- Email: [email protected]
wet air
Journal of Alloys and Compounds 574 (2013) 437–442
Contents lists available at SciVerse ScienceDirect
Journal of Alloys and Compounds journal homepage: www.elsevier.com/locate/jalcom
Oxidation mechanism of Fe–16Cr alloy as SOFC interconnect in dry/wet air Zhi-Yuan Chen a, Li-Jun Wang a,⇑, Fu-Shen Li b, Kuo-Chih Chou a a b
School of Metallurgical and Ecological Engineering, University of Science and Technology Beijing, Beijing 100083, PR China School of Material Science Engineering, University of Science and Technology Beijing, Beijing 100083, PR China
a r t i c l e
i n f o
Article history: Received 13 March 2013 Received in revised form 4 May 2013 Accepted 12 May 2013 Available online 20 May 2013 Keywords: Break-away oxidation Real physical picture model Oxidization SOFC interconnector
a b s t r a c t Experimental study on the oxidation corrosions of Fe–16Cr alloy was carried out at 800–1100 °C under dry/wet air conditions. Faster oxidation rate was observed at higher temperature and water vapor content. The degradation time td between two stages in oxidation process showed an exponential relationship with elevating corrosion temperature in dry air, and a linear relationship with the water content in the case of water vapor introduced to the system. The mechanism of oxidation corrosions of Fe–16Cr alloy was suggested by the Real Physical Picture (RPP) model. It was found that the break-away oxidation in stage II was controlled by diffusion at initial both in dry and wet air, then became linear with the exposure time, which implied that the oxidation rate was then controlled by chemical reaction of the interface between the metal and the oxidized scale. Moreover, the effect of water in the oxidation process is not only to supply more oxygen into system, but also to modify the structures of oxide scale due to the existence of hydrogen atom, which results in the accelerated corrosions. Ó 2013 Elsevier B.V. All rights reserved.
1. Introduction Alloy interconnect attracts a great deal of attention for commercial use of SOFC due to its several advantages compared to ceramics, such as gas tightness, and machinability. Alloy interconnect can be used below 800 °C where Fe–Cr alloys exhibit good oxidation resistance. Fe–16Cr alloy is one of promising alloy interconnects for IT-SOFC. A number of studies [1–4] have been carried out on the oxidation behavior of the Fe–16Cr alloy under various atmospheres. As it is known, the oxidation rate of the metallic interconnect at relative high temperature is of great importance to the stable operation of SOFC. The oxidation process of Fe–Cr alloy usually contains two stages—one is a smaller mass-gain rate stage, the other is with larger rate followed [1,3,5]. The mass-gain rate of the second stage denotes the broken rate of the dense scale. And the degradation time td of the first stage to the second can be regarded as the beginning of protective layer breaking. So it might be a key parameter to define the oxidation resistance other than rate constant. Although oxidation kinetics of the Fe–Cr alloys has been investigated in many researches [1,6–8], information on the degradation time td of the first stage to the second one was still limited [5,9].
⇑ Corresponding author. Tel.: +86 10 6233 3622; fax: +86 6233 2570. E-mail address: [email protected] (L.-J. Wang). 0925-8388/$ - see front matter Ó 2013 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.jallcom.2013.05.072
Researchers [8–11] pointed out the mass-gain of Fe–Cr alloys was parabolic to the exposure time, which indicated that the diffusion of the ions through the oxidate layer was the rate-determining step according to Wagner’s theory. Meanwhile, in view of the vaporization of the chromia in the alloy oxidation process, a more complex kinetic law was also addressed [1,12]. The cathode of SOFC is operating at atmosphere more or less containing water vapor. When water was introduced into the air, the growth rate of alloy scale was observed to be increased with the water vapor content at high temperature [5,13]. It suggested the oxidation rate of Fe–16Cr alloy is accelerated in the atmosphere containing water vapor due to the break-down of protective layer. Water vapor can affect the oxidation behavior in several approaches, such as influencing the transport processes of ions in the alloy and modifying the void formation in the oxide scale [14,15]. The aim of the current work is to study the kinetics of the second breakaway stage of Fe–16Cr alloy oxidation at dry/wet air, and obtain the relationship of the degradation time td with two factors: temperature and water vapor content.
2. Experimental The chemical composition of Fe–16Cr alloy used in the present study is shown in Table 1. The main elements were also analyzed by National Analysis Center of Iron and Steel of China. The samples were cut to the rectangular sheet, 21 42 1.3 mm, and were ground with SiC abrasive paper up to #2000. After polishing, the surfaces of the samples were cleaned by alcohol.
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Z.-Y. Chen et al. / Journal of Alloys and Compounds 574 (2013) 437–442
Table 1 Chemical composition of Fe–16Cr alloy. Elements
C
Si
Cr
Ni
Mn
S
P
Fe
Mass% Checked mass%
0.05 unchecked
0.40 0.31
16.05 15.98
0.05 0.089
0.25 0.30
0.005 unchecked
0.012 unchecked
Remnant Remnant
Fig. 3. SEM photo of the cross section of Fe–16Cr alloy scale oxidized in air for 1 h at 1000 °C.
Table 2 EDS results corresponding to the points of the cross section Fe–16Cr alloy scale oxidized in air for 1 h at 1000 °C at Fig. 3. Elements (wt%)
1
2
3
4
Fe Cr O
84.80 15.20 –
29.60 38.43 31.97
66.32 6.23 27.45
65.85 0.79 33.36
Fig. 1. The device of the alloy sample hanging on the balance.
Fig. 2. The mass changing curves of Fe–16Cr alloy in dry air at different temperatures.
