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Changes in mechanical properties of irradiated Zircaloy-2 fuel cladding due to short term annealing
ELSEVIER
JoillTlalof NuclearMaterials 238 (1996) 169-174
Changes in mechanical properties of irradiated Zircaloy-2 fuel cladding due to short term annealing Tadahiko Torimaru *, Takayoshi Yasuda
1, Masafumi Nakatsuka
Nippon Nuclear Fuel Development Co., Ltd, 2163, Narita-cho, Oarai-machi, Higashi-ibarald-gun, Ibaraki-ken, 311-13, Japan Received 13 February 1995; accepted 19 July 1996
Abstract
Zirconium-lined fuel cladding tubes irradiated to 2.7 x 1 0 25 n / m 2 (E > 1 MeV) in a BWR, which had experienced recrystallized annealing in the final process in their manufacture, were heat treated at 500-700°C for 5-600 s to simulate short term dry-out. Tensile tests, hardness measurements, fatigue tests and X-ray analyses were made on those specimens. The irradiation hardening in hardness at room temperature and ultimate tensile strength at 343*(2 recovered to approximately 80% of that after heat treatment at 600-700°C for less than 15 s. Fatigue life and half value width of X-ray analysis recovered to these of unirradiated cladding tube after annealing for 15 s at 600°C. These recovery rates were faster than those on cold worked and stress relieved zirconium alloys. An equation to predict the remaining fraction of hardening was proposed by using the regression analysis on tensile strength and hardness values.
1. Introduction
Zr liner fuel claddings, which are lined with a zirconium barrier on the inner surface of the base cladding of Zircaloy-2 and heat treated under the recrystallization condition, are mainly used in BWRs. When a drop in cooling ability occurs for a short time in a BWR, the temperature of the fuel rods increases by a boiling transition or dry-out. The rise in temperature changes the mechanical properties of the cladding that has already been hardened by neutron induced damage. Therefore, detailed data on the mechanical properties of irradiated Zr liner fuel cladding tubes after short term annealing are indispensable for evaluation of fuel performance. However, earlier studies have only dealt with recovery and recrystallization behavior of coldworked material [1-3] and with recovery behavior of irradiated Zircaloy-4 in the stress relief condition during the final manufacturing process [4].
* Corresponding author. Tel.: +81-29 267 4141; fax: +81-29 267 7135; e-mail: [email protected]. l Present address: Hitachi Nuclear EngineeringCo., Ltd.
The purposes of the present study are to clarify the recovery behavior of irradiated Zr liner fuel cladding from several mechanical properties and half value widths of X-ray diffraction lines after short term annealing, and to formulate the fraction of recovery as a function of annealing temperature and time.
2. Experimental details 2.1. Test material Test materials were retrieved from irradiated Zr liner fuel cladding for which fluence and mean burn up had amounted to 2.7X 1025 n / m 2 ( E > 1 MeV) and 18 G W d / t in a commercial BWR. This cladding had a hydrogen content of 50 ppm and average oxide layer thickness of 28 ~m on the outer surface. Ring specimens at a point approximately 2500 mm from the rod lower end of the cladding were cut to 5 mm in length, and fuel pellets were removed. Dimensions, irradiation conditions and main chemical composition of the cladding are summarized in Tables 1 and 2.
0022-3115/96/$15.00 Copyright © 1996 Elsevier Science B.V. All rights reserved. PII S0022-3115(96)00451-5
170
T. Torimaru et al. / Journal of Nuclear Materials 238 (1996) 169-174
Table 1 Dimensions and irradiation conditions of fuel rods Outer and inner diameters (mm)
Material
Thickness of liner (Ixm)
Irradiation temperature (K)
Neutron fluence (n/m e) (E > 1 MeV)
12.27 10.55
Zircaloy-2
~ 80
~ 620
2.6 to 2.8 X 1025
2.2. Heat treatment
A radiation heating furnace with infrared lamps installed as a heat source was employed for short time isothermal annealing to simulate the temperature changes during dry-out. Since mechanical properties of irradiated cladding were thought to change quickly over 500°C, it was desirable for the heating and cooling rates to be as fast as possible. The furnace could provide rapid heating from 500°C to 700°C at a rate of 4 0 ° C / s and rapid cooling from 700°C to 5000C at 30°C/s. Heat treatment was conducted in argon gas to prevent specimen oxidation. A thermocouple was spot welded on the cross-section of each ring specimen, and their temperatures during heat treatment were measured. The deviation from the desired temperature was held to within + 5°C.
to the loading rods at two positions which were 180 degrees apart from each other. Fully reversed bending strains were induced with 1 Hz in the specimen by the reciprocal movement of the loading rods. The tests were conducted at 350°C in argon gas. The time when the load dropped to approximately 60% of the initial load was judged to represent fatigue failure, and the number of cycles then was recorded. The details of the fatigue tester and test method have been described elsewhere [5]. The tube cross-sections of the specimens used for hardness tests were polished once again for X-ray diffraction. Diffraction intensity curves were obtained by using an X-ray diffractmeter with a Cu K s ray. Other operating conditions were: tube voltage, 40 kV; tube current, 30 mA; beam diameter, 1.0 mm; detector scan speed, 0.5°C/rain. The half value width for each diffraction plane was measured from the diffraction curves.
