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Stress analysis of the brazing joints of tubular ceramic oxygen-permeable membranes and metal supports
Ceramics International 45 (2019) 1545–1553
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Ceramics International journal homepage: www.elsevier.com/locate/ceramint
Stress analysis of the brazing joints of tubular ceramic oxygen-permeable membranes and metal supports ⁎
Jialu Zhang, Jian Zhang, Lanqian Li, Chao Zhang, Yuwen Zhang , Xionggang Lu
T
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State Key Laboratory of Advanced Special Steel & Shanghai Key Laboratory of Advanced Ferrometallurgy & School of Materials Science and Engineering, Shanghai University, Shanghai 200444, China
A R T I C LE I N FO
A B S T R A C T
Keywords: Stress analysis Brazing joints Tubular ceramic oxygen-permeable membrane Nonlinear material properties
This study investigates the effect of joint structure and temperature dependent mechanical properties on the residual stress states in ceramic/metal brazing joints. A closed-one-end BaCo0.7Fe0.2Nb0.1O3 (BCFN) membrane tube is joined to 310S by using silver filler. Finite element simulation of different BCFN/310S joint assemblies is conducted to evaluate the stresses and deformations of the components at an isothermal state. The results indicate that the outer sleeve structure with low residual stress and good air tightness is more suitable for joining ceramics to metal supports than other joint structures in four joint assemblies. And thermal cycling experiments also verify the functional reliability of the outer sleeve joint assembly. Besides, nonlinear material properties of BCFN can not only cause the sudden change in ceramic stress at 300 °C, but also be the important basis for the stress contrast of different joint structures.
1. Introduction Mixed-conducting ceramic oxygen-permeable membranes show the attractive applications in pure oxygen production, membrane reactors for catalytic reforming, and oxy-fuel combustion [1,2]. To develop a commercial process, the large dimension membrane modules with gastight sealing to their metal supports must be constructed [3]. The critical issue of fabricating high-performance joints is the development of the proper joint structure and technique for joining membranes to metal supports. Metal brazing has been paid more and more attention for joining the oxygen-permeable membranes to metal supports. Without the use of fluxes and other special atmosphere, reactive air brazing (RAB) has been investigated extensively in the field of sealing the oxygenpermeable membranes. Due to the ductility and oxidation resistance, Ag-based alloys have been used as metallic sealing fillers. The oxygenpermeable membranes, such as (La0.6Sr0.4)(Co0.2Fe0.8)O3-δ (LSCF) [4,5], BaCo0.4Fe0.4Zr0.2O3-δ (BCFZ) [6], BaCo0.7Fe0.2Nb0.1O3-δ (BCFN) [7], Ba0.5Sr0.5Co0.8Fe0.2O3-δ (BSCF) [8], were used in these brazing experiments. Many studies on RAB have been conducted on the wettability, interfacial reactions and joining strength between the ceramic membranes and the brazes [4–11]. In RAB, the interfacial reactive can improve the wettability, adhesion and bonding strength. RAB was also used to join the tubular membranes to metal supports in high-
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temperature resistance furnaces [12] or by an induction heating method [13]. Due to the mismatch of the thermal expansion of the ceramic membranes and metal supports, the joining stresses are very high and the membranes are easily destructed. In order to minimize joining stresses, it is necessary to optimize the joint structures. A combined model of the process parameters with the simulation of the material behavior would be very helpful for the structure optimization and the reliability assessment of the membrane modules. Compared with the extensive study on the stress simulation for SOFC [14–19], there are a few of reports on the stress analysis for joining tubular mixed-conducting oxygen-permeable ceramic to its metal support. Ragnar Kiebach investigated the residual stress of BSCF/Haynes 214 butt and lap joint assemblies [12]. The stress of the lap joint geometry is larger. Dabbarh et al. analyzed the mechanical reliability of BSCF/Alloy 310Si joints by the wetting experiments and stress simulation. They found that the stress could be reduced by changing from the butt joint to the sleeve joint [20]. Kirsten Bobzin calculated the residual stress of BSCF/AISI 314 joint and BSCF/Crofer 22H joint by using Abaqus and established a visco-plastic model to analyze the effect of silver creep on the residual stress [11]. The similar TECs of the components are useful to reduce the residual stress and the solder creep can also release the stress. Obviously, there are some discrepancies on the simulation and optimization of the joint structures from the different reports. The variation in the sizes of the solder and metal
Corresponding authors E-mail addresses: [email protected] (Y. Zhang), [email protected] (X. Lu).
https://doi.org/10.1016/j.ceramint.2018.10.028 Received 28 September 2018; Received in revised form 2 October 2018; Accepted 3 October 2018 Available online 10 October 2018 0272-8842/ © 2018 Elsevier Ltd and Techna Group S.r.l. All rights reserved.
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supports has not been analyzed and optimized in these studies. Especially, the ceramic properties used in these reports were usually considered as the constant or linear values as a function of temperature. In fact, with the temperature changing, the physicochemical and mechanical properties of mixed-conducting ceramic oxygen-permeable membranes are nonlinear and complicated [2,13,21]. The anomalies in the properties may cause damage of the membranes, and special consideration should be given to the design and the operation of the membrane reactors. The aim of this study was to seek an optimized joint structure for joining the tubular membranes to their metal supports by conducting a systematic stress simulation. The brazing joints of tubular ceramic oxygen-permeable membranes BCFN and 310S stainless steel supports (310S) with different geometric structures were analyzed and compared. COMSOL software was utilized to model and analyze the stresses and strains that arise in the mismatch of the joining components. Nonlinear physicochemical and mechanical characteristics of mixedconducting BCFN membrane material were used and paid more attention in the stress modeling and analysis.
Fig. 2. Simplified diagram of BCFN ceramic membrane tube.
(965 °C). Then, the residual stresses would appear in the assemblies during cooling down from high temperature (965 °C) to room temperature (25 °C) because of different TECs. The residual stresses are calculated by a physical simulation software COMSOL without considering the influence of creep and the elastic-plastic deformation being governed by the braze model. Simultaneously, the total strain of the ceramic membrane can be decomposed into elastic strain and thermal strain. And the strain in ceramic can be described in the Eqs. (1) and (2), where D is an elastic matrix, which is related with Young's modulus and Poisson's ratio. The thermal strain mainly derived from the differences of TEC and temperature.
