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Effect of temperature on hardness of binary U–15%Pu alloy and T91 cladding
Journal of Nuclear Materials 429 (2012) 341–345
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Journal of Nuclear Materials journal homepage: www.elsevier.com/locate/jnucmat
Effect of temperature on hardness of binary U–15%Pu alloy and T91 cladding T.R.G. Kutty ⇑, K. Ravi, Santu Kaity, S.K. Swarnkar, Arun Kumar Radiometallurgy Division, Bhabha Atomic Research Centre, Mumbai 400 085, India
h i g h l i g h t s " Hardness–temperature plot of T91 alloy showed a transition at 778 K. " Hardness–temperature plot of
a phase of U–15%Pu showed a transition at 408 K.
" b Phase of U–15%Pu alloy showed a discontinuity in hardness data above 828 K. "
c Phase region of U–15%Pu was found to be very soft.
" At reactor operating temperatures, U–15%Pu fuel is softer than T91 cladding.
a r t i c l e
i n f o
Article history: Received 18 April 2012 Accepted 15 June 2012 Available online 23 June 2012
a b s t r a c t The high temperature hardness behaviour of U–15%Pu and T91 cladding was studied with the help of a hot hardness tester. The hardness versus temperature plot for the T91 alloy consists of two straight lines with different slopes intersecting at 778 K. The hot hardness data of orthorhombic a phase of U–15%Pu showed a transition at around 408 K. The appearance of the tetragonal b-phase showed a sudden increase in the hardness above 828 K. c-Phase of U–15%Pu was found to very soft. Since U–15%Pu fuel is softer than the T91 cladding above 473 K, the probability of the clad damage by the fuel due to its interaction with the cladding is very less. Ó 2012 Elsevier B.V. All rights reserved.
1. Introduction For the rapid growth of fast reactor programme in India, it is essential to shift to the use of metallic fuels in fast breeder reactors (FBR), which gives a higher breeding ratio (BR) and lower doubling time (DT) [1,2]. In FBR, the basic thermal and neutronic performance of metallic fuels is better than oxide ceramic fuels. The harder spectrum in the metallic fuelled core results in high breeding ratio [3,4]. India is going ahead with two types of fast reactor fuel element designs, which are under active consideration, viz. (i) sodium bonded U–15%Pu–6%Zr alloy and (ii) mechanically bonded binary U–15%Pu binary alloy [1,2]. For the mechanically bonded binary alloy, a Zr liner is proposed between the fuel pellet and clad to reduce the fuel–clad interaction. Modified T91, which is a 9Cr1Mo ferritic martensitic steel, has been used as the cladding material. Smear density, which is a key parameter for accommodation of fuel swelling, varies between 80% and 85% for mechanical bonded fuel and 75% for sodium bonded fuel. Since achieving a higher breeding ratio is one of the goals of India’s fast reactor programme, and breeding ratio increases linearly with the reduction of zirconium content in the fuel, a ternary fuel with a lower Zr ⇑ Corresponding author. Tel.: +91 22 2559 5361; fax: +91 22 2550 5151. E-mail Kutty).
addresses:
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0022-3115/$ - see front matter Ó 2012 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.jnucmat.2012.06.025
(T.R.G.
content is the preferred alloy [1]. In this study, emphasis has been given to study the mechanical properties of binary U–15%Pu alloy. It is well known that metallic fuel swells when irradiated due to the accumulation of fission products, especially the inert gaseous fission products. When total fuel swelling reaches the allowable limit due to gap closure, the contact stress between the fuel and the cladding starts to increase with the burnup [5–7]. During this period, swelling by closed bubbles decreases due to increase in external pressure on the closed bubble. The volume fraction of porosity remains relatively the same, until they are filled by solid fission products. It should be mentioned that the build up of the solid fission products is a very important part of the fuel swelling. Fission gas swelling saturates early on irradiation, whereas solid fission products continue to building up uniformly due to fission event causing volume expansion. As a result, fuel/clad contact pressure increases causing creep induced hot pressing of the fuel and open porosity decrease of the fuel [5,6]. It may be noted that there should be a dynamic equilibrium, where open pores are filled, new ones are formed and closed pores connect each other. Operation with an incompressible fuel may cause a very significant fuel clad mechanical interaction (FCMI). Hence one need to adjust the fuel smear density carefully so that fuel will be rather compressible by the end of the life [7]. FCMI is a significant issue during normal operation as well as in case of transient overpower accidents. Interaction between fuel and cladding that may lead to
