed by EDS analysis, as peaks of Ba, Cr, and O are detected in the EDS spectrum (Fig. 4). Note that in the EDS spectrum of Fig. 4, Pt is also detected as it is sputtered on the fracture surface to increase conductivity for SEM observation. As needle-shape crystalline phases (a-Ba(Al2Si2O8)) of GC-9 glassceramic are also observed in Fig. 3(a) and Si, Al, and Ca elements are detected by EDS (Fig. 4), GC-9 glass-ceramic also exists in the micrograph of Fig. 3(a). Note that a-Ba(Al2Si2O8) is the main crystalline phase in GC-9 glass-ceramic [35]. Therefore, the yellow peripheral region in the lower fracture surface of Fig. 2(a) is a BaCrO4 layer mixed with GC-9 glass-ceramic. In Fig. 3(b), needleshape a-Ba(Al2Si2O8) phases are observed and Cr2O3 is detected by EDS analysis, indicating the light green central region in the lower fracture surface of Fig. 2(a) is a Cr2O3 layer mixed with GC-9
glass-ceramic. To further confirm existence of BaCrO4, Cr2O3, and aBa(Al2Si2O8) phases, XRD is employed to analyze the phase structure in the lower fracture surface of Fig. 2(a). As shown in Fig. 5, the diffraction peaks match those of BaCrO4, Cr2O3, and a-Ba(Al2Si2O8) phases. SEM micrographs of two selected areas in the upper fracture surface of Fig. 2(a), one at the peripheral region and the other at the central region, are shown in Fig. 6. Note the micrographs shown in Fig. 6 are the counterparts of Fig. 3. Based on EDS analysis results and existence of needle-shape crystalline phases in Fig. 6(a), the peripheral region in the upper fracture surface of Fig. 2(a) is a thin peeled BaCrO4 layer mixed with GC-9 glass-ceramic. Similarly, EDS analysis results and existence of chromia structure in Fig. 6(b) indicate the central region in the upper fracture surface of Fig. 2(a) is a peeled Cr2O3 layer mixed with GC-9 glass-ceramic. As the joining process of the sandwich joint specimen was conducted in air, a BaCrO4 layer is formed along the edge of joining area. However, as the central region is far from the edge, oxygen from the ambient air is not easy to diffuse into the deep interior. Therefore, chromate layer is barely observed in the central part of the joining area. This observation is in agreement with that in the study of Yang [37]. Barium chromate is produced by the reaction of oxygen with BaO and Cr2O3 [37]. The fracture surfaces of (0.4, 0.6) shear TMC loading with a life of <1 cycle are shown in Fig. 2(b). Its failure pattern is almost the same as that of the sample under (0.6, 0.4) shear TMC loading (Fig. 2(a)), indicating a similar fracture mechanism. The only difference between them is that a greater amount of BaCrO4 (a larger yellow peripheral region) is observed on the fracture surface of (0.6, 0.4) sample (Fig. 2(a)). Apparently, for a longer testing time in air, more BaCrO4 is formed on the joint interface at the peripheral region. Based on the fractography observations described above for the joint samples under shear TMC loadings, cracks were initiated at the BaCrO4/GC-9 interface in the periphery, followed by propagation and final fracture at the Cr2O3/GC-9 interface in the central joining area, as shown in Fig. 2. Apparently, shear fracture prefers to take place at the interfaces between chromate, chromia, and glassceramic layers due to a larger CTE mismatch. Note that the CTE values for BaCrO4, GC-9 glass-ceramic, Cr2O3, and Crofer 22 H alloy are of 20.4 106 C1 (at 20e813 C) [38], 13.1 106 C1 (at 20e650 C) [32], 9.6 106 C1 (at 20e1400 C) [39], and 11.8 106 C1 (at 20e800 C) [40], respectively. A large thermal mismatch could generate thermal stress and result in formation of defects and/or microcracks which would serve as the fracture origin and/or provide a weak path for cracking [41]. This is why cracking begins at the BaCrO4/GC-9 interface in the peripheral region and switches to the Cr2O3/GC-9 interface in the central region of the joining area when subjected to shear TMC loading.
3.2. Thermo-mechanical cycling under tensile loading Table 3 lists the number of cycles to failure for joint specimens
C.-K. Lin et al. / Renewable Energy 138 (2019) 1205e1213
(a)
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(b)
Fig. 3. SEM micrographs of fracture surface in the lower part of Fig. 2(a): (a) peripheral region (BaCrO4 with GC-9); (b) central region (Cr2O3 with GC-9).
cycles to failure increases with a decrease in applied JSR at 40 C. It indicates that the tensile TMC life is controlled not only by the end stress applied at 800 C but also by that applied at 40 C. This is different from the trend of shear TMC life which is dominated only by the JSR applied at 800 C. As described in the following failure analysis, fracture mainly takes place within the glass-ceramic layer in tensile specimens. Glass-ceramic is more sensitive to tensile loading at low temperature due to its brittle behavior. Therefore, the JSR applied at 40 C also plays a certain role in determining the tensile TMC life. Similar to shear loading, the estimated creep rupture time for a given constant tensile stress at 800 C can be calculated through the creep stress-life relation determined in the previous study [17]. The creep stress-life relation under constant tensile load at 800 C in air is expressed as follows [17],
tensile creep loading : str0:074 ¼ 1:85 Fig. 4. EDS analysis result of Fig. 3(a).
Fig. 5. XRD pattern for the lower fracture surface of Fig. 2(a).
subjected to tensile TMC loadings. It shows that the number of cycles to failure under tensile TMC is decreased with an increase in JSR applied at 800 C, with a given JSR applied at 40 C. For a given tensile stress of 2.54 MPa (0.2 JSR) applied at 800 C, the number of
(2)
where s is the applied tensile stress in unit of MPa and tr is time to rupture in unit of h. Table 4 lists the accumulated duration of creep loading at 795 Ce800 C for joint specimens under various tensile TMC loadings. The bottom row in Table 4 is the estimated creep rupture time calculated by Eq. (2). Unlike the shear loading results, the accumulated duration of tensile creep loading at 795 Ce800 C in experiment is not close to the estimated creep rupture time, in particular for JSR of 0.2 applied at 800 C (Table 4). It reveals that the creep effect at 800 C does not play a dominating role in determining the tensile TMC life. Therefore, a true cycling effect is responsible for the TMC failure under tensile loading. Representative failure patterns in tensile specimens with various TMC lives are shown in Fig. 7. Fig. 7(a) exhibits a fracture pattern for TMC life of <1 cycle under (0.4, 0.6) tensile loading. SEM micrographs of two selected areas on the fracture surface of Fig. 7(a), one from the heavy green region and the other from the white region, are shown in Fig. 8. As shown in Fig. 8(a), chromia features are found and confirmed by EDS analysis for the heavy green region in Fig. 7(a). As shown in Fig. 8(b), needle-shape crystalline phases (a-Ba(Al2Si2O8)) are found in the white region of Fig. 7(a) indicating fracture occurred within the glass-ceramic layer. Therefore, for the sample under (0.4, 0.6) tensile TMC loading, cracking started peripherally at the Cr2O3/GC-9 interface, and then proceeded inward into the glass-ceramic layer.
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(a)
(b)
Fig. 6. SEM micrographs of fracture surface in the upper part of Fig. 2(a): (a) peripheral region (BaCrO4 with GC-9); (b) central region (peeled Cr2O3 with GC-9).