2.3. Experimental procedure
The ring shaped specimens had a 5 mm length to allow the plane stress condition to be established and they had a cross-section larger than 5 mm. The specimens were tensile tested using two semicircular tension jigs. The ultimate tensile strength (UTS), the distance of cross-head movement up to failure and the reductions of area were measured. The tests were conducted at 343°C in air with the strain rate of approximately 0.5%/min. The gauge length of a ring specimen was assumed to be one half of the tube diameter. The ring shaped specimens had a 5 mm length were fixed into resin and polished their cross-section. Vickers hardness on the cross-section of each specimen at room temperature was measured with force of 200 g (1.96 N) for Zircaloy-2 and 50 g (0.49 N) for zirconium liner. Cladding tubes, machined into C-shape specimens, were adopted for the bending fatigue test. A specimen was fixed
3. Experimental results Fig. l(a)-(c) show ultimate tensile strength, cross-head movement at failure and reduction of area against the holding time of annealing, respectively. The tensile strength after annealings at 700°C × 5 s and 600-650°C x 15 s, almost recovered to that of unirradiated specimen. Those after the annealing at 500-575°C tended to decrease with annealing time and show the maximum value at an holding time. These maximum values were considered to be due to radiation anneal hardening (RAH). This issue will be discussed later. Vickers hardness changes of the irradiated and unirradiated specimens for Zircaloy-2 and zirconium liner are depicted in Fig. 2. The closed circle and bar show the average value and standard deviation of unirradiated specimens with and without heat treatment. The hardness of the
Table 2 Chemical composition of fuel cladding (wt% or wtppm) Zircaloy-2 a Zr-liner b
Sn
Fe
Cr
Ni
O
1.42% < 10 ppm
0.14% 430 ppm
0.10% 54 ppm
0.05% < 35 ppm
1150 ppm 350 ppm
Chemical composition satisfied ASTM B353, recrystallized before irradiation. b High purity sponge zirconium.
171
T. Torimaru et al. / Journal of Nuclear Materials 238 (1996) 169-174 14" 500"C -4-550"C ~ 575"C-o- 600"C-0--650"C~ ?O0"C]
Holding Time 15s
~ 600
100
....
,
ao
>~
400
60
/
40
/
~i
°i--~
(b)
1.5
o ~
1,0 0,5
As-irradiated
80 ~
.
(c)
40
0
~
................. I0
i |
i|l
I00
I000
HOLDING TIME (s) Fig. I. Changes in the mechanical properties with holding time of annealing in tensile tests at 343°C (616 K).
["e-500"C
260
Unlrradlated
//
~ , I (211) (201) (110) I
/
i .
i
.
i
5oo,~o ~
°
,
i
*1
° ,
i
i
Mo 7m
Fig. 3. Recovery of half-value width with annealing temperature.
2.0 25
0 ~
-/--.
/
i~"
/o, i
I
......
,
~-Irmdlated//. / '
20
3.0
•
4,-.
/*o,. ,
~ 300
**
~ .....
"P'$$0"C
-~'-575"C
"'o--800"(2
"-o-650"C
.,~e-n'z'adiated
.-A.-?O0"C
I
Zry-2
unirradiated specimens showed no great changes, and the effect of thermal shock during annealing could be ignored. For irradiated specimens, the hardness after annealing above 600 for 15 s or more almost recovered to that for the unirradiated specimen. The recovery curves of hardness after the annealing at 500-550°C also showed RAH. Number of cycles to fatigue failures under the total bending strain amplitude of 0.55% were measured. The failure cycles of unirradiated and as-irradiated specimens were 2700 and 20350, respectively. The failure cycle after heat treatment at 600°C for 15 s nearly recovered to that for unirradiated specimen. Fig. 3 plots the recovery in half value width of X-ray diffraction lines against the annealing temperatures for the three diffraction planes. The half value widths after annealing above 600°C nearly recovered to that for unirradiated specimen. These results corresponded to those of the mechanical tests. Fig. 4 shows the half value widths for the (211) plane plotted against the Vickers hardness of the Zircaloy-2. Half value widths and Vickers hardness had a linear relationship. This result meant that the recovery of hardness was due to disappearance of irradiation defects.
9-20 ~ N m
200
Uni~
180 160 140
AB-irradi~ted
Zr-liner
Holding tlme 15s (Temperature °C)
~ '~
0.8
>~
o.s
As-lrradlate~
Unlrradlated
120 I00 80 60
o.,
Unirradiated ~ .......... 10
100
1000
HOLDING TIME (s) Fig. 2. Changes in Vickers hardness at room temperature with holding time of annealing.
0.2
160
i
.
i
.
i
.
I
180 200 220 240 VICKERS HARDNESS (Hv)
i
260
Fig. 4. Relationship between half value width and Vickers hardness.