Sample arrangement for oxidation under the dry/wet air atmosphere is illustrated in Fig. 1. An alumina stick (diameter is 2 mm) through the hole of the sample sheet was connected to the OHAUS AR153CN balance (accuracy 1 mg) using an alumina pipe. The mass gain of the sample sheet was recorded during oxidation. The inner area of the drilled hole was ignored in the following calculation due to the tight fit between the hole and the alumina stick. Oxidation experiments were
Fig. 4. Equilibrium diagram of 100 g Fe–Cr alloy continuously exposure to air at 1000 °C.
carried out in the temperature region of 800–1100 °C under dry air and at 1000 °C under wet air, respectively. The gas flow was controlled to 100 ml/min at 25 °C (accuracy ± 1.8 ml/min), by Alicat Gas Flow Controller. The water was pumped into a preheating chamber in 1–7 ll/min using a MasterFlex peristaltic
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Z.-Y. Chen et al. / Journal of Alloys and Compounds 574 (2013) 437–442
Fig. 5. The micrographs of oxide scales at 1000 °C and 1100 °C for 50 h ((a) scale surface at 1000 °C and (b) scale surface at 1100 °C).
rate-determining step of the reaction. A parameter of rate constant ‘‘k’’ was used, in which all effects of factors involved. This kind of treatment is too complicated to offer an explicit analytic expression and an intuitionistic quantitative discussion. More recently, Chou and Hou [16–22] proposed a Real Physical Picture model (RPP) to solve the problem mentioned above, in which each parameters had clear physical meaning and is quite convenient for the theoretical discussions. A more substantive parameter called ‘‘characteristic oxidation time’’ was introduced into the model to characterize the oxidation reaction kinetics. According to RPP model, the reaction fraction n can be expressed as following [23]:
n¼
Dm pffiffiffiffiffiffiffiffiffiffiffi ¼ t=t U1 m
t U1 ¼ Fig. 6. Degradation time td between the two stages oxidation as a function of temperature. pump (accuracy ± 0.1%), mixing with dry air to obtain wet air, which contained 1.2 vol%, 2.5 vol%, and 8.0 vol% water vapor (1, 2, and 7 ll water (l) per 100 ml air (g)), respectively. Microstructure observation after oxidation was carried out by a scanning electron microscope (SEM, CARL ZEISS EVO MA 10/LS 10 JS) with energy dispersive spectrum (EDS, Thermo NORAN System).
3. Kinetic models The kinetic mechanism of Fe–Cr alloy oxidation was usually suggested that the diffusion of ion in the oxide scale was the
tm L20
eq
DðC C Þ
¼
ð1Þ
tm L20 D0 ðC C eq Þ
exp
DE1 RT
ð2Þ
where tU is named as the ‘‘characteristic oxidation time’’. When t = tU, the reaction fraction n is equal to 1. D is the diffusion coefficient of the certain component, here D0 is a constant. C and Ceq are the diffusion atom activities at alloy surface and alloy/scale interface with equilibrium condition, respectively. The rest of parameters are related to the external physical characteristic, where tm is a coefficient related to the density of the substance and reaction, L0 is the original thickness of the alloy. If the rate-determining step of the oxidation is a chemical reaction, the relationship of the reaction fraction to exposure time is as following [23]:
n ¼ t=t U2 t U2 ¼
ð3Þ
tm L20 KðC 0 C eq Þ
¼
tm L20 K 0 ðC 0 C eq Þ
exp
DE2 RT
ð4Þ
where K is the equilibrium constant, K0 is a constant, C0 is the reactant activity in the interface oxidation reaction. Eqs. (1) and (3) are used to analyze the kinetics of the alloy oxidation in the following chapters. 4. Results and discussion 4.1. Oxidation corrosion in dry air
Fig. 7. The kinetic equations describing the alloy oxidation at different temperatures (fitting results at 850–900 °C used Eq. (1); fitting results at 1000–1100 °C used Eq. (5)).
Fig. 2 shows the TG results of the oxidation study of Fe–16Cr alloy at 800–1000 °C under dry air condition. It was indicated that oxidation corrosion became more serious when elevating temperature. Apparently, the oxidation was turned into a quick mass gain stage after a slow one. In the second stage, the oxidation behavior of Fe–16Cr alloy at 1000 °C and 1100 °C showed parabolic curves initially, and then became straight lines at the remaining time, compared with the pure parabolic situations at 850 °C and 900 °C
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Z.-Y. Chen et al. / Journal of Alloys and Compounds 574 (2013) 437–442
Table 3 The character oxidation time tU of stage II of alloy oxidation in dry air. Unit tU1 (diffusion) tU2 (chem.) tt
ks ks ks
800 °C – – –
850 °C 9.341 10 – –
Fig. 8. The function of degradation time td to water vapor content compared with data estimated from Ref. [5] (experimental data: left y axis; Reference data: right y axis).
during 40 ks. The straight lines illustrated a more rapid oxidation happens. These results were also confirmed by the SEM observations, as shown in Fig. 2. A dense protective scale of Cr2O3 was formed in the first stage. And a thick scale contained Fe2+/Fe3+ was formed after a degradation time td. The second stage showed a break-away oxidation of the Fe– 16Cr alloy due to the destroyed protective layer, where the corrosion section of Fe-contained oxides appeared as Fig. 3 presented. Table 2 is the corresponding compositions of points 1–4 in Fig. 3. The oxides in the scale are FeCr2O4, (Fe, Cr)3O4 and Fe2O3 from points 2 to 4, respectively. The forming sequence of these oxides in the oxidation could be Cr2O3 ? FeCr2O4 ? (Fe, Cr)3O4 ? ferrous oxides, which was in accordance with the results by thermodynamic calculation. Fig. 4 shows the equilibrium mass of 100 g Fe–16Cr alloy continuously exposed to air at 1000 °C. The oxidation sequence of Fe–16Cr alloy in air is suggested as: MnSiO3 ? Mn2SiO4 ? Cr2O3/ SiO 2 ? FeCr 2 O 4 ? FeO/(FeO) 2 (SiO 2 )/Mn 2 SiO 4 ? Fe 3 O 4 ? Fe 2 O 3 / Cr2O3. The increase of Cr2O3 mass curve at beginning was related to the stage I of Fe–16Cr alloy oxidation. The decrease of the Cr2O3 mass curve was corresponded to the damage of the dense Cr2O3 scale in the further oxidation where the amounts of the FeCr2O4 increased. This degradation process corresponded to the transport
900 °C 5
1000 °C 4
3.348 10 – –
1100 °C 4
1.423 10 7.661 102 9.944
1.083 104 4.316 102 4.297
Fig. 10. The thickness of the scales peeling off the oxidized alloys as a function of water vapor content in air.