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cladding failure can be both mechanical and chemical. Fuel–clad contact stresses can be minimised by maintaining a low (<75%) smear density, which guarantees a spongy fuel microstructure with open porosity [3]. In fast reactor, the metallic fuel forms various phases under the influence of the radial temperature gradient [8,9]. Among them, single BCC c-phase, which dominates at the central parts of the fuel, has a high creep rate; therefore, rendering the fuel rather compliant. On the other hand, at outer periphery of the fuel pin, the stiff a phase may dominate, reducing fuel compressibility and delaying bubble growth, coalescence and formation of interconnected porosity [10]. In this study, an attempt has been made to evaluate the strength of the U–15%Pu fuel and T91 cladding at various temperatures by measuring their hardness. In general, hardness of both the fuel and cladding are required to fully describe the gap conductance and contact resistance, once the mechanical contact between the fuel and cladding has occurred [11,12]. Hardness is an easily measured mechanical property that can provide insight into the strength of the fuel and cladding. Since hardness is related to yield stress, UTS, stress exponent and other mechanical properties by simple empirical relations, prediction of these mechanical properties from hardness data will be very useful for radiotoxic materials where conventional testing is very difficult to perform. The hot hardness data further give information on softening behaviour [13] and can be correlated to the Larson–Miller and Sherby–Dorn parameters to predict the long term rupture and creep properties. The main objectives of the present study are: i. Determination of the hot-hardness data of the U–15%Pu alloy from room temperature to 1000 K. ii. Evaluation of hot hardness of T91 cladding from room temperature to 1073 K. iii. From these data, an attempt will be made to assess fuel–clad mechanical interaction between U–15%Pu fuel and T91 cladding. The hardness values for the as-cast ternary U–15%Pu–6.8%Zr (composition in wt.%) was measured by Harbur et al. [14] between 673 and 1073 K using a load of 200 g. However, no studies have been reported on the mechanical properties of binary U–15%Pu alloy. Hence this study will be very useful for the fuel designers. 2. Experimental procedure U–15%Pu alloys were made by melting and casting route. The U and Pu slugs were melted together using induction melting in an yttria coated graphite crucible. The metallic impurities of the alloy were revealed by atomic emission spectroscopy (AES) using DC-arc as an excitation source. The phase analysis of the fuel alloy was carried out using XRD. For this, Cu Ka radiation was used with fixed slit optics and h–h goniometer. Diffraction patterns were obtained by the step-scan method with a step size of 0.01° and a measuring time of 1 s at each step were used. For metallographic analyses, the alloy slugs were cut into disc of suitable thickness using slow speed SiC abrasive cut-off wheel. Standard metallographic techniques were adopted for grinding and polishing using cold setting resin mounts. The chemical composition of T91 used in this study is shown in Table 1. The heat treatment for cladding material, T91, consists of austenization at 1323 K and air quench-
ing followed by tempering at 1023 K for 1 h. Grain size of T91 was found to be around 30 lm and its hardness measured using Vicker’s hardness tester at a test load of 200 g was 220 kg mm2. T91 steels are used in the normalised and tempered condition. Room temperature microhardness measurements were carried out using a Leica Microhardness Tester using a diamond Vickers indenter and a load of 1 N. An average of five indentations was made on each sample and their average value was noted. The hardness (H) was calculated from 2
H ¼ 1854:4 L=d
ð1Þ
where L is load in g and d is diagonal length of the indentation in lm. Hot hardness measurements were carried out in a hot hardness tester (Nikon, Model QM) using a diamond Vickers pyramid indenter. The load was applied at a rate of 0.2 mm min1. The instrument was calibrated using a standard (Cu: SRM; National Bureau of Standards, USA) sample. Five indentations were made on the standard using 1 N load. The hardness obtained was found to be ±0.5% of the actual value. The metallographically polished sample was then loaded into the specimen holder. Care was taken to maintain the sample surface perpendicular to the microscope axis. The load used for hot hardness experiment was 1 N. Hot hardness measurements of T91 were carried out in vacuum (5 104 torr) from room temperature to 1023 K at every 100 K intervals and using a dwell time of 5 s. For U–15%Pu alloy, the hardness was measured from room temperature to 673 K in every 100 K intervals. Above this temperature, hardness measurements of U–15%Pu alloy were carried out in shorter intervals. The sample temperature was kept constant within ±1 K and that of the indenter was kept within ±3 K. 3. Results The variation of hardness with temperature is plotted in Fig. 1 on a semi-logarithmic scale for the T91 alloy. It is clearly brought out that the curves consist of two straight lines with different slopes intersecting at 778 K. The hardness values fall drastically above the transition temperature. Fig. 2 shows the hardness versus temperature plot for U–15%Pu alloy. The plot shows that the hardness falls marginally up to about 408 K and thereafter on further heating hardness falls sharply. A discontinuity was observed in the hardness–temperature plot at around 823 K. In the temperature range of 828–923 K, hardness falls marginally with temperature. At 988 K, hardness falls steeply reaching a value as low as 0.43 kg mm2. At 988 K, hardness value was so low; further measurements could not be carried out above this temperature. 4. Discussion The temperature dependence of hardness of metals and alloys has been reviewed by Westbrook [13]. The summary of his finding is that the temperature dependence of hardness is best represented by the relation of the type
H ¼ A expðBTÞ
ð2Þ
The constants A and B have one set of values at low temperatures (AI and BI) and another set of values at high temperatures (AII and BII), indicating two different deformation mechanisms. The values of A, B and transition temperatures for T91 and U–15%Pu alloy are given in Table 2.