172
T. Torimaru et a l . / Journal of Nuclear Materials 238 (1996) 169-174 600
4. Discussion
I
I I IIIIII
I
i IIIIIII
( ~ )
4.1. Time dependence of softening As shown in Figs. 1 and 2, softening of irradiated specimen rapidly progresses during short term annealings, 5-15 s, and then slows down. Generally, many defects have been found to be uniformly formed in irradiated fuel cladding tubes. Small defects, which cannot be distinguished using TEM, were produced in Zircaloy-2 tubes irradiated at 573 K to a fast neutron fluence of less than 2-3 × 1023 n / m 2 [6]. Very small dislocation loops with an average diameter of approximately 37 nm were formed in cladding tubes irradiated at 50°C to the fluence of 5.5 × 1019 n/cm2[7]. From these reports, it is clear that defects made by neutron irradiation are very small. According to Dollins" recovery model [8], large defects are relatively stable in the initial process during post irradiation annealing, although small loops rapidly dissolve. This analysis explains the fast recovery behavior shown in Figs. 1 and 2. While both tensile strength and hardness drop monotonously with the holding time at high temperatures (650, 700°C), their curves at low temperature (500, 550°C) each show a maximum value and then decrease. The maximum values appear when the holding time is between 10 and 100 s. An increase in strength of Zr-alloy after post-irradiation annealing, which is called radiation anneal hardening (RAH), was produced by the interaction between impurity elements and radiation damage clusters [4,9,10]. The RAH occurred only on the materials which had been irradiated below 325°C, and its maximum value occurred when the heat treatment was done at a temperature approximately 100-150°C higher than the irradiation temperature of 325°C [9]. RAH was also found in Zircaloy-2, which was irradiated at 325°C and experienced post-irradiation annealing of 400-600°C for 3600 s [10]. Finally, RAH was produced under transient annealing conditions of 14-28°C/s when heated up to a maximum temperature of 600°C [4]. Checking the materials and irradiation conditions of the previous data with the temperature range in which the maximum values appeared in Figs. 1 and 2, and considering the oxygen concentration of 0.12% in the specimens, it is expected that RAH also had occurred for specimens in the present study.
4.2. Recove~ process Fig. 5 compares tensile strength obtained in the present study with that of cold worked Zircaloy-4 [1]. Softening of irradiated specimen is faster than that of cold worked material for isothermal annealing at 600°C. This trend remains even at 650°C although the difference is reduced. It has been reported that the recovery of neutron irradiated zirconium alloys was faster than that of cold-worked ones [4,6]. The softening of materials is caused by dissolution of
~ e~
g_~_
500
I
I IIIIIII
I
I IIIIIll
coo 65o'c [] IlUnln'idiataaZry-4[ll
(As-irradiated)
• Thhl Study ( I h - i d l a t e d )
C)
zO 4 0 0
300
~, uJ I'--
200
I I I IIIlll
I
100
I I Illlll
I
101
HOLDING
I I III111
102
TIME
~, I I IIIIII
103
I I I IIII
10'
(s)
Fig. 5. Difference in softening rate between irradiated and cold worked Zircaloys.
point defects, rearrangement of dislocations and a recrystallization process. There was a linear relationship between the half value width and the increase in strength as indicated in Fig. 3. Since the increment of the strength is proportional to the 1 / 2 power of the dislocation density [11] and the half value width is also proportional to the 1/2 power of the dislocation density [12], the change in the fraction of recovery of the half value width is considered to be due to a decrease of dislocation density (dislocations are radiation defects) rather than recrystallization. In the case of irradiated Zircaloy-2 in the present study, it is considered that the reduction of dislocation density is induced by isothermal annealing of 500°C for 15 s, and the dislocation density after heat treatment at 600°C for 15 s is close to that of the unirradiated material.
4.3. Recocer), equation The remaining fraction of irradiation hardening after heat treatment, F, is defined by the following equation,
F = P / ( P o - Pf),
(l)
where P is the mechanical properties of the irradiated specimens after heat treatment, P0 and Pf are the mechanical properties of as-irradiated and unirradiated specimens, respectively. Bauer and Lowry [4] carried out isothermal annealing on Zircaloy-4 fuel cladding irradiated up to 30 G W d / t , and formulated the mechanical properties using Johnson-Mehl-Avrami (JHA) law. In the present study, the JMA law was adopted for regression analysis of the results of tensile strength and hardness. AIhough the amount of recovery for the term of transient heating up to the target temperature and down to room temperature after isothermal annealing is considered to be small [1], a correction of the holding time was performed as follows to refine the equation of regression analysis. A provisional equation of recovery was obtained using the target holding time. The fraction of recovery during transient temperature rise and drop before and after the isothermal annealing was calculated by the provisional equation. This excess fraction
T. Torimaru et al. / Journal of Nuclear Materials 238 (1996) 169-174 was translated into the excess holding time at each target temperature. Finally, the following expression was obtained from the least squares method using the corrected holding time, which was the sum of the target holding time and the excess one, F= exp[- (Kt)°5],
(2)
173
since the density of irradiation defects saturates when fluence is more than 12 × 10 25 n / m 2 under BWR conditions. Therefore, mechanical properties, after boiling transition or dry-out, of irradiated Zircaloy-2 with higher fluences than that adopted in the present study can be estimated well using Eq. (2), although the equation does not include a term for fluence.