from stage I to stage II after the degradation time td. Next, the forming of ferrous oxides indicated the degradation of chormia. The Fe-rich layer could not protect the alloy from oxidation any more, and the corrosion was turning to be a chemical reaction controlling step after the critical turning time tt. The oxidation rate of Fe–16Cr alloy at 800 °C was very slow even when the exposure time was more than 120 h. Moreover, the break-away oxidation at the second kinetic stage was distinct at the temperature above 850 °C. Although Fe–16Cr alloy is usually applied in 600–800 °C region, the accelerated oxidation experiments were currently carried out at 850–1000 °C, since the predominance area diagram in the previous studies [2,14] had proved that the oxidation sequence of the elements in Fe–16Cr would not be modified by the elevating temperature. Fig. 5 showed the micrographs of scales formed after 50 h in dry oxygen at 1000 °C and 1100 °C. The scale surface at 1100 °C was rougher and more porous, illustrating the oxide developing quickly at higher temperature, while forest-like surface was surveyed in both of the two scales as Fig. 5a and b showed. Therefore, higher temperature results in faster oxidation, while the oxidation process obeys the same rule. The degradation time td, which is the beginning of break-away oxidation of Fe–16Cr interconnect in the current work, could also
Fig. 9. The SEM micrographs of the Fe–16Cr alloy oxide scale surface in H2O–Air at 1000 °C for 24 h ((a) surface in dry air and (b) surface in 2 ll H2O (l): 100 ml air).
Z.-Y. Chen et al. / Journal of Alloys and Compounds 574 (2013) 437–442
441
Fig. 11. Photos of the oxidized Fe–16Cr alloys in wet air at 1000 °C (from left to right: 0, 1, 2, 7 ll water (l) in 100 ml air (g)).
Here, an equation combined chemical reaction with diffusion control was used in this work to analyze the critical turning time tt of the rate-determining step in the oxidation. The controlling step was the diffusion of ions in the oxides scale before tt, and then interface chemical reaction follows afterwards.
(
Fig. 12. The kinetic equations (Eq. (5)) describing the alloy oxidation at 1000 °C in wet air.
be considered as an evaluation indicator of interconnect materials. It was found that the degradation time td was shortened with the increasing reaction temperature in the alloy oxidation process, from more than 120 h at 800 °C to 10 min at 1100 °C as Fig. 6 showed. The decrease of the degradation time td was slower at higher temperatures. The exponential relationship of the degradation time td to corrosion temperature was obtained, which is similar to the Tung’s work [9]. In order to figure out the mechanism of the second stage in oxidation, the RPP model was used in the present study, both diffusion controlling mechanism (Eq. (1)) and chemical reaction controlling mechanism (Eq. (3)). The oxidation kinetics of the second stage at 1000–1100 °C did not follow the diffusion mechanism (Eq. (1)) only, even in the case lower than 1000 °C. It is suggested that the diffusion of iron is not the sole control factor in the stage II process at 1000–1100 °C. The thickness of the oxides layer increased with the transformation of various oxides in the process. We suppose that the change of the kinetic mechanism is due to the serious break of the dense oxide scale on the alloy. There was no protective oxide layer to prevent the further oxidizing diffusion after tt, the gaps and holes in the oxidized scale accelerated the diffusion rate of the oxygen in this oxide layer, thus the interface oxidation reaction became the rate-determining step of the oxidation process.
pffiffiffiffiffiffiffiffiffiffiffi t=t U1 ; t 6 tt pffiffiffiffiffiffiffiffiffiffiffiffiffi n ¼ t t =tU1 þ ðt t t Þ=t U2 ; t > tt n¼
ð5Þ
This pattern showed a better fitting result than Eq. (1) as Fig. 7 shown. And a kinetic pattern of the Fe–16Cr alloy at second oxidation reaction stage is suggested: The rate-determining step is diffusion step (parabolic) initially, and then turns into interface chemical reaction (linear) with the increase of exposure time. Oxidation of Fe–16Cr alloy in stage II could be observed clearly in the accelerated oxidation experiments at 1000 °C and 1100 °C. The fitting parameters of Fig. 7 are given in Table 3. Corresponding to the sharply increasing oxidation rate of the diffusion controlling step to the reaction temperature, the characteristic oxidation time tU1 was in an exponential descent. The characteristic time of chemical controlled step tU2 and the turning time tt decreased with the increase of the oxidation temperature also. It is supposed that a porous Fe-richer oxide scale at a higher temperature would result in the shorter turning time tt of rate-determining step of oxidation from diffusion in the scale to the interface chemical reaction and the faster interface chemical reaction. 4.2. Effect of water vapor content in air The oxidation experiments of Fe–16Cr alloy in air with different water vapor contents were also carried out at 1000 °C. The results showed that, water vapor had strong effect on the oxidation rate. The higher contents of water vapor in the air, the faster oxidation rate at the second reaction stage, and the shorter of the degradation time td from the first stage to the second stage. Fig. 8 shows that the degradation time td is sensitive to the water–vapor content in the atmosphere [5]. And the linear relationship of the degradation time td to the water vapor content in the wet air is obtained as following:
td ¼ 1:47 0:102V water
ð6Þ
where the unit of td is ks, Vwater is the water vapor content in the atmosphere with the unit of vol%. Compared with the case in dry air mentioned above, the fact that the water vapor accelerates the oxidation corrosion is confirmed. Moreover, the oxidation sequence
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Z.-Y. Chen et al. / Journal of Alloys and Compounds 574 (2013) 437–442
Table 4 The character oxidation time tU of stage II of alloy oxidation in wet air. Kinetic step tU1 (diffusion) tU2 (chem.) tt
1st Diffusion 2nd Chem. Turning time
Unit
1.2 vol%
ks ks ks