Table 1 Chemical composition of T91 steel (wt.%). Cr 9.161
Mo 0.882
V 0.207
Nb 0.079
Al 0.008
Ti 0.003
Ni 0.197
Cu 0.068
Mn 0.368
Si 0.209
C 0.099
N 0.0457
P 0.015
S 0.0013
Fe Balance
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constant B, which is the slope of ln H versus T plot, indicates the rate at which the lattice is weakened by the thermal energy and is closely related to the coefficient of thermal expansion [16]. The constant B is known as softening coefficient or thermal coefficient of hardness. The low melting temperature of pure plutonium, its large number of phase transformations, its anisotropy in the low temperature phases, and the large, 10% change in volume during the beta to alpha transformation all work towards making pure plutonium an undesirable engineering material [18]. Although U–Pu alloy system forms the basis for fast reactor metallic fuels containing plutonium, few physical property studies have been made of the binary system alone. This is because of the extremely poor properties of samples containing even small amounts of the zeta phase. Understanding the U–Pu binary alloy phase diagram is important in investigating metallic fuel behaviour. The phase transformation behaviour of U–15%Pu binary alloy can be best evaluated with the help of U–Pu binary phase diagram [19]. There is complete mutual solubility of e-Pu and c-U over the entire range of composition, whereas U has only a limited solubility in the remaining Pu allotropes. a-U dissolves a maximum of 15 at.% Pu, whereas b-U dissolves about 20 at.% Pu. According to the U–Pu phase diagram, the addition of Pu to U lowers the allotropic transformation temperature. U and Pu form two intermediate phases: g, which exist at high temperatures, and f, which stable at room temperature. The g phase is formed by a peritectoid reaction (ePu, cU) + (bU) M g at 70 at.% U and 978 K. The g phase decomposes eutectoidally, g M (bPu) + f, at about 3 at.% U and 551 K. The f phase forms by a peritectoid reaction, g + (bU) M f at about 72 at.% and 863 K [19]. Crystal structure of g phase is simple tetragonal with lattice parameters 1.057 and 1.076 nm [19]. f phase has a cubic structure with lattice parameter varying with U content. Lattice parameter of f phase was found decreasing with increase in its U content at room temperature from 1.0692 for 35 at.% U to 1.0651 nm for the f phase containing 70 at.% U [19]. This study indicates that the formation of c phase results in a very low values of hardness which will help fuel to creep and reduce FCMI. Presence of b will lead to some sort of hardening of the fuel which will not harm the clad since clad is stronger than the fuel at these temperatures. It is reported that sample containing small amount of f phase shows very poor properties [18]. Hardness is the mechanical property of fuel that is measured most frequently. The room temperature of hardness of reactor grade U used in this study is 274 kg mm2. The effect of temperature on hardness of polycrystalline U has been studied by the present authors and has been found that at the transition temperature the hardness of b is greater than that of a [20]. Similar observation of higher hardness for b phase was also reported by Chubb et al. [21]. The hardness of as-cast a-Pu containing 300–1000 ppm total impurities is 260–285 kg mm2 [22–26]. The hardness for a-Pu at room temperature, having a density of 19.50–19.70 g cm3 and a total impurity content of less than 1000 ppm is reported as 265 kg mm2 [18]. When alpha transforms to beta, the hardness
300 200
Hardness (kg mm-2)
778 K 100 90 80 70 60 50 40 30 20
10
300
400
500
600
700
800
900
1000
1100
Temperature (K) Fig. 1. log (hardness) versus temperature plot for T91. The transition point is also shown in the figure.
1000
408 K
Hardness (kg mm-2)
100
10
α
1
β
γ
0.1 300
400
500
600
700
800
900
1000
Temperature (K) Fig. 2. log (hardness) versus temperature plot for U–15%Pu alloy showing the variation of hardness in different phases.
Westbrook [13] sought a correlation between constants A and B of Eq. (2) and some fundamental physical parameter. The constant A for the low temperature branch was termed as intrinsic hardness, i.e. extrapolated hardness at absolute zero. The value of A at low temperature is a measure of the inherent strength of the interatomic bond in the metal lattice and is related to the crystal structure and ‘thermal energy of melting’ which is defined as the heat content of the liquid metal at the melting point [13,15–17]. The
Table 2 Values of constants A and B in the relation H = A exp (BT). Sample
A (kg mm2)
B 103 (K1)
Transition temperature, TT (K)
AI
AII
BI
BII
T91
291
16,180
1.100
6.263
778
U–15%Pu (a phase)
326
3464
0.499
6.286
408
U–15%Pu (b phase)
1361
–
4.582
–
–
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decreases to about 60 kg mm2 at 398 K and decreases to about 25 kg mm2 at 453 K. The hardness of c-Pu is less than 20 kg mm2 and that of d-Pu is less than 10 kg mm2 [23]. A limited amount of hardness, tensile, and compressive data are available on Pu–U alloys. Hardness as-cast U–10%Pu has been reported in the literature as 205 DPH [18]. The hardness of rolled U rich U–Pu alloys has been reported in the literature [23]. The reported values of hardness are 300, 400, 300 DPH for U–10%Pu, U–15%Pu and U–20%Pu alloys, respectively. The hardness of the U–Pu alloys increases from 261 DPH at 5%Pu to 300 at 10%Pu, where the phase boundary between single phase a-U and the two phase (a-U + f) region exists. The room temperature hardness of as-cast U–15%Pu–6.8%Zr was reported to be 440 kg mm2 [14]. For the above alloy the hardness at 673 and 973 K was found to be 256 and 6.4 kg mm2, respectively. The above result shows that binary U–15%Pu alloy is softer than the ternary U–15%Pu–6.8%Zr alloy. As mentioned earlier, plutonium is soluble in concentrations up to 15% in a uranium, and the a-uranium phase retains its anisotropic characteristics. Alloys with up to 15% plutonium, like the present U–15%Pu alloy, can, therefore, be expected to have swelling behaviour similar to uranium [18,27]. In the a phase, cavitation dominates, whereas rapid growth of fission-gas bubbles dominates in the c phase. It appears that the addition of plutonium enhances the swelling rate in each of the phases, presumably because plutonium increases the diffusivity in the alloy and decreases the creep strength. Let us analyse the results of the present study, on the basis of the above observation. From the U–Pu phase diagram, the equilibrium phases for U–15%Pu are a-U and intermediate f-phase. The XRD analysis of U–15%Pu alloy showed the peaks correspond to a-phase of uranium, since the peritectoid reaction, g + (bU) M f is sluggish nature [19]. The average room temperature hardness of U–15%Pu alloy obtained in this