K = 2.24 × 10 l' exp( - 2 . 0 9 × 105/RT), where 0.5 and 2.24 × 10 ~' are material constants, 2.09 × 105 is activation energy (J/tool), R is gas constant, 8.314 ( J / t o o l / K ) , T is annealing temperature (K), and t is holding time (s). The calculated curves and test results in the present study show small deviations compared to those for cold-worked materials [1,2]. These deviations would be caused by a complex softening mechanism of irradiated materials which would include the disappearance of radiation defects and RAH during heat treatment as described above. Hence, the time obtained using Eq. (2) to recover the irradiation hardening up to 95% of that was close to the full recovery time for irradiated Zirconium alloys [4,13]. Therefore, it is possible to use Eq. (2) for evaluation of mechanical properties of irradiated Zircaly-2 which have experienced recrystallized annealing. According to Eq. (2), 209 kJ/mol was obtained as the activation energy. Isothermal recrystallization data of cold-worked Zircaloy [2,14] could be successfully expressed using chemical rate equation; activation energy for the recrystallization process was found to be 345 kJ/mol[2]. B ~ o et al. [1] also reported that the activation energy and recovery equation obtained from short time recrystallization tests on cold-worked Zircaloy-4 agreed with that reported by Hunt and Schulson [2]. Bauer and Lowry [4] have reported that the activation energy for both recovery and recrystallization obtained by fitting data was 247 kJ/mol, which was relatively close to that for self-diffusion in alpha zirconium, for Zircaloy-4 fuel cladding tubes irradiated up to approximately 4 × 1025 n / m 2. Activation energies for dislocation glide and dislocation climb have been reported as 174 kJ/mol and 270 kJ/mol, respectively [15]. Activation energies of 338 kJ/mol for the growth of recrystallized grains, and 270 kJ/mol for the self-diffusion of Zr [16] have also been reported; Dollins expected that activation energy would increase with recovery progress based on the numerical analysis. The activation energy of 209 kJ/mol obtained in the present study is roughly 60% of the value for cold worked material and ranks between that for dislocation glide and self-diffusion in alpha zirconium. From these findings, it is presumed that the recovery process which irradiated Zircaloy-2 with recrystallized annealing experienced is not recrystallization, but disappearance of dislocation loops generated during radiation. The recovery process and rate for irradiated Zircaloy-2, which experienced recrystallized annealing during manufacturing will not change at fluences over 2 × 1025 n / m 2
5. Conclusions (1) The irradiation hardening in hardness at room temperature and ultimate tensile strength at 343°C recovered to approximately 80% of that after heat treatment at 600700°C for less than 15 s. Fatigue life and half value width of X-ray analysis line after annealing at 600°C for 15 s recovered to those values of unirradiated cladding tube. (2) The maximum values appeared after annealing at less than 550°C for 10-100 s on the curves of mechanical properties against holding time. But it was not observed for any heating at 600-700°C. This hardening phenomenon was considered to be due to radiation anneal hardening. (3) As the recovery of mechanical properties of irradiated Zr liner cladding with recrystallized annealing is caused by disappearance of radiation defects such as dislocation loops, the recovery rate was faster than that of cold worked and stress relieved materials. (4) From the results of tensile strength and hardness, an equation to estimate the remaining fraction of radiation hardening of irradiated Zr liner cladding after boiling transition or dry-out was proposed:
F = e x p [ - ( K,)°'5], K = 2.24 × 1011 exp( - 2 . 0 9 × 105/RT), where F, t, R, and T are the remaining fraction of irradiation hardening after heat treatment, the holding time (s) during annealing, gas constant ( J / m o l / K ) and the temperature (K), respectively.
Acknowledgements Most of this work was conducted under the contracts with Tokyo Electric Power Co., Tohoku Electric Power Co., Chubu Electric Power Co., Hokuriku Electric Power Co., Chugoku Electric Power Co., Japan Atomic Power Co., Hitachi Ltd., Toshiba Corp., and Nippon Nuclear Fuel Development Co., Ltd. The authors are indebted to these member organizations. The authors express their heartfelt thanks to Mr T. Anegawa of TEPCO, Mr T. Hirose of Hitachi Ltd., and Mr T. Matsumoto of Toshiba Corp. for helpful advice on determining the experimental conditions.
174
T. Torimaru et al. / Journal of Nuclear Materials 238 (1996) 169-174
References [1] G. B~o, H.G. Wessjohann and A.R. Nicoll, ASTM STP 824 (1984) 200. [2] C.E.L. Hunt and E.M. Schulson, J. Nucl. Mater. 92 (1980) 184. [3] E.D. Hindle, Dimensional Stability and Mechanical Behavior of Irradiated Metal and Alloys (British Nuclear Society, 1983) p. 183. [4] A.A. Bauer and L.M. Lowry, Nucl. Tech. 41 (1978) 359. [5] M. Nakatsuka, T. Kubo and Y. Hayashi, ASTM STP 1132 (1991) 230. [6] G.J.C. Carpenter and J.F. Watters, ASTM STP 551 (1974) 400.
[7] G.J.C. Carpenter and D.O. Northwood, J. Nucl. Mater. 56 (1975) 260. [8] C.C. Dollins, Radiat. Eft. 16 (1972) 271. [9] K.U. Snowden and K. Veevers, Radiat. Eft. 20 (1973) 169. [10] K. Veevers, W.B. Rotsey and K.U. Snowden, ASTM STP 458 (1969) 194. [11] R.A. Holt, J. Nucl. Mater. 59 (1976) 234. [12] M. Wilkens, Phys. Status Solidi (a)2 (1970) 359. [13] E.E. Juenke and J.F. White, General Electric Co., Report GEMP-731 (1970). [14] D. Lee, J. Nucl. Mater. 37 (1970) 159. [I 5] P. Merle, C. Vauglin, G. Fantozzi, J.L. Derep and D. Charquet, ASTM STP 939 (1987) 555. [16] C.C. Dollins, J. Nucl. Mater. 59 (1975) 61.