2. Modeling and experiments 2.1. Geometric model To join the membrane tubes to their metal supports suitably, four geometric structures of the butt and sleeve joints are designed as shown in Fig. 1: (1) butt joint without brazing meniscus (Figs. 1(a)), (2) butt joint with brazing meniscus (Figs. 1(b)), (3) inner sleeve joint (Figs. 1(c)) and (4) outer sleeve joint (Fig. 1(d)). During the simulation process, a closed-one-end BCFN membrane tube with an outer diameter of 10 mm and a thickness of 1 mm are used. Similarly, silver rings with different thickness of 0.6 mm, 0.35 mm and 2 mm are prepared for the butt joint, inner sleeve joints and outer sleeve joints respectively. The simplified diagram of ceramic tube is shown in the Fig. 2, whose bottom part (A1B1, A2B2) represents the brazing part in sleeve joint. Only a half axial section was considered due to the rotationally symmetry of the sealing part along the vertical Z axis. The finite element meshes were generated automatically by the simulation software, as shown in Fig. 3. Notably, the maximum mesh sizes of four kinds of assemblies are same. The paths A1B1 (Fig. 3(c)) and A2B2 (Fig. 3(c)) represent the Ag/BCFN interfaces of the inner sleeve joint and the outer sleeve joint respectively.
σ = Dεel + σ0
(1)
εel = ε − ε0 − εth
(2)
2.3. Material properties The parameters of the properties of silver and 310S are provided by COMSOL Material Library. The parameters of the BCFN properties are characterized by our group. Fig. 4 demonstrates the properties of the different materials used in the simulation. And it is clear that the variation of BCFN material properties is nonlinear with the temperature varying, especially its TEC and heat capacity. 2.4. Boundary and constrained conditions The two-dimensional axisymmetric boundary conditions were applied in the axial sections. The Ag/BCFN interface and Ag/310S interface were defined as the identity boundary pairs. All the nodes on the
2.2. Simulation procedure At first, all assemblies are in a stress-free state at initial temperature
Fig. 1. Cross-sectional schematics of (a) the butt joint assembly without brazing meniscus, (b) the butt joint assembly with brazing meniscus, (c) the inner sleeve joint assembly and (d) the outer sleeve joint assembly. 1546
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Fig. 3. Finite element meshing of (a) the butt joint assembly without brazing meniscus, (b) the butt joint assembly with brazing meniscus, (c) the inner sleeve joint assembly and (d) the outer sleeve joint assembly.
was heated from room temperature to brazing temperature (965 °C) within 12 min and then was directly cooled to 900 °C within 1 min. With the subsequent cooling following, a cooling rate of 20 °C/min was taken during 900–600 °C and a cooling rate of 10 °C/min for 600–300 °C step. Below 300 °C, the assembly was cooled down to room temperature. To examine the reliability of the optimized joints, thermal cycle testing was conducted. The integrity of the joints was tested using a setup described in Ref. [8]. The leakage of N2 in the air into the sweeping gas (Ar) was continuously monitored by an online mass spectrometer (Hidden HPR20) to confirm that the membrane tube and the seal remained intact. The brazed module was heated to 875 °C, holding at 875 °C for 30 min, and cooling to 200 °C. A dwelling time of 20 min was
bottom face of the supported metal were constrained in z-direction. 2.5. Brazing and thermal cycling test Based on the simulation and analysis, the optimized joints were designed and brazing BCFN tubes to their metal supports were carried out. Considering the complex physicochemical properties of Ag-Cu compared to those for pure Ag, pure Ag was used in simulation. Good wettability of the brazing alloy in the joining components is necessary. Pure silver displayed no wetting on both the BCFN and 310S substrates. 1 mol%Cu was added to Ag to enhance the wetting of the filler upon melting [22]. Ag-1Cu was used as a filler for RAB. The optimized joint
Fig. 4. Material properties of BCFN, Ag filler and 310S in simulation. 1547
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Fig. 5. First principal stresses of ceramic predicted upon cooling from a stress-free state at 965–25 °C for (a) the butt joint without brazing meniscus, (b) the butt joint with brazing meniscus, (c) the inner sleeve joint and (d) the outer sleeve joint.
types of joint structure in three-dimensional space. In the entire assembly, the stresses mainly concentrate in the brazing alloy. The stress for the sleeve joints is generally higher than that for the butt joints. The stress transition of the ceramic membrane in the sleeve joint is uniform (Fig. 5(c) and (d)). The stress mutation of the one in the inner sleeve joint is obvious at the end point of the interface (Fig. 5(c)). The stress of butt joint with brazing meniscus is lower than the one without meniscus. The stress of ceramic in butt joint with brazing meniscus is the lowest one. The maximum stress of tubular ceramic membranes for different joint structures is summarized in Table 1. It is remarkable that the calculated stress is higher than the tested magnitude and still lower than the strength of ceramics. Based on the results of Fig. 5 and Table 1, it can been seen that the stress of ceramic membrane tubes is not only relate with the constraint state of the different joint structures, but also with the geometry, especially the thickness of brazing filler. The shape and thickness of silver can affect its shrinkage and direction. Silver ring would shrink centripetally if it was extremely thin, as Fig. 5(c). It would also contract in axial and radial, as Fig. 5(d). Compared with the size of filler in a butt joint, the silver in the sleeve joint is much big, which can increase the stress of the whole structure. The stress difference of Fig. 5(a) and (b) is
set at 200 °C. The heating and cooling rates were both set at 10 °C/min. Testing was conducted for 11 cycles. Around 1 bar pressure was maintained on both sides of the membrane. When the temperature was held at 875 °C, the flow rates of Ar and air were set to 1000 ml/min. The outlet compositions were measured using a gas chromatograph (VARIAN, CP3800). The oxygen permeation flux was determined from the content of oxygen in the effluent gas. At other temperatures, the flow rates of Ar and air were reduced to 300 ml/min. 3. Results and discussion 3.1. Residual stress distribution The residual stresses can develop in the brazing process during cooling from the braze temperature due to the difference of TECs. The stress calculation of brazing components is represented with the following expression: (1) first principal stress in Figs. 5, (2) axial stress in Figs. 6, (3) radial stress in Figs. 7 and 10. The axial stress and radial stress are used to analyze the residual stress states in different directions. Fig. 5 shows the maximum principal stress distributions of four 1548
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Fig. 6. Axial stress component distribution predicted upon cooling from a stress-free state at 965–25 °C for (a) the butt joint without brazing meniscus, (b) the butt joint with brazing meniscus, (c) the inner sleeve joint and (d) the outer sleeve joint.