study is 281 kg mm2. This value is lower than the only reported value of 400 kg mm2 for U–15%Pu alloy by Kittel et al. [18]. Since their alloy was rolled, no comparison could be made with present study. The a-phase extends up to 833 K. The hot hardness data of orthorhombic a phase consists of two straight-line with different slope intersecting at around 408 K. The sudden increase in the hardness above 828 K is due to appearance of the tetragonal b-phase. The b-phase extends up to 978 K and in this region fall in hardness with increase in temperature is meager. As mentioned earlier, the b-phase of U is reported to be harder than a-phase which is found to be true for this alloy also [21]. Above 978 K, the alloy enters c phase region which was found to very soft as can be seen from the hot hardness data. In the literature, a high temperature tensile strength value of U–10%Pu has been reported. Tensile strength of U–10%Pu decreased from 26 kg mm2 at 473 K to 0.5 kg mm2 at 1073 K [18]. No attempt was made here to compare this with present data since information on heat treatment, phase contents and impurity contents of the alloy are not given. A detailed study of the microstructure of the irradiated U–Pu–Zr metallic alloy reported three distinct concentric zones [28]. Similar behaviour is expected in U–15%Pu binary fuel. At the beginning of irradiation, this fuel will be extremely soft at the central zone due to the presence of extremely soft c-phase as indicated by the hot hardness data of this study. The intermediate zone will be harder due to presence of b dominated phases. The outer peripheral zone will be comparatively softer due to presence of a phase. As a result of a higher swelling rate and fluid like behaviour in the hotter central part of the fuel, the fuel in this region is in a hydrostatic stress state, while the cooler and stronger outer shell is in a state of tensile stress [10]. The effect of these tensile stresses on swelling and creep in the outer fuel zone explains the larger radial than axial swelling and, thus, observed anisotropy in swelling [7]. The outer, anisotropic swelling leads to large radial cracks, presumably
because there is insufficient time to relax radial stresses by means of creep. It is reported that the fuel and clad get ‘locked’ axially in the upper section of the fuel [10]. Across the radius of the fuel, the central region is in c phase and rather compliant. The middle and outer region are colder and stiffer. Also, the fuel–cladding contact and thermal gradients would cause the fuel slug to ratchet up the fuel column, further exacerbated by fuel creep during irradiation that would prevent the fuel from dropping to the original position during shutdown. Furthermore, during transients, one has to guard against fuel–clad interaction, which may result in the damage of the clad [10,29]. Fuel–cladding mechanical interaction (FCMI) arises from applied stress when the cladding is designed to restrain fuel swelling, and may result in plastic deformation of the cladding. From the hardness data of the cladding and fuel, it may be able to predict whether fuel can damage the cladding due to its interaction. To check the interaction between fuel and cladding, we plotted the hardness of T91 and U–15%Pu together and are shown in Fig. 3. The figure shows the hardness of fuel is lower than the clad at temperatures above 473 K. From Fig. 3, it is clear that the hardness in the b-phase region comes closer to the cladding and diverges on entering into c region. This confirms at the reactor operating temperatures, U–15%Pu fuel is softer than the T91 cladding and hence if any mechanical interaction occurs between fuel and cladding, the probability of the clad damage by the fuel is very much less since the fuel is softer than the cladding. Because of early pin designs did not allow the fuel sufficient room for free swelling, gas-bubble pressure was transmitted directly to the cladding [30]. As a result, all early designs suffered from cladding deformation and rupture at modest burnups. However, as discussed earlier, a smeared density of about 75% allows free fuel swelling of approximately 33%, at which point porosity becomes largely interconnected and open to the outside of the fuel, releasing a large fraction of the fission gas to a suitably large plenum at the top of the pin. The gas pressure in the open pores is then determined by the volume and temperature of the plenum above the fuel. The real source of the stress during steady state and transients is fuel clad mechanical interaction due to fuel swelling and fuel thermal expansion and the plenum pressure. The main reason for the return of FCMI comes from the swelling caused by solid fission products. Solid fission products are not released like the fission gas so the production of them creates a stress on the cladding. It also closes off some of the open porosity and the gas swelling will
300
T91 U-15 wt.% Pu Fitted (T91) Fitted (U-15 wt.% Pu)
250
Hardness (kg mm-2)
344
200
150
100
50
0 300
400
500
600
700
800
900
1000
1100
Temperature (K) Fig. 3. Hardness versus temperature plots for T91 and U–15%Pu together showing U–15%Pu fuel is softer than the T91 cladding above 473 K.
T.R.G. Kutty et al. / Journal of Nuclear Materials 429 (2012) 341–345
enhance FCMI. The amount of open porosity decreases with the amount of solid fission products. The fact that the fuel may also stiffens during this process [31]. If the magnitude of the fast neutron flux is high or the neutron spectrum is excessively hard, the dose to the clad may induce void swelling at high burnup, resulting in further embrittlement of the clad. This study does not take into account of irradiation effects on fuel pin. It has shown that the U–15%Pu fuel cannot deform the T91 cladding at the reactor operating temperatures since fuel is much softer than the cladding.
5. Conclusions
[3] [4] [5] [6] [7]
[8] [9] [10] [11] [12] [13] [14] [15]
The high temperature hardness behaviour of U–15%Pu and T91 cladding was studied with the help of a hot hardness tester and the following conclusions are drawn: 1. The hardness versus temperature plot for the T91 alloy consists of two straight lines with different slope intersecting at 778 K. 2. The hot hardness data of orthorhombic a phase of U–15%Pu showed a transition at around 408 K. 3. The appearance of the tetragonal b-phase showed a sudden increase in the hardness above 828 K. 4. In the b-phase, which extends up to 978 K, fall in hardness with increase in temperature is meager. 5. c phase region of U–15%Pu was found to be very soft. 6. At the reactor operating temperatures, U–15%Pu fuel is softer than the T91 cladding and hence if any mechanical interaction occurs between fuel and cladding, the probability of the clad damage by the fuel is very much less.