of Fe–16Cr alloy isn’t modified by water based on the thermodynamic calculation. Regarding the effect of water vapor on the oxidation products, two functions are suggested, which result in the accelerated corrosions. One is providing extra oxygen in the alloy oxidation process; the other is affecting and even modifying the structures of oxide scale due to the existence of hydrogen atom. The second function can be observed through the SEM micrographs (Fig. 9) of the oxide scale surfaces of the Fe–16Cr alloy exposed in H2O–Air at 1000 °C for 24 h, in which surface topography of the Fe-rich oxide scale is modified by the water vapor. The forest-like surface is collapsing with the increase of moisture capacity. The scales with coarser grains and cavern are observed in the wet air as well. The thicknesses of oxidized scales stripped off are measured with the accuracy of 0.005 mm. The average values of every 12 measurements are shown in Fig. 10. The increasing corrosion mass gain is not directly related to the increased thickness of oxidized scale. As Fig. 11 shows, the color1 of scale is turning from velvety black to dark gray, then to steel gray finally with the increase of water vapor content in the corrosion process. The collapse of the microscopic ‘‘forest’’ changes the surface gloss. It is corresponding to Fig. 9. The alloy under the oxide scale is exposed at parts of the alloy where the scale is stripped off. The color of alloy is from golden to gray with the increase of the moisture capacity in the high temperature oxidation. In other words, the oxidation product sticking on the metal surface is changing from ferric chromate to ferriferous oxide. Oxidation of Fe–Cr alloy contains external oxidation occurred at the interface of air/scale as well as internal oxidation of the alloy components at the interface of scale/alloy [24,25]. The oxidized scale becomes coarser (Fig. 9) and its thickness is reduced (Fig. 10), which results in the worse internal alloy oxidation. This is also confirmed by the observed precipitation of the subscale oxide in Fig. 11. This interesting phenomenon suggests that more oxygen atoms diffuse into the alloy, and a more serious internal oxidation at scale/alloy interface happens, which results in the faster mass-gain of the whole alloy. It is evident that higher water vapor content in wet air makes the corrosion of Fe–16Cr alloy more serious. The theoretical descriptions of Fe–16Cr alloy oxidation kinetics are shown in Fig. 12. The shapes of the mass gain curves are parabolic initially, and then followed by a linear process. It is indicated that the oxidation stage II in wet air is similar to the kinetics in the dry air. The broken ferrous scale results in the mechanism change of rate-determining step from diffusion of iron in the oxide scale to the internal chemical reaction at scale/alloy interface. A good fitting of Eq. (5) is also shown in Fig. 12. tU1, tU2 and tt described in Eq. (5) increase with the increase of H2O% in wet air. These parameters are shown in Table 4. It can be found that the mass gain curve follows parabolic relationship at high water vapor content corrosion experiment, which reflects that diffusion of ions in oxidized scale is the main ratedetermining step in the second oxidation stage instead of chemical reaction. However, it seems contrary to the more serious corrosion observed in the study. The characteristic time tU1 of the diffusion controlled step is decreased sharply with the water vapor com1 For interpretation of color in Fig. 11, the reader is referred to the web version of this article.
2.5 vol% 3
2.256 10 9.899 102 37.043
8.0 vol% 3
2.054 10 8.471 102 32.250
1.659 103 6.268 102 59.118
pared with the tiny variation tU2 of the chemical reaction controlled step. It could be explained that the influence water vapor to interface chemical reaction is less than it to internal diffusion step in the Fe–16Cr alloy oxidized scale. 5. Conclusions The oxidation experimental studies of Fe–16Cr alloy in dry and wet air were carried out in the current work. The complete oxidation process of the alloy includes two stages, in which the degradation time td from the slow oxidation stage to the break-away oxidation stage exponentially decreases with elevating corrosion temperature in the dry air. In the case of oxidation in wet air, the degradation time td shows a linear relationship to water content. The oxides transformation order calculated in the oxidation of Fe–16Cr alloy in dry and wet air is consistent with the experimental results. RPP model is applied in analyzing the mechanism of oxidation Fe–16Cr alloy, and exhibits advantages for clear physical meaning of parameters. The kinetics of Fe–16Cr alloy oxidation stage II follows diffusion controlling mechanism at initial, then interface chemical reaction controlling mechanism. Acknowledgement The authors gratefully acknowledge the financial support of National Program on Key Basic Research Project of China (Grant No. 2012CB215405). References [1] [2] [3] [4] [5] [6] [7] [8] [9] [10] [11] [12] [13] [14] [15] [16] [17] [18] [19] [20] [21] [22] [23] [24] [25]
K. Honda, T. Maruyama, T. Atake, Y. Saito, Oxid. Met. 35 (1992) 347. H. Kurokawa, Solid State Ionics 168 (2004) 13. P. Jian, L. Jian, H. Bing, G. Xie, J. Power Sources 158 (2006) 354. T. Horita, Y. Xiong, H. Kishimoto, K. Yamaji, N. Sakai, H. Yokokawa, J. Power Sources 131 (2004) 293. S. Jianian, Z. Longjiang, L. Tiefan, Oxid. Met. 48 (1997) 347. T. Horita, Y. Xiong, K. Yamaji, N. Sakai, H. Yokokawa, J. Power Sources 118 (2003) 35. T. Brylewski, J. Dabek, K. Przybrlski, J. Therm. Anal. Calorim. 77 (2004) 207. T. Brylewski, T. Maruyama, M. Nanko, K. Przybylski, J. Therm. Anal. Calorim. 55 (1999) 681. H.-M. Tung, J.F. Stubbins, Mater. Sci. Eng. A 538 (2012) 1. C. Tedmon, J. Electrochem. Soc. 114 (1967) 788. Z. Yang, K.S. Weil, D.M. Paxton, J.W. Stevenson, J. Electrochem. Soc. 150 (2003) A1188. J.W. Fergus, Mater. Sci. Eng. A 397 (2005) 271. B. Pujilaksono, T. Jonsson, H. Heidari, M. Halvarsson, J.E. Svensson, L.G. Johansson, Oxid. Met. 75 (2011) 183. _ M. Hänsel, W. Quadakkers, Oxid. Met. 69 (2008) E. Essuman, G. Meier, J. Zurek, 143. J. Gong, Y.M. Jiang, B. Deng, C. Zhong, D.M. Sun, J. Li, Metall. Mater. Trans. A 40 (2009) 2511. K.C. Chou, Q. Li, Q. Lin, L. Jiang, K. Xu, Int. J. Hydrogen Energy 30 (2005) 301. K.C. Chou, J. Am. Ceram. Soc. 89 (2006) 1568. K.C. Chou, K. Xu, Intermetallics 15 (2007) 767. X.M. Hou, K.C. Chou, X. Zhong, S. Seetharaman, J. Alloys Comp. 465 (2008) 90. X.M. Hou, K.C. Chou, X.J. Hu, H.l. Zhao, J. Alloys Comp. 459 (2008) 123. X.M. Hou, K.C. Chou, J. Eur. Ceram. Soc. 29 (2009) 517. X.M. Hou, K.C. Chou, M. Zhang, Int. J. Appl. Ceram. Technol. 7 (2010) 248. X.M. Hou, K.C. Chou, F.S. Li, Ceram. Int. 35 (2009) 603. T. Brylewski, M. Nanko, T. Maruyama, K. Przybylski, Solid State Ionics 143 (2001) 131. J. Hammer, S. Laney, R. Jackson, K. Coyne, F. Pettit, G. Meier, Oxid. Met. 67 (2007) 1.