3.3. Overall comparison
Table 3 Number of cycles to failure for various tensile TMC loadings. Stress (JSR) applied at 40 C
Number of cycles to failure
Stress (JSR) applied at 800 C
4.6 MPa (0.2) 9.2 MPa (0.4) 13.8 MPa (0.6)
2.54 MPa (0.2)
5.08 MPa (0.4)
7.62 MPa (0.6)
11 3 1
<1 <1 <1
<1 <1 <1
Fig. 7(b) shows the fracture pattern for TMC life of 11 cycles under (0.2, 0.2) tensile loading. Similarly, by means of EDS analysis and microstructural observation, the heavy green and white regions in Fig. 7(b), again, are identified as Cr2O3 and GC-9, respectively. However, additional yellow regions are observed at the bottom edges of both fracture surfaces presented in Fig. 7(b). Note that these yellow regions are the counterparts of the heavy green regions at top edges of the two mirror-symmetric fracture surfaces of Fig. 7(b). By means of EDS analysis, elements O, Si, Cr and Ba are detected and needle-shape crystalline phases of GC-9 are observed in the yellow regions, as shown in Fig. 9. It indicates the yellow regions at the bottom edges of Fig. 7(b) are BaCrO4 mixed with GC9. Therefore, for tensile TMC loading with a life of 11 cycles, crack was initiated at the peripheral interface between Cr2O3 and BaCrO4 layers and subsequently penetrated into the GC-9 layer. As the testing time of (0.2, 0.2) tensile TMC loading is longer than that of (0.4, 0.6), i.e. 11 cycles vs. < 1 cycle, it is expected more BaCrO4 is formed at the periphery of joining area. Consequently, fracture of (0.2, 0.2) specimen originated at the Cr2O3/BaCrO4 interface rather than at the Cr2O3/GC-9 interface.
As described in previous sections, TMC life under both shear and tensile loadings is decreased with an increase in end stress (JSR) applied at 800 C, with a given end stress applied at 40 C. However, given an end stress applied at 800 C, the tensile TMC life is increased with decreasing end stress applied at 40 C, while the shear TMC life is independent of end stress applied at 40 C. As shown in Tables 1 and 3, for a given combination of cyclic end stresses (x, y), the number of cycles to failure under tensile TMC loading is significantly smaller than the shear counterpart. It reveals that the given Crofer 22 H/GC-9/Crofer 22 H joint suffers from tensile loading to a greater extent than that from shear loading, with regard to TMC effect on structural integrity. Such a difference in the trend of TMC life between shear and tensile loadings also appears in their failure patterns. As shown in Fig. 2, fracture of TMC under shear loading takes place initially at the BaCrO4/GC-9 interface in the peripheral region and then proceeds along the Cr2O3/GC-9 interface in the interior of joining area. For fracture of TMC under tensile loading, cracking starts peripherally at the Cr2O3/GC-9 interface (<1 cycle) or at the Cr2O3/BaCrO4 interface (11 cycles), and then penetrates inward into the GC-9 glass-ceramic layer, as shown in Fig. 7. As glass-ceramic becomes viscous at temperature above its glass transition temperature, viscous deformation and relevant damage, such as creep, are more susceptible to shear loading. Therefore, development of cracking in shear TMC loading presumably takes place mostly in the high-temperature range of a thermo-mechanical loading cycle. As a result, the shear TMC life is controlled by the end stress applied at 800 C, irrespective of the end stress applied at 40 C. Note the glass transition temperature and softening temperature of GC-9 glass-
Table 4 Accumulated duration of creep loading at 795 Ce800 C under various tensile TMC loadings. Stress (JSR) applied at 40 C
Accumulated duration of creep loading at 795 Ce800 C (min) Stress (JSR) applied at 800 C
4.6 MPa (0.2) 9.2 MPa (0.4) 13.8 MPa (0.6) Estimated creep rupture time by Eq. (2)
2.54 MPa (0.2)
5.08 MPa (0.4)
7.62 MPa (0.6)
22 6 2 0.88
<2 <2 <2 <0.1
<2 <2 <2 <0.1
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(a)
Fig. 9. SEM micrograph of the yellow region at bottom edge of the lower fracture surface in Fig. 7(b).
(b) Fig. 7. Representative failure patterns in tensile specimens of various TMC lives: (a) < 1 cycle; (b) 11 cycles.
ceramic are of 668 C and 745 C, respectively [33]. As glass-ceramic itself is a brittle material at low temperature, it is more vulnerable to tensile stress in terms of fracture resistance. Therefore, cracking also takes place within the GC-9 layer under tensile TMC loading. Therefore, TMC damage develops during the entire tensile thermomechanical loading cycle. In this regard, tensile TMC life is dependent on both end stresses applied at 800 C and 40 C. This is consistent with the aforementioned failure patterns in which shear TMC failure all takes place at the GC-9/oxide interfaces while the tensile TMC fracture occurs mainly within the GC-9 layer in addition to initiation at the interfaces. It is required that glass-ceramic sealant for SOFC applications
(a)
provide a good gas tightness during sealing and operating processes at high temperatures. Leak rate for a comparable Crofer 22 APU/GC-9/Crofer 22 APU joint was measured under pure thermal cycling (RTe800 C) in a previous study [32]. Note that Crofer 22 APU is a former generation of Crofer 22 H alloy for SOFC interconnect, with addition of Nb and W in the latter. For the sealed Crofer 22 APU/GC9/Crofer 22 APU coupon subjected to 50 repetitions of pure thermal cycling from RT to 800 C, the leakage rate during the test kept a low value which is well below the allowable leak rate limit [32]. It indicates the GC-9 glass-ceramic performs well in sealing metallic interconnects and/or frames for pSOFC stacks when subjected to pure thermal cycling. However, results of the present work reveal that combining thermal cycling with cyclic mechanical loading could affect the mechanical durability of the sealed joint and deteriorate the sealing performance under certain TMC conditions. For example, the given Crofer 22 H/GC-9/Crofer 22 H joint could survive more than 50 cycles under a (0.2, 0.2), (0.4,
(b)
Fig. 8. SEM micrographs of fracture surface in Fig. 7(a): (a) heavy green region (Cr2O3); (b) white region (GC-9).
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0.2), or (0.6, 0.2) shear TMC loading, but would lose its structural integrity after 11 or a smaller number of cycles under a (0.2, 0.2), (0.4, 0.2), or (0.6, 0.2) tensile TMC loading. The pure thermal cycling tests carried out in previous studies [22e24,26,27,32] are suitable for assessing the long-term chemical stability and interfacial compatibility with adjacent components for SOFC sealants. However, as working temperature and thermal stress change simultaneously in a working cycle of SOFC system, it is thus suggested that TMC test should be applied in quantitatively assessing the mechanical durability of metal/sealant/metal joints for design of a durable pSOFC stack. Previous studies have proved that GC-9 glass-ceramic is suitable for SOFC sealing application as it shows good thermal properties, chemical stability and compatibility, hermetic properties, and mechanical properties [30e36]. The TMC testing method developed in the present study is able to quantitatively determine the mechanical durability of SOFC sealant/interconnect joints in oxidizing atmosphere, which has not yet been reported in the literature. However, the interaction between glass-ceramic sealant and metallic interconnect in oxidizing atmosphere may differ from that in reducing atmosphere which may influence the TMC behavior in a different manner. Therefore, a counterpart study in a reducing environment is being conducted to investigate the effect of reducing atmosphere on the TMC behavior of the given GC-9/ Crofer 22 H joint, which is considered an extension of the present study and the results will be published in the near future. The TMC testing results of sealed Crofer 22 H/GC-9/Crofer 22 H coupons presented in this study in conjunction with other mechanical properties such as bonding strength, creep strength, and interfacial fracture resistance obtained in prior work [16e19] could provide a comprehensive mechanical data base for design and development of reliable pSOFC stacks using GC-9 glass-ceramic sealant.