[16] [17] [18]
[19] [20] [21]
[22]
[23] [24] [25] [26] [27] [28]
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Effect of temperature on hardness of binary U–15%Pu alloy and T91 cladding T.R.G. Kutty ⇑, K. Ravi, Santu Kaity, S.K. Swarnkar, Arun Kumar Radiometallurgy Division, Bhabha Atomic Research Centre, Mumbai 400 085, India
h i g h l i g h t s " Hardness–temperature plot of T91 alloy showed a transition at 778 K. " Hardness–temperature plot of
a phase of U–15%Pu showed a transition at 408 K.
" b Phase of U–15%Pu alloy showed a discontinuity in hardness data above 828 K. "
c Phase region of U–15%Pu was found to be very soft.
" At reactor operating temperatures, U–15%Pu fuel is softer than T91 cladding.
a r t i c l e
i n f o
Article history: Received 18 April 2012 Accepted 15 June 2012 Available online 23 June 2012
a b s t r a c t The high temperature hardness behaviour of U–15%Pu and T91 cladding was studied with the help of a hot hardness tester. The hardness versus temperature plot for the T91 alloy consists of two straight lines with different slopes intersecting at 778 K. The hot hardness data of orthorhombic a phase of U–15%Pu showed a transition at around 408 K. The appearance of the tetragonal b-phase showed a sudden increase in the hardness above 828 K. c-Phase of U–15%Pu was found to very soft. Since U–15%Pu fuel is softer than the T91 cladding above 473 K, the probability of the clad damage by the fuel due to its interaction with the cladding is very less. Ó 2012 Elsevier B.V. All rights reserved.
1. Introduction For the rapid growth of fast reactor programme in India, it is essential to shift to the use of metallic fuels in fast breeder reactors (FBR), which gives a higher breeding ratio (BR) and lower doubling time (DT) [1,2]. In FBR, the basic thermal and neutronic performance of metallic fuels is better than oxide ceramic fuels. The harder spectrum in the metallic fuelled core results in high breeding ratio [3,4]. India is going ahead with two types of fast reactor fuel element designs, which are under active consideration, viz. (i) sodium bonded U–15%Pu–6%Zr alloy and (ii) mechanically bonded binary U–15%Pu binary alloy [1,2]. For the mechanically bonded binary alloy, a Zr liner is proposed between the fuel pellet and clad to reduce the fuel–clad interaction. Modified T91, which is a 9Cr1Mo ferritic martensitic steel, has been used as the cladding material. Smear density, which is a key parameter for accommodation of fuel swelling, varies between 80% and 85% for mechanical bonded fuel and 75% for sodium bonded fuel. Since achieving a higher breeding ratio is one of the goals of India’s fast reactor programme, and breeding ratio increases linearly with the reduction of zirconium content in the fuel, a ternary fuel with a lower Zr ⇑ Corresponding author. Tel.: +91 22 2559 5361; fax: +91 22 2550 5151. E-mail Kutty).
addresses:
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0022-3115/$ - see front matter Ó 2012 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.jnucmat.2012.06.025
(T.R.G.
content is the preferred alloy [1]. In this study, emphasis has been given to study the mechanical properties of binary U–15%Pu alloy. It is well known that metallic fuel swells when irradiated due to the accumulation of fission products, especially the inert gaseous fission products. When total fuel swelling reaches the allowable limit due to gap closure, the contact stress between the fuel and the cladding starts to increase with the burnup [5–7]. During this period, swelling by closed bubbles decreases due to increase in external pressure on the closed bubble. The volume fraction of porosity remains relatively the same, until they are filled by solid fission products. It should be mentioned that the build up of the solid fission products is a very important part of the fuel swelling. Fission gas swelling saturates early on irradiation, whereas solid fission products continue to building up uniformly due to fission event causing volume expansion. As a result, fuel/clad contact pressure increases causing creep induced hot pressing of the fuel and open porosity decrease of the fuel [5,6]. It may be noted that there should be a dynamic equilibrium, where open pores are filled, new ones are formed and closed pores connect each other. Operation with an incompressible fuel may cause a very significant fuel clad mechanical interaction (FCMI). Hence one need to adjust the fuel smear density carefully so that fuel will be rather compressible by the end of the life [7]. FCMI is a significant issue during normal operation as well as in case of transient overpower accidents. Interaction between fuel and cladding that may lead to
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cladding failure can be both mechanical and chemical. Fuel–clad contact stresses can be minimised by maintaining a low (<75%) smear density, which guarantees a spongy fuel microstructure with open porosity [3]. In fast reactor, the metallic fuel forms various phases under the influence of the radial temperature gradient [8,9]. Among them, single BCC c-phase, which dominates at the central parts of the fuel, has a high creep rate; therefore, rendering the fuel rather compliant. On the other hand, at outer periphery of the fuel pin, the stiff a phase may dominate, reducing fuel compressibility and delaying bubble growth, coalescence and formation of interconnected porosity [10]. In this study, an attempt has been made to evaluate the strength of the U–15%Pu fuel and T91 cladding at various temperatures by