JoillTlalof NuclearMaterials 238 (1996) 169-174
Changes in mechanical properties of irradiated Zircaloy-2 fuel cladding due to short term annealing Tadahiko Torimaru *, Takayoshi Yasuda
1, Masafumi Nakatsuka
Nippon Nuclear Fuel Development Co., Ltd, 2163, Narita-cho, Oarai-machi, Higashi-ibarald-gun, Ibaraki-ken, 311-13, Japan Received 13 February 1995; accepted 19 July 1996
Abstract
Zirconium-lined fuel cladding tubes irradiated to 2.7 x 1 0 25 n / m 2 (E > 1 MeV) in a BWR, which had experienced recrystallized annealing in the final process in their manufacture, were heat treated at 500-700°C for 5-600 s to simulate short term dry-out. Tensile tests, hardness measurements, fatigue tests and X-ray analyses were made on those specimens. The irradiation hardening in hardness at room temperature and ultimate tensile strength at 343*(2 recovered to approximately 80% of that after heat treatment at 600-700°C for less than 15 s. Fatigue life and half value width of X-ray analysis recovered to these of unirradiated cladding tube after annealing for 15 s at 600°C. These recovery rates were faster than those on cold worked and stress relieved zirconium alloys. An equation to predict the remaining fraction of hardening was proposed by using the regression analysis on tensile strength and hardness values.
1. Introduction
Zr liner fuel claddings, which are lined with a zirconium barrier on the inner surface of the base cladding of Zircaloy-2 and heat treated under the recrystallization condition, are mainly used in BWRs. When a drop in cooling ability occurs for a short time in a BWR, the temperature of the fuel rods increases by a boiling transition or dry-out. The rise in temperature changes the mechanical properties of the cladding that has already been hardened by neutron induced damage. Therefore, detailed data on the mechanical properties of irradiated Zr liner fuel cladding tubes after short term annealing are indispensable for evaluation of fuel performance. However, earlier studies have only dealt with recovery and recrystallization behavior of coldworked material [1-3] and with recovery behavior of irradiated Zircaloy-4 in the stress relief condition during the final manufacturing process [4].
* Corresponding author. Tel.: +81-29 267 4141; fax: +81-29 267 7135; e-mail: [email protected]. l Present address: Hitachi Nuclear EngineeringCo., Ltd.
The purposes of the present study are to clarify the recovery behavior of irradiated Zr liner fuel cladding from several mechanical properties and half value widths of X-ray diffraction lines after short term annealing, and to formulate the fraction of recovery as a function of annealing temperature and time.
2. Experimental details 2.1. Test material Test materials were retrieved from irradiated Zr liner fuel cladding for which fluence and mean burn up had amounted to 2.7X 1025 n / m 2 ( E > 1 MeV) and 18 G W d / t in a commercial BWR. This cladding had a hydrogen content of 50 ppm and average oxide layer thickness of 28 ~m on the outer surface. Ring specimens at a point approximately 2500 mm from the rod lower end of the cladding were cut to 5 mm in length, and fuel pellets were removed. Dimensions, irradiation conditions and main chemical composition of the cladding are summarized in Tables 1 and 2.
0022-3115/96/$15.00 Copyright © 1996 Elsevier Science B.V. All rights reserved. PII S0022-3115(96)00451-5
170
T. Torimaru et al. / Journal of Nuclear Materials 238 (1996) 169-174
Table 1 Dimensions and irradiation conditions of fuel rods Outer and inner diameters (mm)
Material
Thickness of liner (Ixm)
Irradiation temperature (K)
Neutron fluence (n/m e) (E > 1 MeV)
12.27 10.55
Zircaloy-2
~ 80
~ 620
2.6 to 2.8 X 1025
2.2. Heat treatment
A radiation heating furnace with infrared lamps installed as a heat source was employed for short time isothermal annealing to simulate the temperature changes during dry-out. Since mechanical properties of irradiated cladding were thought to change quickly over 500°C, it was desirable for the heating and cooling rates to be as fast as possible. The furnace could provide rapid heating from 500°C to 700°C at a rate of 4 0 ° C / s and rapid cooling from 700°C to 5000C at 30°C/s. Heat treatment was conducted in argon gas to prevent specimen oxidation. A thermocouple was spot welded on the cross-section of each ring specimen, and their temperatures during heat treatment were measured. The deviation from the desired temperature was held to within + 5°C.
to the loading rods at two positions which were 180 degrees apart from each other. Fully reversed bending strains were induced with 1 Hz in the specimen by the reciprocal movement of the loading rods. The tests were conducted at 350°C in argon gas. The time when the load dropped to approximately 60% of the initial load was judged to represent fatigue failure, and the number of cycles then was recorded. The details of the fatigue tester and test method have been described elsewhere [5]. The tube cross-sections of the specimens used for hardness tests were polished once again for X-ray diffraction. Diffraction intensity curves were obtained by using an X-ray diffractmeter with a Cu K s ray. Other operating conditions were: tube voltage, 40 kV; tube current, 30 mA; beam diameter, 1.0 mm; detector scan speed, 0.5°C/rain. The half value width for each diffraction plane was measured from the diffraction curves.