4. Conclusions (1) For the given joint of GC-9 glass-ceramic sealant/Crofer 22 H steel tested in oxidizing atmosphere, shear TMC life is increased with a decrease in end stress applied at 800 C. However, it is independent of end stress applied at 40 C. The accumulated duration of creep loading at 795 Ce800 C in shear TMC test is comparable with the estimated creep rupture time. It reveals that creep effect plays an important role in determining the shear TMC life. (2) The tensile TMC life is affected by both end stresses applied at 800 C and 40 C, showing an increasing life with decreasing end stresses applied. The accumulated duration of creep loading at 795 Ce800 C in tensile TMC test does not match the estimated creep rupture time. It indicates creep effect is not the sole dominating factor in determining tensile TMC life. (3) Fracture of TMC under shear loading takes place initially at the BaCrO4/GC-9 interface in the peripheral region and then proceeds along the Cr2O3/GC-9 interface in the interior. For fracture of TMC under tensile loading, cracking starts peripherally at the Cr2O3/GC-9 or Cr2O3/BaCrO4 interface, and then penetrates inward into the GC-9 glass-ceramic layer. (4) For the given GC-9/Crofer 22 H joint in oxidizing atmosphere, a difference in the trend of TMC life between shear and tensile loadings is observed. This is consistent with the difference found in failure patterns in which the shear TMC failure takes place at the GC-9/oxide interfaces while the tensile TMC fracture occurs mainly within the GC-9 layer in addition to initiation at the interfaces.
Note The authors declare no competing interests. Acknowledgments This work was supported by the Institute of Nuclear Energy Research (Taiwan) [Grant No. 105-2001-INER-013] and the Ministry of Science and Technology (Taiwan) [Grant No. MOST 103-2221-E008-014-MY2]. References [1] R.N. Singh, Sealing technology for solid oxide fuel cells (SOFC), Int. J. Appl. Ceram. Technol. 4 (2007) 134e144. [2] J.C.W. Mah, A. Muchtar, M.R. Somalu, M.J. Ghazali, Metallic interconnects for solid oxide fuel cell: a review on protective coating and deposition techniques, Int. J. Hydrogen Energy 42 (2017) 9219e9229. [3] N. Shaigan, W. Qu, D.G. Ivey, W. Chen, A review of recent progress in coatings, surface modifications and alloy developments for solid oxide fuel cell ferritic stainless steel interconnects, J. Power Sources 195 (2010) 1529e1542. [4] P.A. Lessing, A review of sealing technologies applicable to solid oxide electrolysis cells, J. Mater. Sci. 42 (2007) 3465e3476. [5] C.-K. Lin, T.-T. Chen, Y.-P. Chyou, L.-K. Chiang, Thermal stress analysis of a planar SOFC stack, J. Power Sources 164 (2007) 238e251. [6] C.-K. Lin, L.-H. Huang, L.-K. Chiang, Y.-P. Chyou, Thermal stress analysis of planar solid oxide fuel cell stacks: effects of sealing design, J. Power Sources 192 (2009) 515e524. [7] F. Smeacetto, M. Salvo, M. Ferraris, V. Casalegno, P. Asinari, A. Chrysanthou, Characterization and performance of glass-ceramic sealant to join metallic interconnects to YSZ and anode-supported-electrolyte in planar SOFCs, J. Eur. Ceram. Soc. 28 (2008) 2521e2527. [8] E.V. Stephens, J.S. Vetrano, B.J. Koeppel, Y. Chou, X. Sun, M.A. Khaleel, Experimental characterization of glass-ceramic seal properties and their constitutive implementation in solid oxide fuel cell stack models, J. Power Sources 193 (2009) 625e631. [9] Y.-S. Chou, J.W. Stevenson, P. Singh, Effect of pre-oxidation and environmental aging on the seal strength of a novel high-temperature solid oxide fuel cell (SOFC) sealing glass with metallic interconnect, J. Power Sources 184 (2008) 238e244. [10] Y.-S. Chou, J.W. Stevenson, P. Singh, Effect of aluminizing of Cr-containing ferritic alloys on the seal strength of a novel high-temperature solid oxide fuel cell sealing glass, J. Power Sources 185 (2008) 1001e1008. [11] J. Malzbender, Y. Zhao, Micromechanical testing of glass-ceramic sealants for solid oxide fuel cells, J. Mater. Sci. 47 (2012) 4342e4347. [12] J. Malzbender, Y. Zhao, Flexural Strength and viscosity of glass ceramic sealants for solid oxide fuel cell stacks, Fuel Cells 12 (2012) 47e53. [13] J. Malzbender, R.W. Steinbrech, L. Singheiser, Determination of the interfacial fracture energies of cathodes and glass ceramic sealants in a planar solidoxide fuel cell design, J. Mater. Res. 18 (2003) 929e934. [14] J. Malzbender, R.W. Steinbrech, L. Singheiser, P. Batfalsky, Fracture energies of brittle sealants for planar solid oxide fuel cells, Ceram. Eng. Sci. Proc. 26 (2005) 285e291. [15] A. Müller, W. Becker, D. Stolten, J. Hohe, A hybrid method to assess interface debonding by finite fracture mechanics, Eng. Fract. Mech. 73 (2006) 994e1008. [16] C.-K. Lin, J.-Y. Chen, J.-W. Tian, L.-K. Chiang, S.-H. Wu, Joint strength of a solid oxide fuel cell glasseceramic sealant with metallic interconnect, J. Power Sources 205 (2012) 307e317. [17] C.-K. Lin, K.-L. Lin, J.-H. Yeh, S.-H. Wu, R.-Y. Lee, Creep rupture of the joint of a solid oxide fuel cell glass-ceramic sealant with metallic interconnect, J. Power Sources 245 (2014) 787e795. [18] C.-K. Lin, W.-H. Shiu, S.-H. Wu, C.-K. Liu, R.-Y. Lee, Interfacial fracture resistance of the joint of a solid oxide fuel cell glass-ceramic sealant with metallic interconnect, J. Power Sources 261 (2014) 227e237. [19] C.-K. Lin, Y.-A. Liu, S.-H. Wu, C.-K. Liu, R.-Y. Lee, Joint strength of a solid oxide fuel cell glass-ceramic sealant with metallic interconnect in a reducing environment, J. Power Sources 280 (2015) 272e288. [20] C.-K. Lin, T.-W. Lin, S.-H. Wu, W.-H. Shiu, C.-K. Liu, R.-Y. Lee, Creep rupture of the joint between a glass-ceramic sealant and lanthanum strontium manganite-coated ferritic stainless steel interconnect for solid oxide fuel cells, J. Eur. Ceram. Soc. 38 (2018) 2417e2429. [21] H. Apfel, M. Rzepka, H. Tu, U. Stimming, Thermal start-up behaviour and thermal management of SOFC’s, J. Power Sources 154 (2006) 370e378. [22] Y.-S. Chou, J.W. Stevenson, Compressive mica seals for solid oxide fuel cells, J. Mater. Eng. Perform. 15 (2006) 414e421. [23] F. Smeacetto, A. Chrysanthou, M. Salvo, T. Moskalewicz, F. D’Herin Bytner, L.C. Ajitdoss, M. Ferraris, Thermal cycling and ageing of a glass-ceramic sealant for planar SOFCs, Int. J. Hydrogen Energy 36 (2011) 11895e11903. [24] F. Smeacetto, A. Chrysanthou, T. Moskalewicz, M. Salvo, Thermal cycling of Crofer22APU-sealant-anode supported electrolyte joined structures for planar