measuring their hardness. In general, hardness of both the fuel and cladding are required to fully describe the gap conductance and contact resistance, once the mechanical contact between the fuel and cladding has occurred [11,12]. Hardness is an easily measured mechanical property that can provide insight into the strength of the fuel and cladding. Since hardness is related to yield stress, UTS, stress exponent and other mechanical properties by simple empirical relations, prediction of these mechanical properties from hardness data will be very useful for radiotoxic materials where conventional testing is very difficult to perform. The hot hardness data further give information on softening behaviour [13] and can be correlated to the Larson–Miller and Sherby–Dorn parameters to predict the long term rupture and creep properties. The main objectives of the present study are: i. Determination of the hot-hardness data of the U–15%Pu alloy from room temperature to 1000 K. ii. Evaluation of hot hardness of T91 cladding from room temperature to 1073 K. iii. From these data, an attempt will be made to assess fuel–clad mechanical interaction between U–15%Pu fuel and T91 cladding. The hardness values for the as-cast ternary U–15%Pu–6.8%Zr (composition in wt.%) was measured by Harbur et al. [14] between 673 and 1073 K using a load of 200 g. However, no studies have been reported on the mechanical properties of binary U–15%Pu alloy. Hence this study will be very useful for the fuel designers. 2. Experimental procedure U–15%Pu alloys were made by melting and casting route. The U and Pu slugs were melted together using induction melting in an yttria coated graphite crucible. The metallic impurities of the alloy were revealed by atomic emission spectroscopy (AES) using DC-arc as an excitation source. The phase analysis of the fuel alloy was carried out using XRD. For this, Cu Ka radiation was used with fixed slit optics and h–h goniometer. Diffraction patterns were obtained by the step-scan method with a step size of 0.01° and a measuring time of 1 s at each step were used. For metallographic analyses, the alloy slugs were cut into disc of suitable thickness using slow speed SiC abrasive cut-off wheel. Standard metallographic techniques were adopted for grinding and polishing using cold setting resin mounts. The chemical composition of T91 used in this study is shown in Table 1. The heat treatment for cladding material, T91, consists of austenization at 1323 K and air quench-
ing followed by tempering at 1023 K for 1 h. Grain size of T91 was found to be around 30 lm and its hardness measured using Vicker’s hardness tester at a test load of 200 g was 220 kg mm2. T91 steels are used in the normalised and tempered condition. Room temperature microhardness measurements were carried out using a Leica Microhardness Tester using a diamond Vickers indenter and a load of 1 N. An average of five indentations was made on each sample and their average value was noted. The hardness (H) was calculated from 2
H ¼ 1854:4 L=d
ð1Þ
where L is load in g and d is diagonal length of the indentation in lm. Hot hardness measurements were carried out in a hot hardness tester (Nikon, Model QM) using a diamond Vickers pyramid indenter. The load was applied at a rate of 0.2 mm min1. The instrument was calibrated using a standard (Cu: SRM; National Bureau of Standards, USA) sample. Five indentations were made on the standard using 1 N load. The hardness obtained was found to be ±0.5% of the actual value. The metallographically polished sample was then loaded into the specimen holder. Care was taken to maintain the sample surface perpendicular to the microscope axis. The load used for hot hardness experiment was 1 N. Hot hardness measurements of T91 were carried out in vacuum (5 104 torr) from room temperature to 1023 K at every 100 K intervals and using a dwell time of 5 s. For U–15%Pu alloy, the hardness was measured from room temperature to 673 K in every 100 K intervals. Above this temperature, hardness measurements of U–15%Pu alloy were carried out in shorter intervals. The sample temperature was kept constant within ±1 K and that of the indenter was kept within ±3 K. 3. Results The variation of hardness with temperature is plotted in Fig. 1 on a semi-logarithmic scale for the T91 alloy. It is clearly brought out that the curves consist of two straight lines with different slopes intersecting at 778 K. The hardness values fall drastically above the transition temperature. Fig. 2 shows the hardness versus temperature plot for U–15%Pu alloy. The plot shows that the hardness falls marginally up to about 408 K and thereafter on further heating hardness falls sharply. A discontinuity was observed in the hardness–temperature plot at around 823 K. In the temperature range of 828–923 K, hardness falls marginally with temperature. At 988 K, hardness falls steeply reaching a value as low as 0.43 kg mm2. At 988 K, hardness value was so low; further measurements could not be carried out above this temperature. 4. Discussion The temperature dependence of hardness of metals and alloys has been reviewed by Westbrook [13]. The summary of his finding is that the temperature dependence of hardness is best represented by the relation of the type
H ¼ A expðBTÞ
ð2Þ
The constants A and B have one set of values at low temperatures (AI and BI) and another set of values at high temperatures (AII and BII), indicating two different deformation mechanisms. The values of A, B and transition temperatures for T91 and U–15%Pu alloy are given in Table 2.