2.3. Experimental procedure
The ring shaped specimens had a 5 mm length to allow the plane stress condition to be established and they had a cross-section larger than 5 mm. The specimens were tensile tested using two semicircular tension jigs. The ultimate tensile strength (UTS), the distance of cross-head movement up to failure and the reductions of area were measured. The tests were conducted at 343°C in air with the strain rate of approximately 0.5%/min. The gauge length of a ring specimen was assumed to be one half of the tube diameter. The ring shaped specimens had a 5 mm length were fixed into resin and polished their cross-section. Vickers hardness on the cross-section of each specimen at room temperature was measured with force of 200 g (1.96 N) for Zircaloy-2 and 50 g (0.49 N) for zirconium liner. Cladding tubes, machined into C-shape specimens, were adopted for the bending fatigue test. A specimen was fixed
3. Experimental results Fig. l(a)-(c) show ultimate tensile strength, cross-head movement at failure and reduction of area against the holding time of annealing, respectively. The tensile strength after annealings at 700°C × 5 s and 600-650°C x 15 s, almost recovered to that of unirradiated specimen. Those after the annealing at 500-575°C tended to decrease with annealing time and show the maximum value at an holding time. These maximum values were considered to be due to radiation anneal hardening (RAH). This issue will be discussed later. Vickers hardness changes of the irradiated and unirradiated specimens for Zircaloy-2 and zirconium liner are depicted in Fig. 2. The closed circle and bar show the average value and standard deviation of unirradiated specimens with and without heat treatment. The hardness of the
Table 2 Chemical composition of fuel cladding (wt% or wtppm) Zircaloy-2 a Zr-liner b
Sn
Fe
Cr
Ni
O
1.42% < 10 ppm
0.14% 430 ppm
0.10% 54 ppm
0.05% < 35 ppm
1150 ppm 350 ppm
Chemical composition satisfied ASTM B353, recrystallized before irradiation. b High purity sponge zirconium.
171
T. Torimaru et al. / Journal of Nuclear Materials 238 (1996) 169-174 14" 500"C -4-550"C ~ 575"C-o- 600"C-0--650"C~ ?O0"C]
Holding Time 15s
~ 600
100
....
,
ao
>~
400
60
/
40
/
~i
°i--~
(b)
1.5
o ~
1,0 0,5
As-irradiated
80 ~
.
(c)
40
0
~
................. I0
i |
i|l
I00
I000
HOLDING TIME (s) Fig. I. Changes in the mechanical properties with holding time of annealing in tensile tests at 343°C (616 K).
["e-500"C
260
Unlrradlated
//
~ , I (211) (201) (110) I
/
i .
i
.
i
5oo,~o ~
°
,
i
*1
° ,
i
i
Mo 7m
Fig. 3. Recovery of half-value width with annealing temperature.
2.0 25
0 ~
-/--.
/
i~"
/o, i
I
......
,
~-Irmdlated//. / '
20
3.0
•
4,-.
/*o,. ,
~ 300
**
~ .....
"P'$$0"C
-~'-575"C
"'o--800"(2
"-o-650"C
.,~e-n'z'adiated
.-A.-?O0"C
I
Zry-2
unirradiated specimens showed no great changes, and the effect of thermal shock during annealing could be ignored. For irradiated specimens, the hardness after annealing above 600 for 15 s or more almost recovered to that for the unirradiated specimen. The recovery curves of hardness after the annealing at 500-550°C also showed RAH. Number of cycles to fatigue failures under the total bending strain amplitude of 0.55% were measured. The failure cycles of unirradiated and as-irradiated specimens were 2700 and 20350, respectively. The failure cycle after heat treatment at 600°C for 15 s nearly recovered to that for unirradiated specimen. Fig. 3 plots the recovery in half value width of X-ray diffraction lines against the annealing temperatures for the three diffraction planes. The half value widths after annealing above 600°C nearly recovered to that for unirradiated specimen. These results corresponded to those of the mechanical tests. Fig. 4 shows the half value widths for the (211) plane plotted against the Vickers hardness of the Zircaloy-2. Half value widths and Vickers hardness had a linear relationship. This result meant that the recovery of hardness was due to disappearance of irradiation defects.
9-20 ~ N m
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Uni~
180 160 140
AB-irradi~ted
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Holding tlme 15s (Temperature °C)
~ '~
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>~
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As-lrradlate~
Unlrradlated
120 I00 80 60
o.,
Unirradiated ~ .......... 10
100
1000
HOLDING TIME (s) Fig. 2. Changes in Vickers hardness at room temperature with holding time of annealing.
0.2
160
i
.
i
.
i
.
I
180 200 220 240 VICKERS HARDNESS (Hv)
i
260
Fig. 4. Relationship between half value width and Vickers hardness.
