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Hoop strength and ductility evaluation of irradiated fuel cladding
Nuclear Engineering and Design 239 (2009) 254–260
Contents lists available at ScienceDirect
Nuclear Engineering and Design journal homepage: www.elsevier.com/locate/nucengdes
Hoop strength and ductility evaluation of irradiated fuel cladding Sun-Ki Kim ∗ , Je-Geon Bang, Dae-Ho Kim, Ik-Sung Lim, Yong-Sik Yang, Kun-Woo Song, Do-Sik Kim Korea Atomic Energy Research Institute, 150 Deokjin-dong, Yuseong-gu, Daejeon 305-353, Republic of Korea
a r t i c l e
i n f o
Article history: Received 9 November 2007 Received in revised form 29 October 2008 Accepted 29 October 2008
a b s t r a c t The objective of this study is to evaluate the hoop-directional mechanical properties comprising strength such as yield strength and ultimate tensile strength as well as mechanical ductility such as uniform elongation and total elongation. Therefore, in this paper, the ring tensile tests were performed in order to evaluate the mechanical properties of high burn-up fuel cladding under a hoop loading condition in a hot cell. The tests were performed with Zircaloy-4 nuclear fuel cladding whose burn-up is approximately 65,000 MWd/tU in the temperature range of room temperature to 800 ◦ C. All the experiments were carried out at a constant strain rate of 0.01/s. On the basis of the ring tensile tests for a high burn-up Zircalay-4 cladding, the following conclusions were drawn. Firstly, the mechanical properties are abruptly degraded beyond 600 ◦ C, which corresponds to a design-basis accident condition such as a RIA. Secondly, the un-irradiated fuel cladding showed ductile fracture behaviors such as 45◦ shear type fracture, cup and cone type fracture, cup and cup type fracture and chisel edge type fracture. While the high burn-up Zircalay-4 cladding showed a brittle fracture behavior even at the high temperatures (e.g. over 600 ◦ C) which are achievable during a RIA. Thirdly, in the case of the high burn-up Zircalay-4 cladding, the strength, ductility and the energy to break are strongly dependent on the material property itself which are degraded by oxidation and hydriding during an operation rather than the temperature. Fourthly, hydride rim formation in the vicinity of metal–oxide interface can play an important role in the degradation of the mechanical properties for high burn-up fuel cladding. © 2008 Elsevier B.V. All rights reserved.
1. Introduction During a steady-state operation of light water reactors, the mechanical behavior of the zirconium-based fuel cladding degrades due to a combination of oxidation, hydriding, and radiation damage. In an effort to increase the operating efficiency through the use of longer fuel cycles, and to reduce the volume of waste associated with the core reloads, utilities have a strong incentive to increase the average discharge burn-up of the fuel assemblies. Further increases in the operating efficiency of power reactors can also be achieved by increasing the coolant outlet temperature. However, both of these changes in a reactor operation enhance the cladding degradation, which may increase the likelihood of a cladding failure during design-basis accidents. One such postulated design-basis accident scenario is the reactivity-initiated accident (RIA) in a pressurized water reactor (PWR) caused by the ejection of a control rod from the core, which would cause a rapid increase of the reactivity and the thermal
energy in the fuel (Meyer et al., 1986). The increase in fuel temperature resulting from an RIA induces a rapid fuel expansion, causing a severe pellet-cladding mechanical interaction (PCMI). This PCMI forces the cladding to experience a multiaxial tension such that the maximum principal strain is in the hoop (i.e., transverse) direction of the cladding tube. The survivability of a fuel cladding irradiated to a high burn-up under postulated RIA conditions is thus a response to a combination of the mechanics of a loading and the material degradation during a reactor operation. While such data is available for the axial deformation behavior of cladding tubes, relatively little has been reported in the open literature on the uniaxial tension behavior in the hoop direction of Zircaloy-4 cladding. Accordingly, it is essential to investigate the uniaxial tension behavior in the hoop direction of high burn-up Zircaloy-4 cladding. In this study, ring tensile tests are applied to obtain the data regarding the uniaxial hoop direction deformation behavior. 2. Experimental procedure
∗ Corresponding author. Tel.: +82 42 868 8661; fax: +82 42 863 0565. E-mail address: [email protected] (S.-K. Kim). 0029-5493/$ – see front matter © 2008 Elsevier B.V. All rights reserved. doi:10.1016/j.nucengdes.2008.10.024
The ring tensile specimen used in this study is a Westinghouse 17 × 17 type (Vantage-5H) Zircaloy-4 cladding irradiated for 3 cycles
S.-K. Kim et al. / Nuclear Engineering and Design 239 (2009) 254–260
255
Fig. 1. A schematic of Zircaloy-4 fuel rod irradiated in Ulchin Unit 2.