C.-K. Lin et al. / Renewable Energy 138 (2019) 1205e1213 SOFCs up to 3000 h, Mater. Lett. 111 (2013) 143e146. [25] Wikipedia, Thermo-mechanical Fatigue, 2018. https://en.wikipedia.org/wiki/ Thermo-mechanical_fatigue. (Accessed 25 January 2018). [26] Y.-S. Chou, E.C. Thomsen, R.T. Williams, J.-P. Choi, N.L. Canfield, J.F. Bonnett, J.W. Stevenson, A. Shyam, E. Lara-Curzio, Compliant alkali silicate sealing glass for solid oxide fuel cell applications: thermal cycle stability and chemical compatibility, J. Power Sources 196 (2011) 2709e2716. [27] K.S. Weil, J.E. Deibler, J.S. Hardy, D.S. Kim, G.-G. Xia, L.A. Chick, C.A. Coyle, Rupture testing as a tool for developing planar solid oxide fuel cell seals, J. Mater. Eng. Perform. 13 (2004) 316e326. [28] Y.-T. Chiu, C.-K. Lin, Effects of Nb and W additions on high-temperature creep properties of ferritic stainless steels for solid oxide fuel cell interconnect, J. Power Sources 198 (2012) 149e157. [29] Y.-T. Chiu, C.-K. Lin, Thermo-mechanical fatigue properties of a ferritic stainless steel for solid oxide fuel cell interconnect, J. Power Sources 219 (2012) 112e119. [30] C.-K. Liu, T.-Y. Yung, K.-F. Lin, R.-Y. Lee, T.-S. Lee, Glass-ceramic sealant for planar solid oxide fuel cells, United States Patent No. 7,897,530 B2, 2011. [31] C.-K. Liu, T.-Y. Yung, K.-F. Lin, R.-Y. Lee, S.-H. Wu, High temperature glassceramic seals for SOFC applications, ECS Trans. 25 (2009) 1491e1500. [32] C.-K. Liu, R.-Y. Lee, K.-C. Tsai, S.-H. Wu, K.-F. Lin, Characterization and performance of a high-temperature glass sealant for solid oxide fuel cell, Ceram. Eng. Sci. Proc. 35 (2015) 65e75. [33] H.-T. Chang, C.-K. Lin, C.-K. Liu, High-temperature mechanical properties of a
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glass sealant for solid oxide fuel cell, J. Power Sources 189 (2009) 1093e1099. [34] H.-T. Chang, C.-K. Lin, C.-K. Liu, Effects of crystallization on the hightemperature mechanical properties of a glass sealant for solid oxide fuel cell, J. Power Sources 195 (2010) 3159e3165. [35] H.-T. Chang, C.-K. Lin, C.-K. Liu, S.-H. Wu, High-temperature mechanical properties of a solid oxide fuel cell glass sealant in sintered forms, J. Power Sources 196 (2011) 3583e3591. [36] C.-K. Lin, K.-L. Lin, J.-H. Yeh, W.-H. Shiu, C.-K. Liu, R.-Y. Lee, Aging effects on high-temperature creep properties of a solid oxide fuel cell glass-ceramic sealant, J. Power Sources 241 (2013) 12e19. [37] Z. Yang, Chemical interactions of barium-calcium-aluminosilicate-based sealing glasses with oxidation resistant alloys, Solid State Ionics 160 (2003) 213e225. [38] C.W.F.T. Pistorius, M.C. Pistorius, Lattice constants and thermal-expansion properties of the chromates and selenates of lead, strontium and barium, Z. Kristallogr. 117 (1962) 259e271. [39] G.V. Samsonov, The Oxide Handbook, second ed., IFI/Plenum Data Company, New York, 1982. [40] Material Data Sheet No. 4050 of Crofer 22 H, VDM Metals GmbH, Werdohl, Germany, 2010. [41] Y.-S. Chou, J.W. Stevenson, P. Singh, Effect of pre-oxidation and environmental aging on the seal strength of a novel high-temperature solid oxide fuel cell (SOFC) sealing glass with metallic interconnect, J. Power Sources 184 (2008) 238e244.
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Renewable Energy journal homepage: www.elsevier.com/locate/renene
Mechanical durability of solid oxide fuel cell glass-ceramic sealant/ steel interconnect joint under thermo-mechanical cycling Chih-Kuang Lin a, *, Kun-Yi Chen a, Si-Han Wu b, Wei-Hong Shiu b, Chien-Kuo Liu b, Ruey-Yi Lee b a b
Department of Mechanical Engineering, National Central University, Jhong-Li District, Tao-Yuan City 32001, Taiwan Division of Nuclear Fuels and Materials, Institute of Nuclear Energy Research, Lung-Tan District, Tao-Yuan City 32546, Taiwan
a r t i c l e i n f o
a b s t r a c t
Article history: Received 4 September 2018 Received in revised form 7 December 2018 Accepted 9 February 2019 Available online 13 February 2019
A testing method is developed to quantitatively determine the thermo-mechanical cycling life in oxidizing atmosphere for the joint between a solid oxide fuel cell glass-ceramic sealant and a ferritic stainless steel interconnect. Thermo-mechanical cycling tests are performed under cyclic shear or tensile loading in conjunction with cyclic temperature variance between 40 C and 800 C. Results reveal thermo-mechanical cycling life under both shear and tensile loadings increases with a decrease in end stress at 800 C, for a certain end stress at 40 C. Nevertheless, for a certain end stress at 800 C, the tensile thermo-mechanical cycling life increases with a decrease in end stress at 40 C, while the shear thermo-mechanical cycling life is independent of end stress at 40 C. A difference in fracture pattern is also observed between shear and tensile loadings. For shear loading, fracture mainly takes place along the interface between glass-ceramic sealant and an oxide layer, such as BaCrO4 or Cr2O3. However, for tensile loading, fracture mainly occurs within the glass-ceramic layer, following crack initiation at the interface of Cr2O3/sealant or Cr2O3/BaCrO4. © 2019 Elsevier Ltd. All rights reserved.
Keywords: Solid oxide fuel cell Glass-ceramic sealant Interconnect Joint Thermo-mechanical cycling
1. Introduction Solid oxide fuel cell (SOFC) is an energy conversion device which directly converts chemical energy into electricity by a series of electrochemical reactions. Compared to other fuel cells developed, SOFC generally has the highest efficiency of energy conversion and a better flexibility of fuel utilization (e.g. pure hydrogen and hydrocarbons) thanks to a high operating temperature [1]. Planar SOFCs (pSOFCs) have recently received more attention than tubular ones due to an easier fabricating process and a higher power density. For an anode-supported cell design with a thinner electrolyte, intermediate-temperature pSOFCs (IT-pSOFCs) can operate at a temperature between 600 C and 800 C due to a lower ohmic loss. A pSOFC stack is generally assembled with multiple unit cells which are connected in series by interconnects and bonded by sealants for a higher voltage and power in practical applications. Hermetic seals are required in joining SOFC components and must be stable at high temperature in both oxidizing and reducing
* Corresponding author. E-mail address: [email protected] (C.-K. Lin). https://doi.org/10.1016/j.renene.2019.02.041 0960-1481/© 2019 Elsevier Ltd. All rights reserved.