Contents lists available at SciVerse ScienceDirect
Journal of Alloys and Compounds journal homepage: www.elsevier.com/locate/jalcom
Oxidation mechanism of Fe–16Cr alloy as SOFC interconnect in dry/wet air Zhi-Yuan Chen a, Li-Jun Wang a,⇑, Fu-Shen Li b, Kuo-Chih Chou a a b
School of Metallurgical and Ecological Engineering, University of Science and Technology Beijing, Beijing 100083, PR China School of Material Science Engineering, University of Science and Technology Beijing, Beijing 100083, PR China
a r t i c l e
i n f o
Article history: Received 13 March 2013 Received in revised form 4 May 2013 Accepted 12 May 2013 Available online 20 May 2013 Keywords: Break-away oxidation Real physical picture model Oxidization SOFC interconnector
a b s t r a c t Experimental study on the oxidation corrosions of Fe–16Cr alloy was carried out at 800–1100 °C under dry/wet air conditions. Faster oxidation rate was observed at higher temperature and water vapor content. The degradation time td between two stages in oxidation process showed an exponential relationship with elevating corrosion temperature in dry air, and a linear relationship with the water content in the case of water vapor introduced to the system. The mechanism of oxidation corrosions of Fe–16Cr alloy was suggested by the Real Physical Picture (RPP) model. It was found that the break-away oxidation in stage II was controlled by diffusion at initial both in dry and wet air, then became linear with the exposure time, which implied that the oxidation rate was then controlled by chemical reaction of the interface between the metal and the oxidized scale. Moreover, the effect of water in the oxidation process is not only to supply more oxygen into system, but also to modify the structures of oxide scale due to the existence of hydrogen atom, which results in the accelerated corrosions. Ó 2013 Elsevier B.V. All rights reserved.
1. Introduction Alloy interconnect attracts a great deal of attention for commercial use of SOFC due to its several advantages compared to ceramics, such as gas tightness, and machinability. Alloy interconnect can be used below 800 °C where Fe–Cr alloys exhibit good oxidation resistance. Fe–16Cr alloy is one of promising alloy interconnects for IT-SOFC. A number of studies [1–4] have been carried out on the oxidation behavior of the Fe–16Cr alloy under various atmospheres. As it is known, the oxidation rate of the metallic interconnect at relative high temperature is of great importance to the stable operation of SOFC. The oxidation process of Fe–Cr alloy usually contains two stages—one is a smaller mass-gain rate stage, the other is with larger rate followed [1,3,5]. The mass-gain rate of the second stage denotes the broken rate of the dense scale. And the degradation time td of the first stage to the second can be regarded as the beginning of protective layer breaking. So it might be a key parameter to define the oxidation resistance other than rate constant. Although oxidation kinetics of the Fe–Cr alloys has been investigated in many researches [1,6–8], information on the degradation time td of the first stage to the second one was still limited [5,9].
⇑ Corresponding author. Tel.: +86 10 6233 3622; fax: +86 6233 2570. E-mail address: [email protected] (L.-J. Wang). 0925-8388/$ - see front matter Ó 2013 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.jallcom.2013.05.072
Researchers [8–11] pointed out the mass-gain of Fe–Cr alloys was parabolic to the exposure time, which indicated that the diffusion of the ions through the oxidate layer was the rate-determining step according to Wagner’s theory. Meanwhile, in view of the vaporization of the chromia in the alloy oxidation process, a more complex kinetic law was also addressed [1,12]. The cathode of SOFC is operating at atmosphere more or less containing water vapor. When water was introduced into the air, the growth rate of alloy scale was observed to be increased with the water vapor content at high temperature [5,13]. It suggested the oxidation rate of Fe–16Cr alloy is accelerated in the atmosphere containing water vapor due to the break-down of protective layer. Water vapor can affect the oxidation behavior in several approaches, such as influencing the transport processes of ions in the alloy and modifying the void formation in the oxide scale [14,15]. The aim of the current work is to study the kinetics of the second breakaway stage of Fe–16Cr alloy oxidation at dry/wet air, and obtain the relationship of the degradation time td with two factors: temperature and water vapor content.
2. Experimental The chemical composition of Fe–16Cr alloy used in the present study is shown in Table 1. The main elements were also analyzed by National Analysis Center of Iron and Steel of China. The samples were cut to the rectangular sheet, 21 42 1.3 mm, and were ground with SiC abrasive paper up to #2000. After polishing, the surfaces of the samples were cleaned by alcohol.