172
T. Torimaru et a l . / Journal of Nuclear Materials 238 (1996) 169-174 600
4. Discussion
I
I I IIIIII
I
i IIIIIII
( ~ )
4.1. Time dependence of softening As shown in Figs. 1 and 2, softening of irradiated specimen rapidly progresses during short term annealings, 5-15 s, and then slows down. Generally, many defects have been found to be uniformly formed in irradiated fuel cladding tubes. Small defects, which cannot be distinguished using TEM, were produced in Zircaloy-2 tubes irradiated at 573 K to a fast neutron fluence of less than 2-3 × 1023 n / m 2 [6]. Very small dislocation loops with an average diameter of approximately 37 nm were formed in cladding tubes irradiated at 50°C to the fluence of 5.5 × 1019 n/cm2[7]. From these reports, it is clear that defects made by neutron irradiation are very small. According to Dollins" recovery model [8], large defects are relatively stable in the initial process during post irradiation annealing, although small loops rapidly dissolve. This analysis explains the fast recovery behavior shown in Figs. 1 and 2. While both tensile strength and hardness drop monotonously with the holding time at high temperatures (650, 700°C), their curves at low temperature (500, 550°C) each show a maximum value and then decrease. The maximum values appear when the holding time is between 10 and 100 s. An increase in strength of Zr-alloy after post-irradiation annealing, which is called radiation anneal hardening (RAH), was produced by the interaction between impurity elements and radiation damage clusters [4,9,10]. The RAH occurred only on the materials which had been irradiated below 325°C, and its maximum value occurred when the heat treatment was done at a temperature approximately 100-150°C higher than the irradiation temperature of 325°C [9]. RAH was also found in Zircaloy-2, which was irradiated at 325°C and experienced post-irradiation annealing of 400-600°C for 3600 s [10]. Finally, RAH was produced under transient annealing conditions of 14-28°C/s when heated up to a maximum temperature of 600°C [4]. Checking the materials and irradiation conditions of the previous data with the temperature range in which the maximum values appeared in Figs. 1 and 2, and considering the oxygen concentration of 0.12% in the specimens, it is expected that RAH also had occurred for specimens in the present study.
4.2. Recove~ process Fig. 5 compares tensile strength obtained in the present study with that of cold worked Zircaloy-4 [1]. Softening of irradiated specimen is faster than that of cold worked material for isothermal annealing at 600°C. This trend remains even at 650°C although the difference is reduced. It has been reported that the recovery of neutron irradiated zirconium alloys was faster than that of cold-worked ones [4,6]. The softening of materials is caused by dissolution of
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100
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101
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TIME
~, I I IIIIII
103
I I I IIII
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(s)
Fig. 5. Difference in softening rate between irradiated and cold worked Zircaloys.
point defects, rearrangement of dislocations and a recrystallization process. There was a linear relationship between the half value width and the increase in strength as indicated in Fig. 3. Since the increment of the strength is proportional to the 1 / 2 power of the dislocation density [11] and the half value width is also proportional to the 1/2 power of the dislocation density [12], the change in the fraction of recovery of the half value width is considered to be due to a decrease of dislocation density (dislocations are radiation defects) rather than recrystallization. In the case of irradiated Zircaloy-2 in the present study, it is considered that the reduction of dislocation density is induced by isothermal annealing of 500°C for 15 s, and the dislocation density after heat treatment at 600°C for 15 s is close to that of the unirradiated material.
4.3. Recocer), equation The remaining fraction of irradiation hardening after heat treatment, F, is defined by the following equation,
F = P / ( P o - Pf),
(l)
where P is the mechanical properties of the irradiated specimens after heat treatment, P0 and Pf are the mechanical properties of as-irradiated and unirradiated specimens, respectively. Bauer and Lowry [4] carried out isothermal annealing on Zircaloy-4 fuel cladding irradiated up to 30 G W d / t , and formulated the mechanical properties using Johnson-Mehl-Avrami (JHA) law. In the present study, the JMA law was adopted for regression analysis of the results of tensile strength and hardness. AIhough the amount of recovery for the term of transient heating up to the target temperature and down to room temperature after isothermal annealing is considered to be small [1], a correction of the holding time was performed as follows to refine the equation of regression analysis. A provisional equation of recovery was obtained using the target holding time. The fraction of recovery during transient temperature rise and drop before and after the isothermal annealing was calculated by the provisional equation. This excess fraction
T. Torimaru et al. / Journal of Nuclear Materials 238 (1996) 169-174 was translated into the excess holding time at each target temperature. Finally, the following expression was obtained from the least squares method using the corrected holding time, which was the sum of the target holding time and the excess one, F= exp[- (Kt)°5],
(2)
173
since the density of irradiation defects saturates when fluence is more than 12 × 10 25 n / m 2 under BWR conditions. Therefore, mechanical properties, after boiling transition or dry-out, of irradiated Zircaloy-2 with higher fluences than that adopted in the present study can be estimated well using Eq. (2), although the equation does not include a term for fluence.