certain value, the axial strain of the silver ring becomes obvious and must be considered. According to TECs, the Ag strain is the most. It leads to compression from Ag on BCFN, which generated from the joining interface and stopped at the interface vertex. The silver in outer sleeve joint is thick and causes the greater contraction in axial [23]. This is the reason that the axial compressive stress of ceramic in outer sleeve joint is bigger than the one in the inner sleeve joint. According to Fig. 7(a) and (b), the compression stress concentration exists near the connection interfaces of the membranes for the butt joints. The compression stress of BCFN in the butt joint with meniscus is greater than that in the butt joint without meniscus. Relating to Fig. 6, the axial stress of ceramic is small and the radial stress is large for the butt joint. The residual stress of the ceramic membrane in the butt joint is parallel to the Ag/BCFN interface. It indicates that the membrane
caused by the brazing meniscus for supporting. The similar result was found in the reference [20]. Fig. 6 shows the residual stress distribution of the tubular ceramic membrane in the axial direction. Ceramic stresses for both the butt joint without meniscus and the inner sleeve joint are higher than that of the others. The stress distribution of the butt joint without meniscus is similar to that with meniscus. However, the stress value of the one with the meniscus is less. The axial stress of ceramics for butt joints is zero near the joining interface (Fig. 6(a) and (b)). The down sides (h1 =hA1–hB1, h2 =hA2–hB2) of ceramic tubes in sleeve joints suffer the axial compressive stress (Fig. 6(c) and (d)). The pressure value of the outer joint of ceramic is high. The tension is exerted near the end point along the BCFN/Ag interface (A1, A2) (Fig. 6(d)). When the height of the silver ring in the sleeve joint is increased to a
Fig. 7. Radial stress component distribution predicted upon cooling from a stress-free state at 965–25 °C for (a) the butt joint without brazing meniscus, (b) the butt joint with brazing meniscus, (c) the inner sleeve joint and (d) the outer sleeve joint. 1549
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Table 1 Maximum stress of ceramic in brazing joint.
Maximum principal stress Maximum axial stress Maximum radial stress
Butt joint without brazing meniscus
Butt joint with brazing meniscus
Inner sleeve joint
Outer sleeve joint
130.2 MPa 118.3 MPa 42.4 MPa
52.2 MPa 50.0 MPa 5.3 MPa
113.3 MPa 102.1 MPa 49.6 MPa
134.3 MPa 45.8 MPa 40.7 MPa
and the linear parameters, it is found that the material properties relevant to the residual stress are Young's modulus, Poisson's ratio and TEC, and the heat capacity has no effect on the result. Fig. 10 shows the maximum first principal stress of BCFN membrane tubes with different joint structure during the cooling process. Because the internal stress of the ceramic membrane increases with the temperature difference between the initial and final temperature. The maximum stress value generally increased with the decrease in temperature. And a stress peak was observed at 300 °C, which is caused by the great mechanical properties difference of BCFN and Ag and the strain accumulation. Besides, it can be known that the stress of the ceramic in butt joint without meniscus is the largest in the four types of foundation structures from Fig. 10. Although the ceramic stress of butt joint with meniscus is minimal, this joint structure is not adopted because of its small joining interfacial area, poor air tightness and the difficulty in obtaining the brazing meniscus. The ceramic stress of outer sleeve joint is greater than the one of inner sleeve joint in the range of 250–350 °C. However, when the inner sleeve joint is made of solder with the same thickness as the outer sleeve joint, the ceramic stress of the inner sleeve joint is much greater than that of the outer sleeve joint and exceeds the rupture strength of BCFN. Considering air tightness, and the effect of thickness of solder on the jointing stress in practical application, it is reasonable to select outer sleeve structure as the joint structure. Complex physicochemical and mechanical behaviors were observed during the thermal cycling and special attention should be given to the heating and cooling processes. To join the membrane tubes to their metal supports, two requirements needed: gas-tightness and the minimized joining stress. As shown in Fig. 1 the butt joint and the sleeve joint were considered. For the butt joints, the Ag-Cu braze in the form of pre-shaped foils or as paste was placed between the membrane tube and the metal support (Fig. 1(a)). Although a complete and uniform brazing meniscus is very useful to enhance the joint surface and strength, it is difficult to achieve gastightness for butt joints because of the technical difficulty. Compared with the butt joint, the joining interfacial area of the sleeve joint is much larger (Fig. 1(c) and (d)). The gas-tightness for the sleeve joints can be ensured by the large joining interfacial area and the ideal reaction layers along the interfaces. To minimize the joining stress, the mismatch in TEC of different
tube near the BCFN/Ag interface suffers shear stress in the butt joint. Besides, the joining interface between the membrane tube and the metal support was very limited, especially for a dense membrane tube with only a thickness of 1 mm. To improve the joining interface, a butt joint with an inner or/and outer brazing meniscus along the joining interface can be used (Fig. 1(b)). The brazing meniscus not only reduces the residual stress of the ceramic but also enhance the joint surface. However, our previous investigation showed that it was very challenging to achieve a uniform meniscus along the joining interface and the maintaining the accuracy between the liquid filler alloy and solidification shrinkage was also very difficult. Due to the capillary forces, the liquid filler alloy can distribute in an undefined manner. It is difficult to achieve gas-tightness for the butt joints. The down sides (h1 =hA1–hB1, h2 =hA2–hB2) of ceramic tubes in sleeve joints mainly suffer the tensile radial stress, which originates from the contact interfaces towards the free surface of the tube (Fig. 7(c) and (d)). The tensile stress area of the inner sleeve brazing ceramic is greater than that of the outer sleeve. In sleeve joints, the different deformation leads to the radial stress pulling BCFN through the interface A1B1/A2B2 (Fig. 8(b) and (c)). The overall radial contraction of the silver ring in the radial direction causes the compression on BCFN for the outer sleeve joint. And the pressure reduces the tensile stress value (Fig. 8). The radial stress of BCFN in the outer sleeve joint is less than that of the inner one. Fig. 9 presents the residual stress distribution of the ceramic tube cross-section in the inner sleeve and the outer sleeve. For the contacting parts (A1B1, A2B2) of ceramics in two sleeve assemblies, the free face of the membrane tubes is stress-free. According to the colour bar, the stress difference between the inner face and the outer face for the inner sleeve joint is higher than that of the outer one. Because of the stressfree state for the tube free part, the shear stress is higher along the cross section (A1A2) in the inner sleeve assembly (Fig. 9(b)). Consequently, this would be the origin of cracking. 3.2. Effect of the membrane characteristics Nonlinear physicochemical and mechanical characteristics of mixed-conducting BCFN membrane material were used in the calculation. Comparing the calculation results with the nonlinear parameters
Fig. 8. Residual stresses of Ag/BCFN interfaces in the radial direction for the sleeve joints. 1550
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Fig. 9. The residual stress directions on the cross section of ceramic tubes: (a) h=A1B1/2 inner sleeve, (b) h=A1B1 inner sleeve, (c) h=A2B2/2 outer sleeve, (d) h=A2B2 outer sleeve.