in the Ulchin Unit 2 pressurized water reactor, whose average burnup was estimated to be approximately 65 GWd/tU (see Fig. 1). The irradiated fuel rod was transferred to PIEF at KAERI, cut into approximately 13 cm length segments with a diamond low speed saw, and then the UO2 pellet inside the fuel rod segment was removed by a mechanical grinding with a drill-attached defueling machine in a hot cell in IMEF. Fig. 2 shows a schematic illustration of the procedure of ring tensile test. The defueling machine is shown in Fig. 3. The specimens for the mechanical properties evaluation were fabricated with a diamond wheel grinder from the defueled cladding segments (see Fig. 4). The dimensions and shape of the ring tensile specimen were designed in order to ensure that any deformation is limited to the gage section of the specimen, so that the uniform uniaxial hoop strain in the gage section could be at its maximal. The dimensions and a photograph of the ring tensile specimen are shown in Figs. 5 and 6, respectively. The gage sections of the specimens were oriented at the top and bottom of the half cylinder of the grip, such that a constant specimen curvature can be maintained during a deformation. A photograph of the jig and the grip is shown in Fig. 7. The interface was lubricated with a graphite-containing vacuum grease lubricant at the beginning of each test to minimize a loss of the applied load. The ring tensile tests were performed in a hot cell with the Instron Servohydraulic System, Model 8562. The tests were performed at 25, 135, 200, 300, 350, 400, 600, and 800 ◦ C, and the initial strain rate was maintained at 0.01/s.
Fig. 3. Photograph of the defueling machine.
The hydride morphologies were observed, which are shown in Fig. 8. As seen in the figure, a hydride rim was formed around the metal substrate/oxide interface. This hydride rim is believed to cause a decrease of the ductility of the cladding tube (Daum et al., 2002).
Fig. 2. A schematic illustration for the procedure of ring tensile tests.
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Fig. 4. Photograph of the machine for the ring tensile specimen fabrication.
Fig. 7. Photograph of the jig and the grip for the ring tensile specimen.
Fig. 5. Dimension of the ring tensile specimen.
3. Results and discussion 3.1. Effects of the lubricant and the gage section position The friction between the inner surface of the ring specimen and the half-cylinder die inserts causes a loss of the applied load. Accordingly, it is desirable to diminish the loss of the applied load by
Fig. 6. Photograph of the ring tensile specimen.
a friction. In this study, the effects of a lubrication were investigated to determine the lubricant for this ring tensile test. The ring tensile tests were performed with Teflon tape and graphite vacuum grease as the lubricant for as-received Zircaloy-4. The load–displacement correlation is shown in Fig. 9. The load–displacement behavior and elastic modulus of the Teflon tape in the elastic region coincides with those of the graphite vacuum grease. On the contrary, the load–displacement behavior is not in accordance with each other in the plastic region. The ultimate tensile strength for the Teflon tape was approximately 2.1 kN, while that of the graphite vacuum grease was approximately 1.8 kN. This difference in the ultimate tensile strength is caused by the difference in their friction coefficients. From a point of view regarding a friction, Teflon tape is more desirable than graphite vacuum grease. But, the maximum allowable temperature of the Teflon tape is approximately 260 ◦ C, so graphite vacuum grease whose maximum allowable temperature is approximately 1000 ◦ C was adopted as a lubricant in this experiment. The load–displacement behavior as a positioning direction of the gauge section was also investigated. As shown in Fig. 10, the ultimate tensile strength and displacement at the fracture showed a different behavior in the up and down positioning direction and the left and right positioning direction. The ultimate tensile strength for the up and down positioning direction was evaluated to be a little higher than that for the left and right positioning direction. And the displacement at fracture direction was evaluated to be a little higher than that for the left and right positioning direction too. When the gauge section of the specimen is located towards the left and right positioning direction in the two half-cylinder inserts system, the elongation of the gauge section may be flattened because there is no circumferential elongation by the central dog-bone insert. Because the specimen grip part in this experiment does not contain the central dog-bone insert, the gauge section of the specimen was located towards the up and down positioning
S.-K. Kim et al. / Nuclear Engineering and Design 239 (2009) 254–260
257
Fig. 10. Load–displacement behavior with the lubricant type.