438
Z.-Y. Chen et al. / Journal of Alloys and Compounds 574 (2013) 437–442
Table 1 Chemical composition of Fe–16Cr alloy. Elements
C
Si
Cr
Ni
Mn
S
P
Fe
Mass% Checked mass%
0.05 unchecked
0.40 0.31
16.05 15.98
0.05 0.089
0.25 0.30
0.005 unchecked
0.012 unchecked
Remnant Remnant
Fig. 3. SEM photo of the cross section of Fe–16Cr alloy scale oxidized in air for 1 h at 1000 °C.
Table 2 EDS results corresponding to the points of the cross section Fe–16Cr alloy scale oxidized in air for 1 h at 1000 °C at Fig. 3. Elements (wt%)
1
2
3
4
Fe Cr O
84.80 15.20 –
29.60 38.43 31.97
66.32 6.23 27.45
65.85 0.79 33.36
Fig. 1. The device of the alloy sample hanging on the balance.
Fig. 2. The mass changing curves of Fe–16Cr alloy in dry air at different temperatures.
Sample arrangement for oxidation under the dry/wet air atmosphere is illustrated in Fig. 1. An alumina stick (diameter is 2 mm) through the hole of the sample sheet was connected to the OHAUS AR153CN balance (accuracy 1 mg) using an alumina pipe. The mass gain of the sample sheet was recorded during oxidation. The inner area of the drilled hole was ignored in the following calculation due to the tight fit between the hole and the alumina stick. Oxidation experiments were
Fig. 4. Equilibrium diagram of 100 g Fe–Cr alloy continuously exposure to air at 1000 °C.
carried out in the temperature region of 800–1100 °C under dry air and at 1000 °C under wet air, respectively. The gas flow was controlled to 100 ml/min at 25 °C (accuracy ± 1.8 ml/min), by Alicat Gas Flow Controller. The water was pumped into a preheating chamber in 1–7 ll/min using a MasterFlex peristaltic
439
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Fig. 5. The micrographs of oxide scales at 1000 °C and 1100 °C for 50 h ((a) scale surface at 1000 °C and (b) scale surface at 1100 °C).
rate-determining step of the reaction. A parameter of rate constant ‘‘k’’ was used, in which all effects of factors involved. This kind of treatment is too complicated to offer an explicit analytic expression and an intuitionistic quantitative discussion. More recently, Chou and Hou [16–22] proposed a Real Physical Picture model (RPP) to solve the problem mentioned above, in which each parameters had clear physical meaning and is quite convenient for the theoretical discussions. A more substantive parameter called ‘‘characteristic oxidation time’’ was introduced into the model to characterize the oxidation reaction kinetics. According to RPP model, the reaction fraction n can be expressed as following [23]:
n¼
Dm pffiffiffiffiffiffiffiffiffiffiffi ¼ t=t U1 m
t U1 ¼ Fig. 6. Degradation time td between the two stages oxidation as a function of temperature. pump (accuracy ± 0.1%), mixing with dry air to obtain wet air, which contained 1.2 vol%, 2.5 vol%, and 8.0 vol% water vapor (1, 2, and 7 ll water (l) per 100 ml air (g)), respectively. Microstructure observation after oxidation was carried out by a scanning electron microscope (SEM, CARL ZEISS EVO MA 10/LS 10 JS) with energy dispersive spectrum (EDS, Thermo NORAN System).
3. Kinetic models The kinetic mechanism of Fe–Cr alloy oxidation was usually suggested that the diffusion of ion in the oxide scale was the
tm L20
eq
DðC C Þ
¼
ð1Þ
tm L20 D0 ðC C eq Þ
exp
DE1 RT
ð2Þ
where tU is named as the ‘‘characteristic oxidation time’’. When t = tU, the reaction fraction n is equal to 1. D is the diffusion coefficient of the certain component, here D0 is a constant. C and Ceq are the diffusion atom activities at alloy surface and alloy/scale interface with equilibrium condition, respectively. The rest of parameters are related to the external physical characteristic, where tm is a coefficient related to the density of the substance and reaction, L0 is the original thickness of the alloy. If the rate-determining step of the oxidation is a chemical reaction, the relationship of the reaction fraction to exposure time is as following [23]:
n ¼ t=t U2 t U2 ¼
ð3Þ
tm L20 KðC 0 C eq Þ
¼
tm L20 K 0 ðC 0 C eq Þ
exp
DE2 RT
ð4Þ
where K is the equilibrium constant, K0 is a constant, C0 is the reactant activity in the interface oxidation reaction. Eqs. (1) and (3) are used to analyze the kinetics of the alloy oxidation in the following chapters. 4. Results and discussion 4.1. Oxidation corrosion in dry air
Fig. 7. The kinetic equations describing the alloy oxidation at different temperatures (fitting results at 850–900 °C used Eq. (1); fitting results at 1000–1100 °C used Eq. (5)).
Fig. 2 shows the TG results of the oxidation study of Fe–16Cr alloy at 800–1000 °C under dry air condition. It was indicated that oxidation corrosion became more serious when elevating temperature. Apparently, the oxidation was turned into a quick mass gain stage after a slow one. In the second stage, the oxidation behavior of Fe–16Cr alloy at 1000 °C and 1100 °C showed parabolic curves initially, and then became straight lines at the remaining time, compared with the pure parabolic situations at 850 °C and 900 °C
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Table 3 The character oxidation time tU of stage II of alloy oxidation in dry air. Unit tU1 (diffusion) tU2 (chem.) tt
ks ks ks
800 °C – – –
850 °C 9.341 10 – –
Fig. 8. The function of degradation time td to water vapor content compared with data estimated from Ref. [5] (experimental data: left y axis; Reference data: right y axis).