components and the brazing technique must be handled carefully. And the filler alloy with an optimized thickness can be used to compensate for the thermal mismatch. It can be seen from Figs. 6 and 7 that the membrane tube in the butt joint suffers shear stress. Compared with the other joint structures, the shearing residual stress of butt joints would cause the failure of brazing joints because it is over the brazing interfacial strength. Based on the preliminary analysis, there is a smaller residual stress for the outer sleeve joint than that for the inner one during the cooling process. And the stress gradient across the membrane in the inner sleeve joint is so large that will break the ceramic Fig. 9(b)). And the nonlinear stress variation of ceramic (Fig. 10) also shows that the structure of the outer sleeve joint is more stable than the others for its good gas-tightness and minimized joining stress.
3.3. Sealing experiments Fig. 10. Maximum first principal stress of ceramic membranes during cooling process.
According to the above discussion, outer sleeve joints were optimized and designed as shown in Fig. 11(a). The thicknesses of the Ag1551
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consistent with calculation and show that the stress of the outer sleeve joint assembly is relatively small and superior structural integrity and perfect air tightness can be achieved. So the outer sleeve joint assembly is a suitable and achievable structure for joining the tubular membranes with the steel supports. 4. Conclusions The stress of four kinds of ceramic/metal joints was analyzed to achieve a reliable joint structure. Because of the shearing stress, the limited interfacial area and the difficulty to achieve a uniform meniscus, the joining interface failure happens easily for the butt joint structures. Due to the large joining interfacial area, sleeve joints are of the better gas tightness than butt joints. However, the ceramic membrane in the inner sleeve joint suffers the tensile stress and is of the greater stress difference between the inner and outer surfaces. The residual stress of the membrane in the outer sleeve joint is lower than that of the inner sleeve joint and the compressive stress from the silver ring rises during the cooling and heating process. The outer sleeve structure is the most reasonable. The nonlinear properties of the ceramic membrane and the stress accumulation generate a stress peak and lead to brazing failure before the temperature reaches to room temperature. It was also found that the residual stress of the inner sleeve joint is much greater than that of the outer sleeve joint when the geometrical dimensions of their components are identical. Thermal cycling experiments were conducted and verified that the outer sleeve joints are the best candidates for joining the tubular membranes to their metal supports.
Fig. 11. Brazing the BCFN membrane tubes to metal supports using Ag braze in air: (a) the different brazing components, (b) the as-brazed membrane modules with the outer sleeve joint.
Acknowledgments The authors would like to thank Zuosheng Lei for COMSOL software. This work was supported by the National Natural Science Foundation of China – the joint research fund of China Bao-Wu Iron and Steel Group Co., Ltd. (Grant nos. U1860108, U1860203) and the National Natural Science Foundation of China (Grant no. 51576164). References Fig. 12. Oxygen permeation flux and leak rate of the brazed membrane module as a function of the number of thermal cycles.
[1] Xueliang Dong, Wanqin Jin, Mixed conducting ceramic membranes for high efficiency power generation with CO2 capture, Curr. Opin. Chem. Eng. 1 (2) (2012) 163–170. [2] W. Li, X. Zhu, S. Chen, et al., Integration of nine steps into one membrane reactor to produce synthesis gases for ammonia and liquid fuel, Angew. Chem. Int. Ed. 55 (30) (2016) 8566–8570. [3] K. Bernd, B. Tilmann, Achievements in reactive air brazing, Adv. Eng. Mater. 16 (12) (2014) 1413–1414. [4] K. Raju, Muksin, S. Kim, et al., Joining of metal-ceramic using reactive air brazing for oxygen transport membrane applications, Mater. Des. 109 (2016) 233–241. [5] K. Raju, Muksin, D.H. Yoon, Reactive air brazing of GDC-LSCF ceramics using Ag-10 wt% CuO paste for oxygen transport membrane applications, Ceram. Int. 42 (14) (2016) 16392–16395. [6] H. Chen, L. Li, R. Kemps, et al., Reactive air brazing for sealing mixed ionic electronic conducting hollow fibre membranes, Acta Mater. 88 (2015) 74–82. [7] Y. Zhang, T. Liu, J. Zhang, C. Wu, X. Lu, W. Ding, Induction brazing BaCo0.7Fe0.2Nb0.1O3-δ membrane tubes to steel supports with Ag-based filler in air, J. Membr. Sci. 533 (2017) 19–27. [8] A. Kaletsch, E.M. Pfaff, C. Broeckmann, Effect of aging on microstructure and mechanical strength of reactive air brazed BSCF/AISI 314-joints, Adv. Eng. Mater. 16 (12) (2015) 1430–1436. [9] Y.K. Jin, M. Engelhard, J.P. Choi, et al., Effects of atmospheres on bonding characteristics of silver and alumina, Int. J. Hydrog. Energy 33 (14) (2008) 4001–4011. [10] J. Brandenberg, B. Kuhn, M. Pausch, Dual-gas stability of RAB-joints in SOFC applications**, Adv. Eng. Mater. 16 (12) (2014) 1415–1422. [11] K. Bobzin, M. Öte, S. Wiesner, A. Kaletsch, C. Broeckmanm, Characterization of reactive air brazed ceramic/metal joints with unadapted thermal expansion behavior, Adv. Eng. Mater. 16 (12) (2015) 1490–1497. [12] R. Kiebach, et al., Joining of ceramic Ba0.5Sr0.5Co0.8Fe0.2O3 membranes for oxygen production to high temperature alloys, J. Membr. Sci. 506 (2016) 11–21. [13] Y. Wang, Z.Z. Duan, G. Chen, et al., Effects of brazing temperature on microstructure and properties of interface between cBN and Co-based active filler metals, Vacuum 145 (2017) 30–38. [14] K.S. Weil, B.J. Koeppel, Comparative finite element analysis of the stress–strain states in three different bonded solid oxide fuel cell seal designs, J. Power Sources
based alloy and the sleeve wall were designed at 2.5 mm. [24] Filler alloy in the form of cylindrical ring was placed in the gap between the membrane tube and the 310S sleeve. The membrane tubes with a length of 100 mm were fabricated by slip casting after the aqueous slurry of BCFN powder was prepared. After assembling, the induction brazing was carried out in air. A number of outer sleeve joints were achieved successfully (Fig. 11(b)) and verified their feasibility as the joint structures. Testing of brazed membrane tube modules under the total pressure gradient of about 1–2 bar at room temperature confirmed gastightness. To further confirm the stability of the brazed joints, the leak rate and oxygen permeation flux of the sealing sample as a function of the number of cycles was characterized. As shown in Fig. 12, no measurable degradation in either hermeticity or the oxygen permeation flux was found during the thermal cycling. The result indicates the mismatch stress along the BCFN membrane/Ag-based braze/310S metal support interfaces during thermal cycles can be mitigated via strain relaxation mechanisms associated with yielding of the joining components and did not reduce the effective stress that the membrane and seal could withstand. At the same time, the BCFN membrane itself can withstand the chemically induced stress under the oxygen chemical gradient of Ar/air. There is no crack or fracture occurred on the BCFN membrane tubes during the thermal cycling testing and the BCFN membrane exhibited good thermal shock resistance. The experimental results are 1552