atmospheres to prevent leakage and fuel-oxidant mixing. Glass and glass-ceramic sealants have been extensively studied and practically used to seal SOFC components as they have a favorable combination of properties to satisfy the requirements [1]. Interconnects not only play a crucial role in electrical connection of repeated cells and separation of fuel and oxidant gas but also provide the necessary structural support. Reducing the operating temperature to 600 Ce800 C in IT-pSOFCs makes it possible to use metallic alloys, especially ferritic stainless steels, as interconnects for lower cost, better workability, higher electrical conductivity, and better mechanical properties [2,3]. Joining glassceramic sealant with metallic interconnect/frame is commonly seen in a rigid type of sealing for IT-pSOFC stack [4]. Although bonding characteristics such as chemical stability and interfacial compatibility are usually considered first for glassceramic sealant/metallic interconnect joint in IT-pSOFC stacks, structural integrity/durability is also an important issue because of thermal stresses generated in the stack. The major sources of thermal stress in a pSOFC stack during thermal cycles of operation include thermal gradient and mismatch of coefficient of thermal expansion (CTE) between adjacent components [5,6]. Therefore, study of mechanical strength and durability of sealant/interconnect
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joint is necessary for improvement in reliability and durability of pSOFC systems. Mechanical properties of SOFC glass-ceramic sealant/metallic interconnect joint have been studied in prior work [7e20]. However, most of those studies were mainly focused on the initial bonding strength and limited work was conducted on the long-term high-temperature creep behavior at steady operation stage [17,20]. In addition to consideration of creep mechanism acting on the SOFC sealant/interconnect joint at a hightemperature operating environment, effect of thermal cycling on the mechanical durability should not be ignored, either. Depending on application type, a pSOFC stack needs to survive possibly several to thousands of thermal cycles (including start-up, steady operation, and shutdown stages) during service life [21e24]. During periodic operation, thermal stresses change simultaneously and cyclically with temperature in each component of a pSOFC stack [5,6]. When an engineering component is subjected to a combined cyclic thermal and mechanical loading, it could suffer from thermomechanical fatigue (TMF) damage [25]. As such cyclic thermomechanical loading damage may accelerate degradation of the sealant/interconnect joint robustness, it is necessary to assess such effect on the structural integrity and cycle life of pSOFC stack. Effects of thermal cycling on the joining characteristics between glass-ceramic sealant and metallic interconnect have been investigated in a few studies [22e24,26,27], but only pure thermal cycling was applied in those studies without considering the synergistic effect of cyclic mechanical loading. For example, Smeacetto et al. [23,24] studied the effect of thermal cycling from room temperature (RT) to 800 C on interfacial stability and adherence between glass-ceramic sealant and metallic interconnect. In another study, Weil et al. [27] compared the rupture strength of glass-ceramic/metallic interconnect joints before and after thermal cycling from <70 C to 800 C. Their results [27] show that the decrease in seal strength of the joint specimens after a number of thermal cycles is attributed to the difference in CTE between crystalline phases and glassy phase and to the interfacial reaction products between glass-ceramic sealant and metallic interconnect. However, study on the mechanical durability of SOFC sealant/ interconnect joint under cyclic thermo-mechanical loading is still lacking. The aim of this study is thus to quantitatively assess the thermo-mechanical cycling (TMC) life in oxidizing atmosphere for the joint between a glass-ceramic sealant (GC-9) and an interconnect steel (Crofer 22 H). It is hoped that results of this study and previous studies [16e20] can provide a comprehensive understanding of the mechanical behavior of such a joint for better design of a reliable and durable IT-pSOFC stack.
stainless steel, Crofer 22 H (ThyssenKrupp VDM GmbH, Werdohl, Germany), which is developed for IT-pSOFC interconnect. Chemical composition and mechanical properties of the given interconnect steel have been studied previously [28,29]. Metallic coupons of the joint specimens are prepared and machined in a way similar to that described in Ref. [16]. A borosilicate glass (designated as GC-9) for pSOFC sealing is applied to joining the two steel coupons in each joint specimen. The major composition in mol% of this glass-ceramic sealant includes 34 BaO, 34 SiO2, 12 CaO, 9.5 B2O3, 5 La2O3, 4.5 Al2O3, and 1 ZrO2 [30]. GC-9 has proved to be suitable for application in pSOFCs thanks to a good combination of thermal properties, chemical compatibility and stability, and hermetic properties [31,32]. Mechanical properties of GC-9 sealant have been investigated in prior work [33e36]. As the sandwich joint specimens are prepared and fabricated following the same way given in a previous study, detailed specimen preparation procedures are described in that study [16]. 2.2. Thermo-mechanical cycling test Previous analyses of thermal stress in a pSOFC stack indicate that the thermal stress generated during periodic operation varies with temperature and location in a component [5,6]. The maximum stress at a certain position at RT is greater than that at hightemperature operation stage [5,6]. It is due to a larger temperature difference between RT and the initial stress-free assembling temperature (800 C) of the given pSOFC stack [5,6]. In a pSOFC stack, glass-ceramic sealants and metallic interconnects are joined together for tight-sealing at high temperature. During joining at high temperature, glass-ceramic is soft enough to create an initial “stress free” condition in the joints. After cooling, thermal stresses are generated within the joints as a result of CTE mismatch and temperature difference/gradient. Such a case in which thermal stress varies cyclically with temperature at a certain position in a pSOFC stack is similar to TMC. To simulate such TMC experienced by the sealant/interconnect joints in a pSOFC stack, a TMC pattern (Fig. 1) is applied to the given sandwich joint specimens. In each TMC test performed in air, cyclic mechanical loading in conjunction with cyclic temperature variance is applied using a commercial closed-loop servo-hydraulic material test machine attached with a furnace. This TMC testing technique has been developed and employed in studying the TMF behavior of Crofer 22 H steel such that details of the testing equipment and procedures can be found
2. Experimental procedures 2.1. Materials and specimens As compressive loads normal to the interface of a joint are not expected to cause interfacial cracking or debonding, only shear and tensile loadings are applied for cyclic thermo-mechanical test in the present study. In this regard, two types of thin-plate sandwich joint specimen (metal/glass-ceramic sealant/metal) with a thickness of 2.5 mm are designed and made, one for shear loading test and the other for tensile loading test. Pin holes are drilled at both ends of each specimen for applying pin loading to minimize bending and twisting effects during mechanical testing. Tensile or shear stress can be generated at the interface of metal/glass-ceramic joint by applying a uni-axial tensile loading. Such specimen designs have been developed for assessing the bonding strength of a similar metal/glass-ceramic joint in a previous study [16] in which detailed specimen geometry and dimensions can be found. The metallic parts of the joint specimens are made of a commercial ferritic
Fig. 1. Schematic diagram of thermo-mechanical loading applied in TMC test.
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in Ref. [29]. As shown in Fig. 1, a cyclic temperature range of 40 Ce800 C with a heating and cooling rate of 5 C min1 is applied to simulating the temperature range between shutdown and steady operation stages in a typical pSOFC system. As restricted by the cooling ability of the furnace, it takes 60 min to cool down from 100 C to 40 C in furnace cooling. Therefore, the total time for one period of TMC is 352 min (5.87 h). Together with the temperature change, the cyclic mechanical loading in Fig. 1 is applied under force control. Thermal stress also varies with location of the sealant/ interconnect joint in periodic operation of SOFC [5,6]. Therefore, several stress levels of specified joint strength ratios (JSRs) at 40 C and 800 C are selected in performing the TMC tests to simulate various combinations of thermal stresses experienced in the joints of a pSOFC stack in periodic operation. JSR is defined as the applied peak or valley stress divided by the corresponding bonding strength at that specific peak or valley temperature. In other words, JSR is a normalized stress level. JSRs of 0.2, 0.4, and 0.6 are selected as the peak or valley stress of mechanical loading applied at 40 C and 800 C in each cycle. In the following, (x, y) is used to represent the applied end stresses in a TMC test. x is the JSR corresponding to the end stress applied at 40 C while y is the JSR corresponding to the end stress applied at 800 C. For example, a (0.4, 0.2) shear TMC loading indicates a shear stress of 2.64 MPa (JSR ¼ 0.4) is applied at 40 C and then changes bi-linearly with temperature to reach a level of 0.94 MPa (JSR ¼ 0.2) at 800 C in a half cycle and vice versa for the other half cycle, as shown in Fig. 1. Note that the average shear bonding strength of the given joint is of 6.6 MPa at RT and 4.7 MPa at 800 C, while the tensile counterpart has a value of 23 MPa at RT and 12.7 MPa at 800 C [16]. As shown in Fig. 1, each TMC test starts from 800 C. Note that in some TMC tests the end stress applied at 40 C is greater than that applied at 800 C and vice versa for the others. Note that each line segment in Fig. 1 represents a constant stress rate or thermal rate through appropriate force and temperature control such that both temperature and stress reach the end levels at the same time. 2.3. Fractography analysis Fracture surfaces of broken specimens after TMC test were observed using optical microscopy (OM) and scanning electron microscopy (SEM). For composition analysis of elemental distribution on the fracture surfaces, an energy dispersive spectrometer (EDS) module attached in the SEM was employed. X-ray diffraction (XRD) was applied to determining the crystalline phases. 3. Results and discussion 3.1. Thermo-mechanical cycling under shear loading Table 1 lists the number of cycles to failure for joint specimens subjected to various shear TMC loadings. Note that the TMC test was terminated if the specimen was not ruptured at 50 or 60 cycles.