Table 1 Chemical composition of T91 steel (wt.%). Cr 9.161
Mo 0.882
V 0.207
Nb 0.079
Al 0.008
Ti 0.003
Ni 0.197
Cu 0.068
Mn 0.368
Si 0.209
C 0.099
N 0.0457
P 0.015
S 0.0013
Fe Balance
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constant B, which is the slope of ln H versus T plot, indicates the rate at which the lattice is weakened by the thermal energy and is closely related to the coefficient of thermal expansion [16]. The constant B is known as softening coefficient or thermal coefficient of hardness. The low melting temperature of pure plutonium, its large number of phase transformations, its anisotropy in the low temperature phases, and the large, 10% change in volume during the beta to alpha transformation all work towards making pure plutonium an undesirable engineering material [18]. Although U–Pu alloy system forms the basis for fast reactor metallic fuels containing plutonium, few physical property studies have been made of the binary system alone. This is because of the extremely poor properties of samples containing even small amounts of the zeta phase. Understanding the U–Pu binary alloy phase diagram is important in investigating metallic fuel behaviour. The phase transformation behaviour of U–15%Pu binary alloy can be best evaluated with the help of U–Pu binary phase diagram [19]. There is complete mutual solubility of e-Pu and c-U over the entire range of composition, whereas U has only a limited solubility in the remaining Pu allotropes. a-U dissolves a maximum of 15 at.% Pu, whereas b-U dissolves about 20 at.% Pu. According to the U–Pu phase diagram, the addition of Pu to U lowers the allotropic transformation temperature. U and Pu form two intermediate phases: g, which exist at high temperatures, and f, which stable at room temperature. The g phase is formed by a peritectoid reaction (ePu, cU) + (bU) M g at 70 at.% U and 978 K. The g phase decomposes eutectoidally, g M (bPu) + f, at about 3 at.% U and 551 K. The f phase forms by a peritectoid reaction, g + (bU) M f at about 72 at.% and 863 K [19]. Crystal structure of g phase is simple tetragonal with lattice parameters 1.057 and 1.076 nm [19]. f phase has a cubic structure with lattice parameter varying with U content. Lattice parameter of f phase was found decreasing with increase in its U content at room temperature from 1.0692 for 35 at.% U to 1.0651 nm for the f phase containing 70 at.% U [19]. This study indicates that the formation of c phase results in a very low values of hardness which will help fuel to creep and reduce FCMI. Presence of b will lead to some sort of hardening of the fuel which will not harm the clad since clad is stronger than the fuel at these temperatures. It is reported that sample containing small amount of f phase shows very poor properties [18]. Hardness is the mechanical property of fuel that is measured most frequently. The room temperature of hardness of reactor grade U used in this study is 274 kg mm2. The effect of temperature on hardness of polycrystalline U has been studied by the present authors and has been found that at the transition temperature the hardness of b is greater than that of a [20]. Similar observation of higher hardness for b phase was also reported by Chubb et al. [21]. The hardness of as-cast a-Pu containing 300–1000 ppm total impurities is 260–285 kg mm2 [22–26]. The hardness for a-Pu at room temperature, having a density of 19.50–19.70 g cm3 and a total impurity content of less than 1000 ppm is reported as 265 kg mm2 [18]. When alpha transforms to beta, the hardness
300 200
Hardness (kg mm-2)
778 K 100 90 80 70 60 50 40 30 20
10
300
400
500
600
700
800
900
1000
1100
Temperature (K) Fig. 1. log (hardness) versus temperature plot for T91. The transition point is also shown in the figure.
1000
408 K
Hardness (kg mm-2)
100
10
α
1
β
γ
0.1 300
400
500
600
700
800
900
1000
Temperature (K) Fig. 2. log (hardness) versus temperature plot for U–15%Pu alloy showing the variation of hardness in different phases.
Westbrook [13] sought a correlation between constants A and B of Eq. (2) and some fundamental physical parameter. The constant A for the low temperature branch was termed as intrinsic hardness, i.e. extrapolated hardness at absolute zero. The value of A at low temperature is a measure of the inherent strength of the interatomic bond in the metal lattice and is related to the crystal structure and ‘thermal energy of melting’ which is defined as the heat content of the liquid metal at the melting point [13,15–17]. The
Table 2 Values of constants A and B in the relation H = A exp (BT). Sample
A (kg mm2)
B 103 (K1)
Transition temperature, TT (K)
AI
AII
BI
BII
T91
291
16,180
1.100
6.263
778
U–15%Pu (a phase)
326
3464
0.499
6.286
408
U–15%Pu (b phase)
1361
–
4.582
–
–
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decreases to about 60 kg mm2 at 398 K and decreases to about 25 kg mm2 at 453 K. The hardness of c-Pu is less than 20 kg mm2 and that of d-Pu is less than 10 kg mm2 [23]. A limited amount of hardness, tensile, and compressive data are available on Pu–U alloys. Hardness as-cast U–10%Pu has been reported in the literature as 205 DPH [18]. The hardness of rolled U rich U–Pu alloys has been reported in the literature [23]. The reported values of hardness are 300, 400, 300 DPH for U–10%Pu, U–15%Pu and U–20%Pu alloys, respectively. The hardness of the U–Pu alloys increases from 261 DPH at 5%Pu to 300 at 10%Pu, where the phase boundary between single phase a-U and the two phase (a-U + f) region exists. The room temperature hardness of as-cast U–15%Pu–6.8%Zr was reported to be 440 kg mm2 [14]. For the above alloy the hardness at 673 and 973 K was found to be 256 and 6.4 kg mm2, respectively. The above result shows that binary U–15%Pu alloy is softer than the ternary U–15%Pu–6.8%Zr alloy. As mentioned earlier, plutonium is soluble in concentrations up to 15% in a uranium, and the a-uranium phase retains its anisotropic characteristics. Alloys with up to 15% plutonium, like the present U–15%Pu alloy, can, therefore, be expected to have swelling behaviour similar to uranium [18,27]. In the a phase, cavitation dominates, whereas rapid growth of fission-gas bubbles dominates in the c phase. It appears that the addition of plutonium enhances the swelling rate in each of the phases, presumably because plutonium increases the diffusivity in the alloy and decreases the creep strength. Let us analyse the results of the present study, on the basis of the above observation. From the U–Pu phase diagram, the equilibrium phases for U–15%Pu are a-U and intermediate f-phase. The XRD analysis of U–15%Pu alloy showed the peaks correspond to a-phase of uranium, since the peritectoid reaction, g + (bU) M f is sluggish nature [19]. The average room temperature hardness of U–15%Pu alloy obtained in this study is 281 kg mm2. This value is lower than the