To obtain the mechanical strength, the 0.2% offset yield strength (YS) and ultimate tensile strength (UTS) were evaluated, and the
uniform elongation (UE) and total elongation (TE) were also evaluated for the ductility. The hoop stress–strain curves at various temperatures are shown in Fig. 11. The evaluation results of the 0.2% offset YS and the UTS are shown in Figs. 12 and 13. From the figure, it is confirmed that the 0.2% offset YS and the UTS abruptly decrease with an increasing temperature. The UTS was evaluated to be 942.70 MPa at RT, 678.83 MPa at 400 ◦ C, but, it is abruptly diminished to 282.64 MPa at 600 ◦ C, which is achievable in the RIA condition. Especially, it decreases to 58.30 MPa at 800 ◦ C, an extreme condition, which corresponds to 6% of the UTS at RT. This means that the mechanical strength of the high burn-up Zircaloy-4 cladding sharply decreases in the RIA-relevant temperature ranges. The evaluation results of the UE and TE are shown in Figs. 14 and 15. The results show that both the UE and TE increase with an increasing temperature. Especially, they abruptly increase at 600 ◦ C, but become lower beyond this temperature. This peculiar behavior was also observed in the PROMETRA program (Averty et al., 2003) which is a mechanical property relevant test program in conjunction with the CABRI program simulating a RIA. The results of hoop-directional mechanical strength and ductility of irradiated Zircaloy-4 with comparison with PROMETRA database are shown in Figs. 16 and 17, respectively. As shown in the figure, the behavior of this study is consistent with the PROMETRA data. It is believed that this behavior is caused by the elongation minimum phenomenon by
Fig. 9. Load–displacement behavior with the lubricant type.
Fig. 11. Hoop stress–strain curves at various test temperatures.
Fig. 8. Optical microscopy of the Zircaloy-4 from Ulchin Unit 2.
direction to avoid a linear elongation in some regions of the gauge section. 3.2. Evaluation of the mechanical strength and the ductility
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Fig. 12. Yield strength of the un-irradiated and high burn-up fuel cladding.
Fig. 15. Total elongation of the un-irradiated and high burn-up fuel cladding.
Fig. 16. Comparison of mechanical strength with the PROMETRA database.
the dynamic strain aging of the Zircaloy-4 cladding material beyond 600 ◦ C. Fig. 13. Ultimate tensile strength of the un-irradiated and high burn-up fuel cladding.
3.3. Fracture characteristics of the ring tensile tests The stereoscope photographs of the specimens after the ring tensile test are shown in Fig. 18. Fig. 18 represents the fracture patterns of the non-irradiated and high burn-up Zircaloy-4 cladding
Fig. 14. Uniform elongation of the un-irradiated and high burn-up fuel cladding.
Fig. 17. Comparison of mechanical ductility with the PROMETRA database.
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Fig. 18. Stereoscope photograph of the specimens after the ring tensile test.
specimens. As shown in the figure, four fracture patterns such as 45◦ shear type fracture, cup and cone type fracture, cup and cup type fracture and chisel edge type fracture were observed in the non-irradiated cladding.
The fracture type was found to depend strongly on the deformation temperature. At room temperature non-irradiated Zircaloy-4 cladding tends to be fractured by 45◦ shear type fracture or cup and cone type fracture. Cup and cone type fracture was dominant at 135
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Table 1 Fracture types for the ring tensile tests of the Zircaloy-4 fuel cladding.