K = 2.24 × 10 l' exp( - 2 . 0 9 × 105/RT), where 0.5 and 2.24 × 10 ~' are material constants, 2.09 × 105 is activation energy (J/tool), R is gas constant, 8.314 ( J / t o o l / K ) , T is annealing temperature (K), and t is holding time (s). The calculated curves and test results in the present study show small deviations compared to those for cold-worked materials [1,2]. These deviations would be caused by a complex softening mechanism of irradiated materials which would include the disappearance of radiation defects and RAH during heat treatment as described above. Hence, the time obtained using Eq. (2) to recover the irradiation hardening up to 95% of that was close to the full recovery time for irradiated Zirconium alloys [4,13]. Therefore, it is possible to use Eq. (2) for evaluation of mechanical properties of irradiated Zircaly-2 which have experienced recrystallized annealing. According to Eq. (2), 209 kJ/mol was obtained as the activation energy. Isothermal recrystallization data of cold-worked Zircaloy [2,14] could be successfully expressed using chemical rate equation; activation energy for the recrystallization process was found to be 345 kJ/mol[2]. B ~ o et al. [1] also reported that the activation energy and recovery equation obtained from short time recrystallization tests on cold-worked Zircaloy-4 agreed with that reported by Hunt and Schulson [2]. Bauer and Lowry [4] have reported that the activation energy for both recovery and recrystallization obtained by fitting data was 247 kJ/mol, which was relatively close to that for self-diffusion in alpha zirconium, for Zircaloy-4 fuel cladding tubes irradiated up to approximately 4 × 1025 n / m 2. Activation energies for dislocation glide and dislocation climb have been reported as 174 kJ/mol and 270 kJ/mol, respectively [15]. Activation energies of 338 kJ/mol for the growth of recrystallized grains, and 270 kJ/mol for the self-diffusion of Zr [16] have also been reported; Dollins expected that activation energy would increase with recovery progress based on the numerical analysis. The activation energy of 209 kJ/mol obtained in the present study is roughly 60% of the value for cold worked material and ranks between that for dislocation glide and self-diffusion in alpha zirconium. From these findings, it is presumed that the recovery process which irradiated Zircaloy-2 with recrystallized annealing experienced is not recrystallization, but disappearance of dislocation loops generated during radiation. The recovery process and rate for irradiated Zircaloy-2, which experienced recrystallized annealing during manufacturing will not change at fluences over 2 × 1025 n / m 2
5. Conclusions (1) The irradiation hardening in hardness at room temperature and ultimate tensile strength at 343°C recovered to approximately 80% of that after heat treatment at 600700°C for less than 15 s. Fatigue life and half value width of X-ray analysis line after annealing at 600°C for 15 s recovered to those values of unirradiated cladding tube. (2) The maximum values appeared after annealing at less than 550°C for 10-100 s on the curves of mechanical properties against holding time. But it was not observed for any heating at 600-700°C. This hardening phenomenon was considered to be due to radiation anneal hardening. (3) As the recovery of mechanical properties of irradiated Zr liner cladding with recrystallized annealing is caused by disappearance of radiation defects such as dislocation loops, the recovery rate was faster than that of cold worked and stress relieved materials. (4) From the results of tensile strength and hardness, an equation to estimate the remaining fraction of radiation hardening of irradiated Zr liner cladding after boiling transition or dry-out was proposed:
F = e x p [ - ( K,)°'5], K = 2.24 × 1011 exp( - 2 . 0 9 × 105/RT), where F, t, R, and T are the remaining fraction of irradiation hardening after heat treatment, the holding time (s) during annealing, gas constant ( J / m o l / K ) and the temperature (K), respectively.
Acknowledgements Most of this work was conducted under the contracts with Tokyo Electric Power Co., Tohoku Electric Power Co., Chubu Electric Power Co., Hokuriku Electric Power Co., Chugoku Electric Power Co., Japan Atomic Power Co., Hitachi Ltd., Toshiba Corp., and Nippon Nuclear Fuel Development Co., Ltd. The authors are indebted to these member organizations. The authors express their heartfelt thanks to Mr T. Anegawa of TEPCO, Mr T. Hirose of Hitachi Ltd., and Mr T. Matsumoto of Toshiba Corp. for helpful advice on determining the experimental conditions.
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T. Torimaru et al. / Journal of Nuclear Materials 238 (1996) 169-174
References [1] G. B~o, H.G. Wessjohann and A.R. Nicoll, ASTM STP 824 (1984) 200. [2] C.E.L. Hunt and E.M. Schulson, J. Nucl. Mater. 92 (1980) 184. [3] E.D. Hindle, Dimensional Stability and Mechanical Behavior of Irradiated Metal and Alloys (British Nuclear Society, 1983) p. 183. [4] A.A. Bauer and L.M. Lowry, Nucl. Tech. 41 (1978) 359. [5] M. Nakatsuka, T. Kubo and Y. Hayashi, ASTM STP 1132 (1991) 230. [6] G.J.C. Carpenter and J.F. Watters, ASTM STP 551 (1974) 400.
[7] G.J.C. Carpenter and D.O. Northwood, J. Nucl. Mater. 56 (1975) 260. [8] C.C. Dollins, Radiat. Eft. 16 (1972) 271. [9] K.U. Snowden and K. Veevers, Radiat. Eft. 20 (1973) 169. [10] K. Veevers, W.B. Rotsey and K.U. Snowden, ASTM STP 458 (1969) 194. [11] R.A. Holt, J. Nucl. Mater. 59 (1976) 234. [12] M. Wilkens, Phys. Status Solidi (a)2 (1970) 359. [13] E.E. Juenke and J.F. White, General Electric Co., Report GEMP-731 (1970). [14] D. Lee, J. Nucl. Mater. 37 (1970) 159. [I 5] P. Merle, C. Vauglin, G. Fantozzi, J.L. Derep and D. Charquet, ASTM STP 939 (1987) 555. [16] C.C. Dollins, J. Nucl. Mater. 59 (1975) 61.