during 40 ks. The straight lines illustrated a more rapid oxidation happens. These results were also confirmed by the SEM observations, as shown in Fig. 2. A dense protective scale of Cr2O3 was formed in the first stage. And a thick scale contained Fe2+/Fe3+ was formed after a degradation time td. The second stage showed a break-away oxidation of the Fe– 16Cr alloy due to the destroyed protective layer, where the corrosion section of Fe-contained oxides appeared as Fig. 3 presented. Table 2 is the corresponding compositions of points 1–4 in Fig. 3. The oxides in the scale are FeCr2O4, (Fe, Cr)3O4 and Fe2O3 from points 2 to 4, respectively. The forming sequence of these oxides in the oxidation could be Cr2O3 ? FeCr2O4 ? (Fe, Cr)3O4 ? ferrous oxides, which was in accordance with the results by thermodynamic calculation. Fig. 4 shows the equilibrium mass of 100 g Fe–16Cr alloy continuously exposed to air at 1000 °C. The oxidation sequence of Fe–16Cr alloy in air is suggested as: MnSiO3 ? Mn2SiO4 ? Cr2O3/ SiO 2 ? FeCr 2 O 4 ? FeO/(FeO) 2 (SiO 2 )/Mn 2 SiO 4 ? Fe 3 O 4 ? Fe 2 O 3 / Cr2O3. The increase of Cr2O3 mass curve at beginning was related to the stage I of Fe–16Cr alloy oxidation. The decrease of the Cr2O3 mass curve was corresponded to the damage of the dense Cr2O3 scale in the further oxidation where the amounts of the FeCr2O4 increased. This degradation process corresponded to the transport
900 °C 5
1000 °C 4
3.348 10 – –
1100 °C 4
1.423 10 7.661 102 9.944
1.083 104 4.316 102 4.297
Fig. 10. The thickness of the scales peeling off the oxidized alloys as a function of water vapor content in air.
from stage I to stage II after the degradation time td. Next, the forming of ferrous oxides indicated the degradation of chormia. The Fe-rich layer could not protect the alloy from oxidation any more, and the corrosion was turning to be a chemical reaction controlling step after the critical turning time tt. The oxidation rate of Fe–16Cr alloy at 800 °C was very slow even when the exposure time was more than 120 h. Moreover, the break-away oxidation at the second kinetic stage was distinct at the temperature above 850 °C. Although Fe–16Cr alloy is usually applied in 600–800 °C region, the accelerated oxidation experiments were currently carried out at 850–1000 °C, since the predominance area diagram in the previous studies [2,14] had proved that the oxidation sequence of the elements in Fe–16Cr would not be modified by the elevating temperature. Fig. 5 showed the micrographs of scales formed after 50 h in dry oxygen at 1000 °C and 1100 °C. The scale surface at 1100 °C was rougher and more porous, illustrating the oxide developing quickly at higher temperature, while forest-like surface was surveyed in both of the two scales as Fig. 5a and b showed. Therefore, higher temperature results in faster oxidation, while the oxidation process obeys the same rule. The degradation time td, which is the beginning of break-away oxidation of Fe–16Cr interconnect in the current work, could also
Fig. 9. The SEM micrographs of the Fe–16Cr alloy oxide scale surface in H2O–Air at 1000 °C for 24 h ((a) surface in dry air and (b) surface in 2 ll H2O (l): 100 ml air).
Z.-Y. Chen et al. / Journal of Alloys and Compounds 574 (2013) 437–442
441
Fig. 11. Photos of the oxidized Fe–16Cr alloys in wet air at 1000 °C (from left to right: 0, 1, 2, 7 ll water (l) in 100 ml air (g)).
Here, an equation combined chemical reaction with diffusion control was used in this work to analyze the critical turning time tt of the rate-determining step in the oxidation. The controlling step was the diffusion of ions in the oxides scale before tt, and then interface chemical reaction follows afterwards.
(
Fig. 12. The kinetic equations (Eq. (5)) describing the alloy oxidation at 1000 °C in wet air.
be considered as an evaluation indicator of interconnect materials. It was found that the degradation time td was shortened with the increasing reaction temperature in the alloy oxidation process, from more than 120 h at 800 °C to 10 min at 1100 °C as Fig. 6 showed. The decrease of the degradation time td was slower at higher temperatures. The exponential relationship of the degradation time td to corrosion temperature was obtained, which is similar to the Tung’s work [9]. In order to figure out the mechanism of the second stage in oxidation, the RPP model was used in the present study, both diffusion controlling mechanism (Eq. (1)) and chemical reaction controlling mechanism (Eq. (3)). The oxidation kinetics of the second stage at 1000–1100 °C did not follow the diffusion mechanism (Eq. (1)) only, even in the case lower than 1000 °C. It is suggested that the diffusion of iron is not the sole control factor in the stage II process at 1000–1100 °C. The thickness of the oxides layer increased with the transformation of various oxides in the process. We suppose that the change of the kinetic mechanism is due to the serious break of the dense oxide scale on the alloy. There was no protective oxide layer to prevent the further oxidizing diffusion after tt, the gaps and holes in the oxidized scale accelerated the diffusion rate of the oxygen in this oxide layer, thus the interface oxidation reaction became the rate-determining step of the oxidation process.