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[20] S. Dabbarh, E. Pfaff, A. Ziombra, A. Bezold, Brazing of MIEC ceramics to high temperature metals, in: D. Singh, E. Pfaff, A. Ziombra, A. Bezold (Eds.), Design, Development, and Applications of Engineering Ceramics and Composites, John Wiley & Sons, Inc, 2010, pp. 213–223. [21] Q. Li, X. Zhu, Y. He, et al., Effects of sintering temperature on properties of dualphase oxygen permeable membranes, J. Membr. Sci. 367 (1) (2011) 134–140. [22] K. Bobzin, T. Schlaefer, L. Zhao, et al., Brazing of ceramic-to-ceramic and ceramicto-metal joints in air, Front. Mech. Eng. Chin. 5 (2) (2010) 125–129. [23] K.M. Erskine, A.M. Meier, V.V. Joshi, S.M. Pilgrim, The effect of braze interlayer thickness on the mechanical strength of alumina brazed with Ag–CuO braze alloys, Adv. Eng. Mater. 16 (12) (2015) 1442–1447. [24] B. Chen, C. Cheng, L. Xiong, L. Wang, Transformation toughening of Al2O3/ZrO2 laminated ceramics with residual compressive stress, J. Univ. Sci. Technol. Beijing, Min., Metall., Mater. 14 (5) (2007) 449–453.
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Stress analysis of the brazing joints of tubular ceramic oxygen-permeable membranes and metal supports ⁎
Jialu Zhang, Jian Zhang, Lanqian Li, Chao Zhang, Yuwen Zhang , Xionggang Lu
T
⁎
State Key Laboratory of Advanced Special Steel & Shanghai Key Laboratory of Advanced Ferrometallurgy & School of Materials Science and Engineering, Shanghai University, Shanghai 200444, China
A R T I C LE I N FO
A B S T R A C T
Keywords: Stress analysis Brazing joints Tubular ceramic oxygen-permeable membrane Nonlinear material properties
This study investigates the effect of joint structure and temperature dependent mechanical properties on the residual stress states in ceramic/metal brazing joints. A closed-one-end BaCo0.7Fe0.2Nb0.1O3 (BCFN) membrane tube is joined to 310S by using silver filler. Finite element simulation of different BCFN/310S joint assemblies is conducted to evaluate the stresses and deformations of the components at an isothermal state. The results indicate that the outer sleeve structure with low residual stress and good air tightness is more suitable for joining ceramics to metal supports than other joint structures in four joint assemblies. And thermal cycling experiments also verify the functional reliability of the outer sleeve joint assembly. Besides, nonlinear material properties of BCFN can not only cause the sudden change in ceramic stress at 300 °C, but also be the important basis for the stress contrast of different joint structures.
1. Introduction Mixed-conducting ceramic oxygen-permeable membranes show the attractive applications in pure oxygen production, membrane reactors for catalytic reforming, and oxy-fuel combustion [1,2]. To develop a commercial process, the large dimension membrane modules with gastight sealing to their metal supports must be constructed [3]. The critical issue of fabricating high-performance joints is the development of the proper joint structure and technique for joining membranes to metal supports. Metal brazing has been paid more and more attention for joining the oxygen-permeable membranes to metal supports. Without the use of fluxes and other special atmosphere, reactive air brazing (RAB) has been investigated extensively in the field of sealing the oxygenpermeable membranes. Due to the ductility and oxidation resistance, Ag-based alloys have been used as metallic sealing fillers. The oxygenpermeable membranes, such as (La0.6Sr0.4)(Co0.2Fe0.8)O3-δ (LSCF) [4,5], BaCo0.4Fe0.4Zr0.2O3-δ (BCFZ) [6], BaCo0.7Fe0.2Nb0.1O3-δ (BCFN) [7], Ba0.5Sr0.5Co0.8Fe0.2O3-δ (BSCF) [8], were used in these brazing experiments. Many studies on RAB have been conducted on the wettability, interfacial reactions and joining strength between the ceramic membranes and the brazes [4–11]. In RAB, the interfacial reactive can improve the wettability, adhesion and bonding strength. RAB was also used to join the tubular membranes to metal supports in high-
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temperature resistance furnaces [12] or by an induction heating method [13]. Due to the mismatch of the thermal expansion of the ceramic membranes and metal supports, the joining stresses are very high and the membranes are easily destructed. In order to minimize joining stresses, it is necessary to optimize the joint structures. A combined model of the process parameters with the simulation of the material behavior would be very helpful for the structure optimization and the reliability assessment of the membrane modules. Compared with the extensive study on the stress simulation for SOFC [14–19], there are a few of reports on the stress analysis for joining tubular mixed-conducting oxygen-permeable ceramic to its metal support. Ragnar Kiebach investigated the residual stress of BSCF/Haynes 214 butt and lap joint assemblies [12]. The stress of the lap joint geometry is larger. Dabbarh et al. analyzed the mechanical reliability of BSCF/Alloy 310Si joints by the wetting experiments and stress simulation. They found that the stress could be reduced by changing from the butt joint to the sleeve joint [20]. Kirsten Bobzin calculated the residual stress of BSCF/AISI 314 joint and BSCF/Crofer 22H joint by using Abaqus and established a visco-plastic model to analyze the effect of silver creep on the residual stress [11]. The similar TECs of the components are useful to reduce the residual stress and the solder creep can also release the stress. Obviously, there are some discrepancies on the simulation and optimization of the joint structures from the different reports. The variation in the sizes of the solder and metal
Corresponding authors E-mail addresses: [email protected] (Y. Zhang), [email protected] (X. Lu).
https://doi.org/10.1016/j.ceramint.2018.10.028 Received 28 September 2018; Received in revised form 2 October 2018; Accepted 3 October 2018 Available online 10 October 2018 0272-8842/ © 2018 Elsevier Ltd and Techna Group S.r.l. All rights reserved.