Table 1 Number of cycles to failure for various shear TMC loadings. Stress (JSR) applied at 40 C
Number of cycles to failure Stress (JSR) applied at 800 C
1.32 MPa (0.2) 2.64 MPa (0.4) 3.96 MPa (0.6) a
Runout test.
0.94 MPa (0.2)
1.88 MPa (0.4)
2.82 MPa (0.6)
>50a >60a >50a
6 8 6
<1 <1 1
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Accordingly, it is considered a runout test and the TMC life is expressed as > 50 or >60 cycles in Table 1. As shown in Table 1, the number of cycles to failure under shear cyclic loading is increased with a decrease in JSR applied at 800 C, with a given JSR applied at 40 C. However, the end stress applied at 40 C does not affect the shear TMC life significantly as the TMC life is comparable for a certain JSR applied at 800 C regardless of the JSR applied at 40 C. Therefore, shear TMC life is controlled by the end stress applied at 800 C and may be considered an accumulated duration of creep loading at 800 C. As a combination of high temperature environment and mechanical stress may generate creep damage, the TMC life is compared with the estimated creep rupture time through a creep stress-life relation given in a previous study [17]. To make a comparison of TMC life with accumulated duration of creep loading, the estimated creep rupture time is calculated by the creep stress-life relation determined for a similar Crofer 22 H/GC-9/Crofer 22 H joint tested at 800 C [17]. The creep stress-life relation under constant shear loading at 800 C is expressed as follows [17],
Shear creep loading : ttr0:066 ¼ 1:68
(1)
where t is the applied shear stress in unit of MPa and tr is time to rupture in unit of h. As the creep stress-life relation above is obtained under constant shear loading at 800 C, the estimated creep rupture time for each shear TMC test is calculated by substituting into Eq. (1) the end stress applied at 800 C. In consideration of deviation in temperature control during mechanical test in the present study and Ref. [17], a small temperature range of 795 Ce800 C around the peak temperature is selected and considered for generating a similar creep loading effect at 800 C. In this way, the sample is assumed to experience a similar creep loading effect of 800 C for 2 min at 795 Ce800 C in each thermo-mechanical loading cycle, due to a heating and cooling rate of 5 C min1. In other words, a TMC cycle is equivalent to 2 min of creep loading at 800 C. Table 2 shows the accumulated duration of creep loading at 795 Ce800 C for various shear TMC loadings and the corresponding estimated creep rupture time. According to Eq. (1), the shear stress applied at 800 C for a creep rupture time over 1000 h is about 1 MPa which is slightly higher than 0.94 MPa. It is thus expected that the shear TMC life for an end stress of 0.94 MPa (0.2 JSR) applied at 800 C is much longer than 60 cycles (352.2 h). For shear loading of 0.4 JSR (1.88 MPa) applied at 800 C combined with various JSRs applied at 40 C, the TMC life is of 6e8 cycles which is equivalent to a duration of 12e16 min of creep loading at 795 Ce800 C. An estimated creep rupture time by Eq. (1) for a shear stress of 1.88 MPa applied at 800 C is about 12 min which is comparable with the accumulated duration of creep loading at 795 Ce800 C for shear TMC tests with 0.4 JSR (1.88 MPa) applied at 800 C. For a shear TMC test with 0.6 JSR (2.86 MPa) applied at 800 C, the estimated creep rupture time is 0.03 min, so the TMC test is expected not to run more than 1 cycle. As shown in Table 2, the accumulated duration of creep loading in shear TMC test is in good agreement with the estimated creep rupture time. It indicates that creep mechanism at 800 C may play an important role in determining the TMC life under shear loading. Fig. 2 shows the representative failure patterns of TMC under shear loading. Note that in Fig. 2 and the following similar figures, fracture surfaces on both sides of each broken specimen are presented with mirror symmetry. Fig. 2(a) shows the fracture pattern for TMC life of 6 cycles under (0.6, 0.4) shear loading. SEM micrographs of two selected areas in the lower fracture surface of Fig. 2(a), one at the peripheral region and the other at the central region, are shown in Fig. 3. Note that on each fracture surface shown in Fig. 2, the peripheral region and the central region are
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Table 2 Accumulated duration of creep loading at 795 Ce800 C for various shear TMC loadings. Stress (JSR) applied at 40 C
Accumulated duration of creep loading at 795 Ce800 C (min) Stress (JSR) applied at 800 C
1.32 MPa (0.2) 2.64 MPa (0.4) 3.96 MPa (0.6) Estimated creep rupture time by Eq. (1) a
0.94 MPa (0.2)
1.88 MPa (0.4)
2.82 MPa (0.6)
>100a >120a >100a 404,098
12 16 12 11.9
<2 <2 2 0.03
Runout test.
(a)
(b) Fig. 2. Representative failure patterns in shear specimens of various TMC lives: (a) 6 cycles; (b) < 1 cycle.