4. Conclusions
Temperature
On the basis of the ring tensile tests for the high burn-up Zircalay-4 cladding from Ulchin Unit 2 and the as-received nonirradiated Zircalay-4 cladding, the following conclusions were drawn. Firstly, the mechanical properties are abruptly degraded beyond 600 ◦ C, which corresponds to a design-basis accident condition such as a RIA. Secondly, the un-irradiated fuel cladding showed ductile fracture behaviors such as 45◦ shear type fracture, cup and cone type fracture, cup and cup type fracture and chisel edge type fracture, while the high burn-up Zircalay-4 cladding showed a brittle fracture behavior even at the high temperatures (e.g. over 600 ◦ C) which are achievable during a RIA. Thirdly, in the case of the high burn-up Zircalay-4 cladding, the strength, ductility and the energy to break are strongly dependent on the material property itself which are degraded by oxidation and hydriding during an operation rather than the temperature. Fourthly, the hydride rim formation in the vicinity of the metal-oxide interface can play an important role in the degradation of the mechanical properties for the high burn-up fuel cladding.
Fracture type Non-irradiated cladding
Room temperature 135 ◦ C 200 ◦ C 300 ◦ C 350 ◦ C 400 ◦ C 600 ◦ C 800 ◦ C
45◦ shear type fracture, cup and cone type fracture Cup and cone type fracture Cup and cone type fracture Cup and cone type fracture, cup and cup type fracture Cup and cup type fracture Cup and cup type fracture Chisel edge type fracture Unidentified
High burn-up cladding
Brittle fracture
and 200 ◦ C. Both cup and cone type fracture and cup and cup type fracture was observed at 300 ◦ C. Only cup and cup type fracture was observed at 350 and 400 ◦ C. At 600 ◦ C chisel edge type fracture was observed unlike at lower temperatures with few fracture surface area. At 800 ◦ C the fracture type is unclear. These fracture types are summarized in Table 1. In conclusion, the fracture type of non-irradiated Zircaloy-4 cladding depends on the temperature, and the behavior is governed by a ductile fracture with a necking. On the contrary, the fracture patterns of the high burn-up Zircaloy-4 cladding showed completely different fracture patterns from the non-irradiated Zircaloy-4 cladding. The fracture type was observed to be vertical in the tensile direction without a necking at all of the test temperatures, which is convincing evidence of the brittle fracture behavior of high burn-up fuel cladding regardless of temperature. This means that even at a high temperature, 600 or 800 ◦ C, the fracture pattern showed a brittle fracture behavior. Accordingly, it was found that the high burn-up Zircaloy-4 cladding becomes very brittle even at the high temperatures achievable during a designbasis accident (Daum et al., 2002).
Acknowledgment This work was financially supported by the Ministry of Education, Science and Technology in Korea. References Meyer, R.O., et al., 1986. A regulatory assessment of test data for reactivity initiated accidents. Nuclear Safety 37 (4), 271–288. Averty, X., et al., 2003. Tensile tests on ring specimens machined in M5 cladding irradiated 6 cycles. IRSN 2003/50. Daum, R.S., et al., 2002. On the embrittlement of Zircaloy-4 under RIA-relevant conditions. In: Zirconium in the nuclear industry: thirteenth international symposium, ASTM STP 1423, PA, pp. 702–719.
Contents lists available at ScienceDirect
Nuclear Engineering and Design journal homepage: www.elsevier.com/locate/nucengdes
Hoop strength and ductility evaluation of irradiated fuel cladding Sun-Ki Kim ∗ , Je-Geon Bang, Dae-Ho Kim, Ik-Sung Lim, Yong-Sik Yang, Kun-Woo Song, Do-Sik Kim Korea Atomic Energy Research Institute, 150 Deokjin-dong, Yuseong-gu, Daejeon 305-353, Republic of Korea
a r t i c l e
i n f o
Article history: Received 9 November 2007 Received in revised form 29 October 2008 Accepted 29 October 2008
a b s t r a c t The objective of this study is to evaluate the hoop-directional mechanical properties comprising strength such as yield strength and ultimate tensile strength as well as mechanical ductility such as uniform elongation and total elongation. Therefore, in this paper, the ring tensile tests were performed in order to evaluate the mechanical properties of high burn-up fuel cladding under a hoop loading condition in a hot cell. The tests were performed with Zircaloy-4 nuclear fuel cladding whose burn-up is approximately 65,000 MWd/tU in the temperature range of room temperature to 800 ◦ C. All the experiments were carried out at a constant strain rate of 0.01/s. On the basis of the ring tensile tests for a high burn-up Zircalay-4 cladding, the following conclusions were drawn. Firstly, the mechanical properties are abruptly degraded beyond 600 ◦ C, which corresponds to a design-basis accident condition such as a RIA. Secondly, the un-irradiated fuel cladding showed ductile fracture behaviors such as 45◦ shear type fracture, cup and cone type fracture, cup and cup type fracture and chisel edge type fracture. While the high burn-up Zircalay-4 cladding showed a brittle fracture behavior even at the high temperatures (e.g. over 600 ◦ C) which are achievable during a RIA. Thirdly, in the case of the high burn-up Zircalay-4 cladding, the strength, ductility and the energy to break are strongly dependent on the material property itself which are degraded by oxidation and hydriding during an operation rather than the temperature. Fourthly, hydride rim formation in the vicinity of metal–oxide interface can play an important role in the degradation of the mechanical properties for high burn-up fuel cladding. © 2008 Elsevier B.V. All rights reserved.