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supports has not been analyzed and optimized in these studies. Especially, the ceramic properties used in these reports were usually considered as the constant or linear values as a function of temperature. In fact, with the temperature changing, the physicochemical and mechanical properties of mixed-conducting ceramic oxygen-permeable membranes are nonlinear and complicated [2,13,21]. The anomalies in the properties may cause damage of the membranes, and special consideration should be given to the design and the operation of the membrane reactors. The aim of this study was to seek an optimized joint structure for joining the tubular membranes to their metal supports by conducting a systematic stress simulation. The brazing joints of tubular ceramic oxygen-permeable membranes BCFN and 310S stainless steel supports (310S) with different geometric structures were analyzed and compared. COMSOL software was utilized to model and analyze the stresses and strains that arise in the mismatch of the joining components. Nonlinear physicochemical and mechanical characteristics of mixedconducting BCFN membrane material were used and paid more attention in the stress modeling and analysis.
Fig. 2. Simplified diagram of BCFN ceramic membrane tube.
(965 °C). Then, the residual stresses would appear in the assemblies during cooling down from high temperature (965 °C) to room temperature (25 °C) because of different TECs. The residual stresses are calculated by a physical simulation software COMSOL without considering the influence of creep and the elastic-plastic deformation being governed by the braze model. Simultaneously, the total strain of the ceramic membrane can be decomposed into elastic strain and thermal strain. And the strain in ceramic can be described in the Eqs. (1) and (2), where D is an elastic matrix, which is related with Young's modulus and Poisson's ratio. The thermal strain mainly derived from the differences of TEC and temperature.
2. Modeling and experiments 2.1. Geometric model To join the membrane tubes to their metal supports suitably, four geometric structures of the butt and sleeve joints are designed as shown in Fig. 1: (1) butt joint without brazing meniscus (Figs. 1(a)), (2) butt joint with brazing meniscus (Figs. 1(b)), (3) inner sleeve joint (Figs. 1(c)) and (4) outer sleeve joint (Fig. 1(d)). During the simulation process, a closed-one-end BCFN membrane tube with an outer diameter of 10 mm and a thickness of 1 mm are used. Similarly, silver rings with different thickness of 0.6 mm, 0.35 mm and 2 mm are prepared for the butt joint, inner sleeve joints and outer sleeve joints respectively. The simplified diagram of ceramic tube is shown in the Fig. 2, whose bottom part (A1B1, A2B2) represents the brazing part in sleeve joint. Only a half axial section was considered due to the rotationally symmetry of the sealing part along the vertical Z axis. The finite element meshes were generated automatically by the simulation software, as shown in Fig. 3. Notably, the maximum mesh sizes of four kinds of assemblies are same. The paths A1B1 (Fig. 3(c)) and A2B2 (Fig. 3(c)) represent the Ag/BCFN interfaces of the inner sleeve joint and the outer sleeve joint respectively.
σ = Dεel + σ0
(1)
εel = ε − ε0 − εth
(2)
2.3. Material properties The parameters of the properties of silver and 310S are provided by COMSOL Material Library. The parameters of the BCFN properties are characterized by our group. Fig. 4 demonstrates the properties of the different materials used in the simulation. And it is clear that the variation of BCFN material properties is nonlinear with the temperature varying, especially its TEC and heat capacity. 2.4. Boundary and constrained conditions The two-dimensional axisymmetric boundary conditions were applied in the axial sections. The Ag/BCFN interface and Ag/310S interface were defined as the identity boundary pairs. All the nodes on the
2.2. Simulation procedure At first, all assemblies are in a stress-free state at initial temperature
Fig. 1. Cross-sectional schematics of (a) the butt joint assembly without brazing meniscus, (b) the butt joint assembly with brazing meniscus, (c) the inner sleeve joint assembly and (d) the outer sleeve joint assembly. 1546
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Fig. 3. Finite element meshing of (a) the butt joint assembly without brazing meniscus, (b) the butt joint assembly with brazing meniscus, (c) the inner sleeve joint assembly and (d) the outer sleeve joint assembly.
was heated from room temperature to brazing temperature (965 °C) within 12 min and then was directly cooled to 900 °C within 1 min. With the subsequent cooling following, a cooling rate of 20 °C/min was taken during 900–600 °C and a cooling rate of 10 °C/min for 600–300 °C step. Below 300 °C, the assembly was cooled down to room temperature. To examine the reliability of the optimized joints, thermal cycle testing was conducted. The integrity of the joints was tested using a setup described in Ref. [8]. The leakage of N2 in the air into the sweeping gas (Ar) was continuously monitored by an online mass spectrometer (Hidden HPR20) to confirm that the membrane tube and the seal remained intact. The brazed module was heated to 875 °C, holding at 875 °C for 30 min, and cooling to 200 °C. A dwelling time of 20 min was
bottom face of the supported metal were constrained in z-direction. 2.5. Brazing and thermal cycling test Based on the simulation and analysis, the optimized joints were designed and brazing BCFN tubes to their metal supports were carried out. Considering the complex physicochemical properties of Ag-Cu compared to those for pure Ag, pure Ag was used in simulation. Good wettability of the brazing alloy in the joining components is necessary. Pure silver displayed no wetting on both the BCFN and 310S substrates. 1 mol%Cu was added to Ag to enhance the wetting of the filler upon melting [22]. Ag-1Cu was used as a filler for RAB. The optimized joint
Fig. 4. Material properties of BCFN, Ag filler and 310S in simulation. 1547
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Fig. 5. First principal stresses of ceramic predicted upon cooling from a stress-free state at 965–25 °C for (a) the butt joint without brazing meniscus, (b) the butt joint with brazing meniscus, (c) the inner sleeve joint and (d) the outer sleeve joint.