only reported value of 400 kg mm2 for U–15%Pu alloy by Kittel et al. [18]. Since their alloy was rolled, no comparison could be made with present study. The a-phase extends up to 833 K. The hot hardness data of orthorhombic a phase consists of two straight-line with different slope intersecting at around 408 K. The sudden increase in the hardness above 828 K is due to appearance of the tetragonal b-phase. The b-phase extends up to 978 K and in this region fall in hardness with increase in temperature is meager. As mentioned earlier, the b-phase of U is reported to be harder than a-phase which is found to be true for this alloy also [21]. Above 978 K, the alloy enters c phase region which was found to very soft as can be seen from the hot hardness data. In the literature, a high temperature tensile strength value of U–10%Pu has been reported. Tensile strength of U–10%Pu decreased from 26 kg mm2 at 473 K to 0.5 kg mm2 at 1073 K [18]. No attempt was made here to compare this with present data since information on heat treatment, phase contents and impurity contents of the alloy are not given. A detailed study of the microstructure of the irradiated U–Pu–Zr metallic alloy reported three distinct concentric zones [28]. Similar behaviour is expected in U–15%Pu binary fuel. At the beginning of irradiation, this fuel will be extremely soft at the central zone due to the presence of extremely soft c-phase as indicated by the hot hardness data of this study. The intermediate zone will be harder due to presence of b dominated phases. The outer peripheral zone will be comparatively softer due to presence of a phase. As a result of a higher swelling rate and fluid like behaviour in the hotter central part of the fuel, the fuel in this region is in a hydrostatic stress state, while the cooler and stronger outer shell is in a state of tensile stress [10]. The effect of these tensile stresses on swelling and creep in the outer fuel zone explains the larger radial than axial swelling and, thus, observed anisotropy in swelling [7]. The outer, anisotropic swelling leads to large radial cracks, presumably
because there is insufficient time to relax radial stresses by means of creep. It is reported that the fuel and clad get ‘locked’ axially in the upper section of the fuel [10]. Across the radius of the fuel, the central region is in c phase and rather compliant. The middle and outer region are colder and stiffer. Also, the fuel–cladding contact and thermal gradients would cause the fuel slug to ratchet up the fuel column, further exacerbated by fuel creep during irradiation that would prevent the fuel from dropping to the original position during shutdown. Furthermore, during transients, one has to guard against fuel–clad interaction, which may result in the damage of the clad [10,29]. Fuel–cladding mechanical interaction (FCMI) arises from applied stress when the cladding is designed to restrain fuel swelling, and may result in plastic deformation of the cladding. From the hardness data of the cladding and fuel, it may be able to predict whether fuel can damage the cladding due to its interaction. To check the interaction between fuel and cladding, we plotted the hardness of T91 and U–15%Pu together and are shown in Fig. 3. The figure shows the hardness of fuel is lower than the clad at temperatures above 473 K. From Fig. 3, it is clear that the hardness in the b-phase region comes closer to the cladding and diverges on entering into c region. This confirms at the reactor operating temperatures, U–15%Pu fuel is softer than the T91 cladding and hence if any mechanical interaction occurs between fuel and cladding, the probability of the clad damage by the fuel is very much less since the fuel is softer than the cladding. Because of early pin designs did not allow the fuel sufficient room for free swelling, gas-bubble pressure was transmitted directly to the cladding [30]. As a result, all early designs suffered from cladding deformation and rupture at modest burnups. However, as discussed earlier, a smeared density of about 75% allows free fuel swelling of approximately 33%, at which point porosity becomes largely interconnected and open to the outside of the fuel, releasing a large fraction of the fission gas to a suitably large plenum at the top of the pin. The gas pressure in the open pores is then determined by the volume and temperature of the plenum above the fuel. The real source of the stress during steady state and transients is fuel clad mechanical interaction due to fuel swelling and fuel thermal expansion and the plenum pressure. The main reason for the return of FCMI comes from the swelling caused by solid fission products. Solid fission products are not released like the fission gas so the production of them creates a stress on the cladding. It also closes off some of the open porosity and the gas swelling will
300
T91 U-15 wt.% Pu Fitted (T91) Fitted (U-15 wt.% Pu)
250
Hardness (kg mm-2)
344
200
150
100
50
0 300
400
500
600
700
800
900
1000
1100
Temperature (K) Fig. 3. Hardness versus temperature plots for T91 and U–15%Pu together showing U–15%Pu fuel is softer than the T91 cladding above 473 K.
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enhance FCMI. The amount of open porosity decreases with the amount of solid fission products. The fact that the fuel may also stiffens during this process [31]. If the magnitude of the fast neutron flux is high or the neutron spectrum is excessively hard, the dose to the clad may induce void swelling at high burnup, resulting in further embrittlement of the clad. This study does not take into account of irradiation effects on fuel pin. It has shown that the U–15%Pu fuel cannot deform the T91 cladding at the reactor operating temperatures since fuel is much softer than the cladding.
5. Conclusions
[3] [4] [5] [6] [7]
[8] [9] [10] [11] [12] [13] [14] [15]
The high temperature hardness behaviour of U–15%Pu and T91 cladding was studied with the help of a hot hardness tester and the following conclusions are drawn: 1. The hardness versus temperature plot for the T91 alloy consists of two straight lines with different slope intersecting at 778 K. 2. The hot hardness data of orthorhombic a phase of U–15%Pu showed a transition at around 408 K. 3. The appearance of the tetragonal b-phase showed a sudden increase in the hardness above 828 K. 4. In the b-phase, which extends up to 978 K, fall in hardness with increase in temperature is meager. 5. c phase region of U–15%Pu was found to be very soft. 6. At the reactor operating temperatures, U–15%Pu fuel is softer than the T91 cladding and hence if any mechanical interaction occurs between fuel and cladding, the probability of the clad damage by the fuel is very much less.
[16] [17] [18]
[19] [20] [21]
[22]
[23] [24] [25] [26] [27] [28]
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