separated by a dash-line loop. SEM micrograph of a peripheral yellow area is shown in Fig. 3(a) and identified as a BaCrO4 layer on top of GC-9 glass-ceramic substrate. BaCrO4 has been found to exhibit a yellow color in a previous study [17]. Existence of BaCrO4 in Fig. 3(a) is also confirmed by EDS analysis, as peaks of Ba, Cr, and O are detected in the EDS spectrum (Fig. 4). Note that in the EDS spectrum of Fig. 4, Pt is also detected as it is sputtered on the fracture surface to increase conductivity for SEM observation. As needle-shape crystalline phases (a-Ba(Al2Si2O8)) of GC-9 glassceramic are also observed in Fig. 3(a) and Si, Al, and Ca elements are detected by EDS (Fig. 4), GC-9 glass-ceramic also exists in the micrograph of Fig. 3(a). Note that a-Ba(Al2Si2O8) is the main crystalline phase in GC-9 glass-ceramic [35]. Therefore, the yellow peripheral region in the lower fracture surface of Fig. 2(a) is a BaCrO4 layer mixed with GC-9 glass-ceramic. In Fig. 3(b), needleshape a-Ba(Al2Si2O8) phases are observed and Cr2O3 is detected by EDS analysis, indicating the light green central region in the lower fracture surface of Fig. 2(a) is a Cr2O3 layer mixed with GC-9
glass-ceramic. To further confirm existence of BaCrO4, Cr2O3, and aBa(Al2Si2O8) phases, XRD is employed to analyze the phase structure in the lower fracture surface of Fig. 2(a). As shown in Fig. 5, the diffraction peaks match those of BaCrO4, Cr2O3, and a-Ba(Al2Si2O8) phases. SEM micrographs of two selected areas in the upper fracture surface of Fig. 2(a), one at the peripheral region and the other at the central region, are shown in Fig. 6. Note the micrographs shown in Fig. 6 are the counterparts of Fig. 3. Based on EDS analysis results and existence of needle-shape crystalline phases in Fig. 6(a), the peripheral region in the upper fracture surface of Fig. 2(a) is a thin peeled BaCrO4 layer mixed with GC-9 glass-ceramic. Similarly, EDS analysis results and existence of chromia structure in Fig. 6(b) indicate the central region in the upper fracture surface of Fig. 2(a) is a peeled Cr2O3 layer mixed with GC-9 glass-ceramic. As the joining process of the sandwich joint specimen was conducted in air, a BaCrO4 layer is formed along the edge of joining area. However, as the central region is far from the edge, oxygen from the ambient air is not easy to diffuse into the deep interior. Therefore, chromate layer is barely observed in the central part of the joining area. This observation is in agreement with that in the study of Yang [37]. Barium chromate is produced by the reaction of oxygen with BaO and Cr2O3 [37]. The fracture surfaces of (0.4, 0.6) shear TMC loading with a life of <1 cycle are shown in Fig. 2(b). Its failure pattern is almost the same as that of the sample under (0.6, 0.4) shear TMC loading (Fig. 2(a)), indicating a similar fracture mechanism. The only difference between them is that a greater amount of BaCrO4 (a larger yellow peripheral region) is observed on the fracture surface of (0.6, 0.4) sample (Fig. 2(a)). Apparently, for a longer testing time in air, more BaCrO4 is formed on the joint interface at the peripheral region. Based on the fractography observations described above for the joint samples under shear TMC loadings, cracks were initiated at the BaCrO4/GC-9 interface in the periphery, followed by propagation and final fracture at the Cr2O3/GC-9 interface in the central joining area, as shown in Fig. 2. Apparently, shear fracture prefers to take place at the interfaces between chromate, chromia, and glassceramic layers due to a larger CTE mismatch. Note that the CTE values for BaCrO4, GC-9 glass-ceramic, Cr2O3, and Crofer 22 H alloy are of 20.4 106 C1 (at 20e813 C) [38], 13.1 106 C1 (at 20e650 C) [32], 9.6 106 C1 (at 20e1400 C) [39], and 11.8 106 C1 (at 20e800 C) [40], respectively. A large thermal mismatch could generate thermal stress and result in formation of defects and/or microcracks which would serve as the fracture origin and/or provide a weak path for cracking [41]. This is why cracking begins at the BaCrO4/GC-9 interface in the peripheral region and switches to the Cr2O3/GC-9 interface in the central region of the joining area when subjected to shear TMC loading.
3.2. Thermo-mechanical cycling under tensile loading Table 3 lists the number of cycles to failure for joint specimens
C.-K. Lin et al. / Renewable Energy 138 (2019) 1205e1213
(a)
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(b)
Fig. 3. SEM micrographs of fracture surface in the lower part of Fig. 2(a): (a) peripheral region (BaCrO4 with GC-9); (b) central region (Cr2O3 with GC-9).
cycles to failure increases with a decrease in applied JSR at 40 C. It indicates that the tensile TMC life is controlled not only by the end stress applied at 800 C but also by that applied at 40 C. This is different from the trend of shear TMC life which is dominated only by the JSR applied at 800 C. As described in the following failure analysis, fracture mainly takes place within the glass-ceramic layer in tensile specimens. Glass-ceramic is more sensitive to tensile loading at low temperature due to its brittle behavior. Therefore, the JSR applied at 40 C also plays a certain role in determining the tensile TMC life. Similar to shear loading, the estimated creep rupture time for a given constant tensile stress at 800 C can be calculated through the creep stress-life relation determined in the previous study [17]. The creep stress-life relation under constant tensile load at 800 C in air is expressed as follows [17],
tensile creep loading : str0:074 ¼ 1:85 Fig. 4. EDS analysis result of Fig. 3(a).
Fig. 5. XRD pattern for the lower fracture surface of Fig. 2(a).
subjected to tensile TMC loadings. It shows that the number of cycles to failure under tensile TMC is decreased with an increase in JSR applied at 800 C, with a given JSR applied at 40 C. For a given tensile stress of 2.54 MPa (0.2 JSR) applied at 800 C, the number of
(2)
where s is the applied tensile stress in unit of MPa and tr is time to rupture in unit of h. Table 4 lists the accumulated duration of creep loading at 795 Ce800 C for joint specimens under various tensile TMC loadings. The bottom row in Table 4 is the estimated creep rupture time calculated by Eq. (2). Unlike the shear loading results, the accumulated duration of tensile creep loading at 795 Ce800 C in experiment is not close to the estimated creep rupture time, in particular for JSR of 0.2 applied at 800 C (Table 4). It reveals that the creep effect at 800 C does not play a dominating role in determining the tensile TMC life. Therefore, a true cycling effect is responsible for the TMC failure under tensile loading. Representative failure patterns in tensile specimens with various TMC lives are shown in Fig. 7. Fig. 7(a) exhibits a fracture pattern for TMC life of <1 cycle under (0.4, 0.6) tensile loading. SEM micrographs of two selected areas on the fracture surface of Fig. 7(a), one from the heavy green region and the other from the white region, are shown in Fig. 8. As shown in Fig. 8(a), chromia features are found and confirmed by EDS analysis for the heavy green region in Fig. 7(a). As shown in Fig. 8(b), needle-shape crystalline phases (a-Ba(Al2Si2O8)) are found in the white region of Fig. 7(a) indicating fracture occurred within the glass-ceramic layer. Therefore, for the sample under (0.4, 0.6) tensile TMC loading, cracking started peripherally at the Cr2O3/GC-9 interface, and then proceeded inward into the glass-ceramic layer.
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(a)
(b)
Fig. 6. SEM micrographs of fracture surface in the upper part of Fig. 2(a): (a) peripheral region (BaCrO4 with GC-9); (b) central region (peeled Cr2O3 with GC-9).
3.3. Overall comparison
Table 3 Number of cycles to failure for various tensile TMC loadings. Stress (JSR) applied at 40 C
Number of cycles to failure
Stress (JSR) applied at 800 C
4.6 MPa (0.2) 9.2 MPa (0.4) 13.8 MPa (0.6)
2.54 MPa (0.2)
5.08 MPa (0.4)
7.62 MPa (0.6)
11 3 1
<1 <1 <1
<1 <1 <1
Fig. 7(b) shows the fracture pattern for TMC life of 11 cycles under (0.2, 0.2) tensile loading. Similarly, by means of EDS analysis and microstructural observation, the heavy green and white regions in Fig. 7(b), again, are identified as Cr2O3 and GC-9, respectively. However, additional yellow regions are observed at the bottom edges of both fracture surfaces presented in Fig. 7(b). Note that these yellow regions are the counterparts of the heavy green regions at top edges of the two mirror-symmetric fracture surfaces of Fig. 7(b). By means of EDS analysis, elements O, Si, Cr and Ba are detected and needle-shape crystalline phases of GC-9 are observed in the yellow regions, as shown in Fig. 9. It indicates the yellow regions at the bottom edges of Fig. 7(b) are BaCrO4 mixed with GC9. Therefore, for tensile TMC loading with a life of 11 cycles, crack was initiated at the peripheral interface between Cr2O3 and BaCrO4 layers and subsequently penetrated into the GC-9 layer. As the testing time of (0.2, 0.2) tensile TMC loading is longer than that of (0.4, 0.6), i.e. 11 cycles vs. < 1 cycle, it is expected more BaCrO4 is formed at the periphery of joining area. Consequently, fracture of (0.2, 0.2) specimen originated at the Cr2O3/BaCrO4 interface rather than at the Cr2O3/GC-9 interface.