types of joint structure in three-dimensional space. In the entire assembly, the stresses mainly concentrate in the brazing alloy. The stress for the sleeve joints is generally higher than that for the butt joints. The stress transition of the ceramic membrane in the sleeve joint is uniform (Fig. 5(c) and (d)). The stress mutation of the one in the inner sleeve joint is obvious at the end point of the interface (Fig. 5(c)). The stress of butt joint with brazing meniscus is lower than the one without meniscus. The stress of ceramic in butt joint with brazing meniscus is the lowest one. The maximum stress of tubular ceramic membranes for different joint structures is summarized in Table 1. It is remarkable that the calculated stress is higher than the tested magnitude and still lower than the strength of ceramics. Based on the results of Fig. 5 and Table 1, it can been seen that the stress of ceramic membrane tubes is not only relate with the constraint state of the different joint structures, but also with the geometry, especially the thickness of brazing filler. The shape and thickness of silver can affect its shrinkage and direction. Silver ring would shrink centripetally if it was extremely thin, as Fig. 5(c). It would also contract in axial and radial, as Fig. 5(d). Compared with the size of filler in a butt joint, the silver in the sleeve joint is much big, which can increase the stress of the whole structure. The stress difference of Fig. 5(a) and (b) is
set at 200 °C. The heating and cooling rates were both set at 10 °C/min. Testing was conducted for 11 cycles. Around 1 bar pressure was maintained on both sides of the membrane. When the temperature was held at 875 °C, the flow rates of Ar and air were set to 1000 ml/min. The outlet compositions were measured using a gas chromatograph (VARIAN, CP3800). The oxygen permeation flux was determined from the content of oxygen in the effluent gas. At other temperatures, the flow rates of Ar and air were reduced to 300 ml/min. 3. Results and discussion 3.1. Residual stress distribution The residual stresses can develop in the brazing process during cooling from the braze temperature due to the difference of TECs. The stress calculation of brazing components is represented with the following expression: (1) first principal stress in Figs. 5, (2) axial stress in Figs. 6, (3) radial stress in Figs. 7 and 10. The axial stress and radial stress are used to analyze the residual stress states in different directions. Fig. 5 shows the maximum principal stress distributions of four 1548
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Fig. 6. Axial stress component distribution predicted upon cooling from a stress-free state at 965–25 °C for (a) the butt joint without brazing meniscus, (b) the butt joint with brazing meniscus, (c) the inner sleeve joint and (d) the outer sleeve joint.
certain value, the axial strain of the silver ring becomes obvious and must be considered. According to TECs, the Ag strain is the most. It leads to compression from Ag on BCFN, which generated from the joining interface and stopped at the interface vertex. The silver in outer sleeve joint is thick and causes the greater contraction in axial [23]. This is the reason that the axial compressive stress of ceramic in outer sleeve joint is bigger than the one in the inner sleeve joint. According to Fig. 7(a) and (b), the compression stress concentration exists near the connection interfaces of the membranes for the butt joints. The compression stress of BCFN in the butt joint with meniscus is greater than that in the butt joint without meniscus. Relating to Fig. 6, the axial stress of ceramic is small and the radial stress is large for the butt joint. The residual stress of the ceramic membrane in the butt joint is parallel to the Ag/BCFN interface. It indicates that the membrane
caused by the brazing meniscus for supporting. The similar result was found in the reference [20]. Fig. 6 shows the residual stress distribution of the tubular ceramic membrane in the axial direction. Ceramic stresses for both the butt joint without meniscus and the inner sleeve joint are higher than that of the others. The stress distribution of the butt joint without meniscus is similar to that with meniscus. However, the stress value of the one with the meniscus is less. The axial stress of ceramics for butt joints is zero near the joining interface (Fig. 6(a) and (b)). The down sides (h1 =hA1–hB1, h2 =hA2–hB2) of ceramic tubes in sleeve joints suffer the axial compressive stress (Fig. 6(c) and (d)). The pressure value of the outer joint of ceramic is high. The tension is exerted near the end point along the BCFN/Ag interface (A1, A2) (Fig. 6(d)). When the height of the silver ring in the sleeve joint is increased to a
Fig. 7. Radial stress component distribution predicted upon cooling from a stress-free state at 965–25 °C for (a) the butt joint without brazing meniscus, (b) the butt joint with brazing meniscus, (c) the inner sleeve joint and (d) the outer sleeve joint. 1549
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Table 1 Maximum stress of ceramic in brazing joint.
Maximum principal stress Maximum axial stress Maximum radial stress
Butt joint without brazing meniscus
Butt joint with brazing meniscus
Inner sleeve joint
Outer sleeve joint
130.2 MPa 118.3 MPa 42.4 MPa
52.2 MPa 50.0 MPa 5.3 MPa
113.3 MPa 102.1 MPa 49.6 MPa
134.3 MPa 45.8 MPa 40.7 MPa
and the linear parameters, it is found that the material properties relevant to the residual stress are Young's modulus, Poisson's ratio and TEC, and the heat capacity has no effect on the result. Fig. 10 shows the maximum first principal stress of BCFN membrane tubes with different joint structure during the cooling process. Because the internal stress of the ceramic membrane increases with the temperature difference between the initial and final temperature. The maximum stress value generally increased with the decrease in temperature. And a stress peak was observed at 300 °C, which is caused by the great mechanical properties difference of BCFN and Ag and the strain accumulation. Besides, it can be known that the stress of the ceramic in butt joint without meniscus is the largest in the four types of foundation structures from Fig. 10. Although the ceramic stress of butt joint with meniscus is minimal, this joint structure is not adopted because of its small joining interfacial area, poor air tightness and the difficulty in obtaining the brazing meniscus. The ceramic stress of outer sleeve joint is greater than the one of inner sleeve joint in the range of 250–350 °C. However, when the inner sleeve joint is made of solder with the same thickness as the outer sleeve joint, the ceramic stress of the inner sleeve joint is much greater than that of the outer sleeve joint and exceeds the rupture strength of BCFN. Considering air tightness, and the effect of thickness of solder on the jointing stress in practical application, it is reasonable to select outer sleeve structure as the joint structure. Complex physicochemical and mechanical behaviors were observed during the thermal cycling and special attention should be given to the heating and cooling processes. To join the membrane tubes to their metal supports, two requirements needed: gas-tightness and the minimized joining stress. As shown in Fig. 1 the butt joint and the sleeve joint were considered. For the butt joints, the Ag-Cu braze in the form of pre-shaped foils or as paste was placed between the membrane tube and the metal support (Fig. 1(a)). Although a complete and uniform brazing meniscus is very useful to enhance the joint surface and strength, it is difficult to achieve gastightness for butt joints because of the technical difficulty. Compared with the butt joint, the joining interfacial area of the sleeve joint is much larger (Fig. 1(c) and (d)). The gas-tightness for the sleeve joints can be ensured by the large joining interfacial area and the ideal reaction layers along the interfaces. To minimize the joining stress, the mismatch in TEC of different