pffiffiffiffiffiffiffiffiffiffiffi t=t U1 ; t 6 tt pffiffiffiffiffiffiffiffiffiffiffiffiffi n ¼ t t =tU1 þ ðt t t Þ=t U2 ; t > tt n¼
ð5Þ
This pattern showed a better fitting result than Eq. (1) as Fig. 7 shown. And a kinetic pattern of the Fe–16Cr alloy at second oxidation reaction stage is suggested: The rate-determining step is diffusion step (parabolic) initially, and then turns into interface chemical reaction (linear) with the increase of exposure time. Oxidation of Fe–16Cr alloy in stage II could be observed clearly in the accelerated oxidation experiments at 1000 °C and 1100 °C. The fitting parameters of Fig. 7 are given in Table 3. Corresponding to the sharply increasing oxidation rate of the diffusion controlling step to the reaction temperature, the characteristic oxidation time tU1 was in an exponential descent. The characteristic time of chemical controlled step tU2 and the turning time tt decreased with the increase of the oxidation temperature also. It is supposed that a porous Fe-richer oxide scale at a higher temperature would result in the shorter turning time tt of rate-determining step of oxidation from diffusion in the scale to the interface chemical reaction and the faster interface chemical reaction. 4.2. Effect of water vapor content in air The oxidation experiments of Fe–16Cr alloy in air with different water vapor contents were also carried out at 1000 °C. The results showed that, water vapor had strong effect on the oxidation rate. The higher contents of water vapor in the air, the faster oxidation rate at the second reaction stage, and the shorter of the degradation time td from the first stage to the second stage. Fig. 8 shows that the degradation time td is sensitive to the water–vapor content in the atmosphere [5]. And the linear relationship of the degradation time td to the water vapor content in the wet air is obtained as following:
td ¼ 1:47 0:102V water
ð6Þ
where the unit of td is ks, Vwater is the water vapor content in the atmosphere with the unit of vol%. Compared with the case in dry air mentioned above, the fact that the water vapor accelerates the oxidation corrosion is confirmed. Moreover, the oxidation sequence
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Table 4 The character oxidation time tU of stage II of alloy oxidation in wet air. Kinetic step tU1 (diffusion) tU2 (chem.) tt
1st Diffusion 2nd Chem. Turning time
Unit
1.2 vol%
ks ks ks
of Fe–16Cr alloy isn’t modified by water based on the thermodynamic calculation. Regarding the effect of water vapor on the oxidation products, two functions are suggested, which result in the accelerated corrosions. One is providing extra oxygen in the alloy oxidation process; the other is affecting and even modifying the structures of oxide scale due to the existence of hydrogen atom. The second function can be observed through the SEM micrographs (Fig. 9) of the oxide scale surfaces of the Fe–16Cr alloy exposed in H2O–Air at 1000 °C for 24 h, in which surface topography of the Fe-rich oxide scale is modified by the water vapor. The forest-like surface is collapsing with the increase of moisture capacity. The scales with coarser grains and cavern are observed in the wet air as well. The thicknesses of oxidized scales stripped off are measured with the accuracy of 0.005 mm. The average values of every 12 measurements are shown in Fig. 10. The increasing corrosion mass gain is not directly related to the increased thickness of oxidized scale. As Fig. 11 shows, the color1 of scale is turning from velvety black to dark gray, then to steel gray finally with the increase of water vapor content in the corrosion process. The collapse of the microscopic ‘‘forest’’ changes the surface gloss. It is corresponding to Fig. 9. The alloy under the oxide scale is exposed at parts of the alloy where the scale is stripped off. The color of alloy is from golden to gray with the increase of the moisture capacity in the high temperature oxidation. In other words, the oxidation product sticking on the metal surface is changing from ferric chromate to ferriferous oxide. Oxidation of Fe–Cr alloy contains external oxidation occurred at the interface of air/scale as well as internal oxidation of the alloy components at the interface of scale/alloy [24,25]. The oxidized scale becomes coarser (Fig. 9) and its thickness is reduced (Fig. 10), which results in the worse internal alloy oxidation. This is also confirmed by the observed precipitation of the subscale oxide in Fig. 11. This interesting phenomenon suggests that more oxygen atoms diffuse into the alloy, and a more serious internal oxidation at scale/alloy interface happens, which results in the faster mass-gain of the whole alloy. It is evident that higher water vapor content in wet air makes the corrosion of Fe–16Cr alloy more serious. The theoretical descriptions of Fe–16Cr alloy oxidation kinetics are shown in Fig. 12. The shapes of the mass gain curves are parabolic initially, and then followed by a linear process. It is indicated that the oxidation stage II in wet air is similar to the kinetics in the dry air. The broken ferrous scale results in the mechanism change of rate-determining step from diffusion of iron in the oxide scale to the internal chemical reaction at scale/alloy interface. A good fitting of Eq. (5) is also shown in Fig. 12. tU1, tU2 and tt described in Eq. (5) increase with the increase of H2O% in wet air. These parameters are shown in Table 4. It can be found that the mass gain curve follows parabolic relationship at high water vapor content corrosion experiment, which reflects that diffusion of ions in oxidized scale is the main ratedetermining step in the second oxidation stage instead of chemical reaction. However, it seems contrary to the more serious corrosion observed in the study. The characteristic time tU1 of the diffusion controlled step is decreased sharply with the water vapor com1 For interpretation of color in Fig. 11, the reader is referred to the web version of this article.
2.5 vol% 3
2.256 10 9.899 102 37.043
8.0 vol% 3
2.054 10 8.471 102 32.250
1.659 103 6.268 102 59.118
pared with the tiny variation tU2 of the chemical reaction controlled step. It could be explained that the influence water vapor to interface chemical reaction is less than it to internal diffusion step in the Fe–16Cr alloy oxidized scale. 5. Conclusions The oxidation experimental studies of Fe–16Cr alloy in dry and wet air were carried out in the current work. The complete oxidation process of the alloy includes two stages, in which the degradation time td from the slow oxidation stage to the break-away oxidation stage exponentially decreases with elevating corrosion temperature in the dry air. In the case of oxidation in wet air, the degradation time td shows a linear relationship to water content. The oxides transformation order calculated in the oxidation of Fe–16Cr alloy in dry and wet air is consistent with the experimental results. RPP model is applied in analyzing the mechanism of oxidation Fe–16Cr alloy, and exhibits advantages for clear physical meaning of parameters. The kinetics of Fe–16Cr alloy oxidation stage II follows diffusion controlling mechanism at initial, then interface chemical reaction controlling mechanism. Acknowledgement The authors gratefully acknowledge the financial support of National Program on Key Basic Research Project of China (Grant No. 2012CB215405). References [1] [2] [3] [4] [5] [6] [7] [8] [9] [10] [11] [12] [13] [14] [15] [16] [17] [18] [19] [20] [21] [22] [23] [24] [25]
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