As described in previous sections, TMC life under both shear and tensile loadings is decreased with an increase in end stress (JSR) applied at 800 C, with a given end stress applied at 40 C. However, given an end stress applied at 800 C, the tensile TMC life is increased with decreasing end stress applied at 40 C, while the shear TMC life is independent of end stress applied at 40 C. As shown in Tables 1 and 3, for a given combination of cyclic end stresses (x, y), the number of cycles to failure under tensile TMC loading is significantly smaller than the shear counterpart. It reveals that the given Crofer 22 H/GC-9/Crofer 22 H joint suffers from tensile loading to a greater extent than that from shear loading, with regard to TMC effect on structural integrity. Such a difference in the trend of TMC life between shear and tensile loadings also appears in their failure patterns. As shown in Fig. 2, fracture of TMC under shear loading takes place initially at the BaCrO4/GC-9 interface in the peripheral region and then proceeds along the Cr2O3/GC-9 interface in the interior of joining area. For fracture of TMC under tensile loading, cracking starts peripherally at the Cr2O3/GC-9 interface (<1 cycle) or at the Cr2O3/BaCrO4 interface (11 cycles), and then penetrates inward into the GC-9 glass-ceramic layer, as shown in Fig. 7. As glass-ceramic becomes viscous at temperature above its glass transition temperature, viscous deformation and relevant damage, such as creep, are more susceptible to shear loading. Therefore, development of cracking in shear TMC loading presumably takes place mostly in the high-temperature range of a thermo-mechanical loading cycle. As a result, the shear TMC life is controlled by the end stress applied at 800 C, irrespective of the end stress applied at 40 C. Note the glass transition temperature and softening temperature of GC-9 glass-
Table 4 Accumulated duration of creep loading at 795 Ce800 C under various tensile TMC loadings. Stress (JSR) applied at 40 C
Accumulated duration of creep loading at 795 Ce800 C (min) Stress (JSR) applied at 800 C
4.6 MPa (0.2) 9.2 MPa (0.4) 13.8 MPa (0.6) Estimated creep rupture time by Eq. (2)
2.54 MPa (0.2)
5.08 MPa (0.4)
7.62 MPa (0.6)
22 6 2 0.88
<2 <2 <2 <0.1
<2 <2 <2 <0.1
C.-K. Lin et al. / Renewable Energy 138 (2019) 1205e1213
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(a)
Fig. 9. SEM micrograph of the yellow region at bottom edge of the lower fracture surface in Fig. 7(b).
(b) Fig. 7. Representative failure patterns in tensile specimens of various TMC lives: (a) < 1 cycle; (b) 11 cycles.
ceramic are of 668 C and 745 C, respectively [33]. As glass-ceramic itself is a brittle material at low temperature, it is more vulnerable to tensile stress in terms of fracture resistance. Therefore, cracking also takes place within the GC-9 layer under tensile TMC loading. Therefore, TMC damage develops during the entire tensile thermomechanical loading cycle. In this regard, tensile TMC life is dependent on both end stresses applied at 800 C and 40 C. This is consistent with the aforementioned failure patterns in which shear TMC failure all takes place at the GC-9/oxide interfaces while the tensile TMC fracture occurs mainly within the GC-9 layer in addition to initiation at the interfaces. It is required that glass-ceramic sealant for SOFC applications
(a)
provide a good gas tightness during sealing and operating processes at high temperatures. Leak rate for a comparable Crofer 22 APU/GC-9/Crofer 22 APU joint was measured under pure thermal cycling (RTe800 C) in a previous study [32]. Note that Crofer 22 APU is a former generation of Crofer 22 H alloy for SOFC interconnect, with addition of Nb and W in the latter. For the sealed Crofer 22 APU/GC9/Crofer 22 APU coupon subjected to 50 repetitions of pure thermal cycling from RT to 800 C, the leakage rate during the test kept a low value which is well below the allowable leak rate limit [32]. It indicates the GC-9 glass-ceramic performs well in sealing metallic interconnects and/or frames for pSOFC stacks when subjected to pure thermal cycling. However, results of the present work reveal that combining thermal cycling with cyclic mechanical loading could affect the mechanical durability of the sealed joint and deteriorate the sealing performance under certain TMC conditions. For example, the given Crofer 22 H/GC-9/Crofer 22 H joint could survive more than 50 cycles under a (0.2, 0.2), (0.4,
(b)
Fig. 8. SEM micrographs of fracture surface in Fig. 7(a): (a) heavy green region (Cr2O3); (b) white region (GC-9).
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0.2), or (0.6, 0.2) shear TMC loading, but would lose its structural integrity after 11 or a smaller number of cycles under a (0.2, 0.2), (0.4, 0.2), or (0.6, 0.2) tensile TMC loading. The pure thermal cycling tests carried out in previous studies [22e24,26,27,32] are suitable for assessing the long-term chemical stability and interfacial compatibility with adjacent components for SOFC sealants. However, as working temperature and thermal stress change simultaneously in a working cycle of SOFC system, it is thus suggested that TMC test should be applied in quantitatively assessing the mechanical durability of metal/sealant/metal joints for design of a durable pSOFC stack. Previous studies have proved that GC-9 glass-ceramic is suitable for SOFC sealing application as it shows good thermal properties, chemical stability and compatibility, hermetic properties, and mechanical properties [30e36]. The TMC testing method developed in the present study is able to quantitatively determine the mechanical durability of SOFC sealant/interconnect joints in oxidizing atmosphere, which has not yet been reported in the literature. However, the interaction between glass-ceramic sealant and metallic interconnect in oxidizing atmosphere may differ from that in reducing atmosphere which may influence the TMC behavior in a different manner. Therefore, a counterpart study in a reducing environment is being conducted to investigate the effect of reducing atmosphere on the TMC behavior of the given GC-9/ Crofer 22 H joint, which is considered an extension of the present study and the results will be published in the near future. The TMC testing results of sealed Crofer 22 H/GC-9/Crofer 22 H coupons presented in this study in conjunction with other mechanical properties such as bonding strength, creep strength, and interfacial fracture resistance obtained in prior work [16e19] could provide a comprehensive mechanical data base for design and development of reliable pSOFC stacks using GC-9 glass-ceramic sealant.
4. Conclusions (1) For the given joint of GC-9 glass-ceramic sealant/Crofer 22 H steel tested in oxidizing atmosphere, shear TMC life is increased with a decrease in end stress applied at 800 C. However, it is independent of end stress applied at 40 C. The accumulated duration of creep loading at 795 Ce800 C in shear TMC test is comparable with the estimated creep rupture time. It reveals that creep effect plays an important role in determining the shear TMC life. (2) The tensile TMC life is affected by both end stresses applied at 800 C and 40 C, showing an increasing life with decreasing end stresses applied. The accumulated duration of creep loading at 795 Ce800 C in tensile TMC test does not match the estimated creep rupture time. It indicates creep effect is not the sole dominating factor in determining tensile TMC life. (3) Fracture of TMC under shear loading takes place initially at the BaCrO4/GC-9 interface in the peripheral region and then proceeds along the Cr2O3/GC-9 interface in the interior. For fracture of TMC under tensile loading, cracking starts peripherally at the Cr2O3/GC-9 or Cr2O3/BaCrO4 interface, and then penetrates inward into the GC-9 glass-ceramic layer. (4) For the given GC-9/Crofer 22 H joint in oxidizing atmosphere, a difference in the trend of TMC life between shear and tensile loadings is observed. This is consistent with the difference found in failure patterns in which the shear TMC failure takes place at the GC-9/oxide interfaces while the tensile TMC fracture occurs mainly within the GC-9 layer in addition to initiation at the interfaces.
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