tube near the BCFN/Ag interface suffers shear stress in the butt joint. Besides, the joining interface between the membrane tube and the metal support was very limited, especially for a dense membrane tube with only a thickness of 1 mm. To improve the joining interface, a butt joint with an inner or/and outer brazing meniscus along the joining interface can be used (Fig. 1(b)). The brazing meniscus not only reduces the residual stress of the ceramic but also enhance the joint surface. However, our previous investigation showed that it was very challenging to achieve a uniform meniscus along the joining interface and the maintaining the accuracy between the liquid filler alloy and solidification shrinkage was also very difficult. Due to the capillary forces, the liquid filler alloy can distribute in an undefined manner. It is difficult to achieve gas-tightness for the butt joints. The down sides (h1 =hA1–hB1, h2 =hA2–hB2) of ceramic tubes in sleeve joints mainly suffer the tensile radial stress, which originates from the contact interfaces towards the free surface of the tube (Fig. 7(c) and (d)). The tensile stress area of the inner sleeve brazing ceramic is greater than that of the outer sleeve. In sleeve joints, the different deformation leads to the radial stress pulling BCFN through the interface A1B1/A2B2 (Fig. 8(b) and (c)). The overall radial contraction of the silver ring in the radial direction causes the compression on BCFN for the outer sleeve joint. And the pressure reduces the tensile stress value (Fig. 8). The radial stress of BCFN in the outer sleeve joint is less than that of the inner one. Fig. 9 presents the residual stress distribution of the ceramic tube cross-section in the inner sleeve and the outer sleeve. For the contacting parts (A1B1, A2B2) of ceramics in two sleeve assemblies, the free face of the membrane tubes is stress-free. According to the colour bar, the stress difference between the inner face and the outer face for the inner sleeve joint is higher than that of the outer one. Because of the stressfree state for the tube free part, the shear stress is higher along the cross section (A1A2) in the inner sleeve assembly (Fig. 9(b)). Consequently, this would be the origin of cracking. 3.2. Effect of the membrane characteristics Nonlinear physicochemical and mechanical characteristics of mixed-conducting BCFN membrane material were used in the calculation. Comparing the calculation results with the nonlinear parameters
Fig. 8. Residual stresses of Ag/BCFN interfaces in the radial direction for the sleeve joints. 1550
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Fig. 9. The residual stress directions on the cross section of ceramic tubes: (a) h=A1B1/2 inner sleeve, (b) h=A1B1 inner sleeve, (c) h=A2B2/2 outer sleeve, (d) h=A2B2 outer sleeve.
components and the brazing technique must be handled carefully. And the filler alloy with an optimized thickness can be used to compensate for the thermal mismatch. It can be seen from Figs. 6 and 7 that the membrane tube in the butt joint suffers shear stress. Compared with the other joint structures, the shearing residual stress of butt joints would cause the failure of brazing joints because it is over the brazing interfacial strength. Based on the preliminary analysis, there is a smaller residual stress for the outer sleeve joint than that for the inner one during the cooling process. And the stress gradient across the membrane in the inner sleeve joint is so large that will break the ceramic Fig. 9(b)). And the nonlinear stress variation of ceramic (Fig. 10) also shows that the structure of the outer sleeve joint is more stable than the others for its good gas-tightness and minimized joining stress.
3.3. Sealing experiments Fig. 10. Maximum first principal stress of ceramic membranes during cooling process.
According to the above discussion, outer sleeve joints were optimized and designed as shown in Fig. 11(a). The thicknesses of the Ag1551
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consistent with calculation and show that the stress of the outer sleeve joint assembly is relatively small and superior structural integrity and perfect air tightness can be achieved. So the outer sleeve joint assembly is a suitable and achievable structure for joining the tubular membranes with the steel supports. 4. Conclusions The stress of four kinds of ceramic/metal joints was analyzed to achieve a reliable joint structure. Because of the shearing stress, the limited interfacial area and the difficulty to achieve a uniform meniscus, the joining interface failure happens easily for the butt joint structures. Due to the large joining interfacial area, sleeve joints are of the better gas tightness than butt joints. However, the ceramic membrane in the inner sleeve joint suffers the tensile stress and is of the greater stress difference between the inner and outer surfaces. The residual stress of the membrane in the outer sleeve joint is lower than that of the inner sleeve joint and the compressive stress from the silver ring rises during the cooling and heating process. The outer sleeve structure is the most reasonable. The nonlinear properties of the ceramic membrane and the stress accumulation generate a stress peak and lead to brazing failure before the temperature reaches to room temperature. It was also found that the residual stress of the inner sleeve joint is much greater than that of the outer sleeve joint when the geometrical dimensions of their components are identical. Thermal cycling experiments were conducted and verified that the outer sleeve joints are the best candidates for joining the tubular membranes to their metal supports.
Fig. 11. Brazing the BCFN membrane tubes to metal supports using Ag braze in air: (a) the different brazing components, (b) the as-brazed membrane modules with the outer sleeve joint.
Acknowledgments The authors would like to thank Zuosheng Lei for COMSOL software. This work was supported by the National Natural Science Foundation of China – the joint research fund of China Bao-Wu Iron and Steel Group Co., Ltd. (Grant nos. U1860108, U1860203) and the National Natural Science Foundation of China (Grant no. 51576164). References Fig. 12. Oxygen permeation flux and leak rate of the brazed membrane module as a function of the number of thermal cycles.
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based alloy and the sleeve wall were designed at 2.5 mm. [24] Filler alloy in the form of cylindrical ring was placed in the gap between the membrane tube and the 310S sleeve. The membrane tubes with a length of 100 mm were fabricated by slip casting after the aqueous slurry of BCFN powder was prepared. After assembling, the induction brazing was carried out in air. A number of outer sleeve joints were achieved successfully (Fig. 11(b)) and verified their feasibility as the joint structures. Testing of brazed membrane tube modules under the total pressure gradient of about 1–2 bar at room temperature confirmed gastightness. To further confirm the stability of the brazed joints, the leak rate and oxygen permeation flux of the sealing sample as a function of the number of cycles was characterized. As shown in Fig. 12, no measurable degradation in either hermeticity or the oxygen permeation flux was found during the thermal cycling. The result indicates the mismatch stress along the BCFN membrane/Ag-based braze/310S metal support interfaces during thermal cycles can be mitigated via strain relaxation mechanisms associated with yielding of the joining components and did not reduce the effective stress that the membrane and seal could withstand. At the same time, the BCFN membrane itself can withstand the chemically induced stress under the oxygen chemical gradient of Ar/air. There is no crack or fracture occurred on the BCFN membrane tubes during the thermal cycling testing and the BCFN membrane exhibited good thermal shock resistance. The experimental results are 1552
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