1. Introduction During a steady-state operation of light water reactors, the mechanical behavior of the zirconium-based fuel cladding degrades due to a combination of oxidation, hydriding, and radiation damage. In an effort to increase the operating efficiency through the use of longer fuel cycles, and to reduce the volume of waste associated with the core reloads, utilities have a strong incentive to increase the average discharge burn-up of the fuel assemblies. Further increases in the operating efficiency of power reactors can also be achieved by increasing the coolant outlet temperature. However, both of these changes in a reactor operation enhance the cladding degradation, which may increase the likelihood of a cladding failure during design-basis accidents. One such postulated design-basis accident scenario is the reactivity-initiated accident (RIA) in a pressurized water reactor (PWR) caused by the ejection of a control rod from the core, which would cause a rapid increase of the reactivity and the thermal
energy in the fuel (Meyer et al., 1986). The increase in fuel temperature resulting from an RIA induces a rapid fuel expansion, causing a severe pellet-cladding mechanical interaction (PCMI). This PCMI forces the cladding to experience a multiaxial tension such that the maximum principal strain is in the hoop (i.e., transverse) direction of the cladding tube. The survivability of a fuel cladding irradiated to a high burn-up under postulated RIA conditions is thus a response to a combination of the mechanics of a loading and the material degradation during a reactor operation. While such data is available for the axial deformation behavior of cladding tubes, relatively little has been reported in the open literature on the uniaxial tension behavior in the hoop direction of Zircaloy-4 cladding. Accordingly, it is essential to investigate the uniaxial tension behavior in the hoop direction of high burn-up Zircaloy-4 cladding. In this study, ring tensile tests are applied to obtain the data regarding the uniaxial hoop direction deformation behavior. 2. Experimental procedure
∗ Corresponding author. Tel.: +82 42 868 8661; fax: +82 42 863 0565. E-mail address: [email protected] (S.-K. Kim). 0029-5493/$ – see front matter © 2008 Elsevier B.V. All rights reserved. doi:10.1016/j.nucengdes.2008.10.024
The ring tensile specimen used in this study is a Westinghouse 17 × 17 type (Vantage-5H) Zircaloy-4 cladding irradiated for 3 cycles
S.-K. Kim et al. / Nuclear Engineering and Design 239 (2009) 254–260
255
Fig. 1. A schematic of Zircaloy-4 fuel rod irradiated in Ulchin Unit 2.
in the Ulchin Unit 2 pressurized water reactor, whose average burnup was estimated to be approximately 65 GWd/tU (see Fig. 1). The irradiated fuel rod was transferred to PIEF at KAERI, cut into approximately 13 cm length segments with a diamond low speed saw, and then the UO2 pellet inside the fuel rod segment was removed by a mechanical grinding with a drill-attached defueling machine in a hot cell in IMEF. Fig. 2 shows a schematic illustration of the procedure of ring tensile test. The defueling machine is shown in Fig. 3. The specimens for the mechanical properties evaluation were fabricated with a diamond wheel grinder from the defueled cladding segments (see Fig. 4). The dimensions and shape of the ring tensile specimen were designed in order to ensure that any deformation is limited to the gage section of the specimen, so that the uniform uniaxial hoop strain in the gage section could be at its maximal. The dimensions and a photograph of the ring tensile specimen are shown in Figs. 5 and 6, respectively. The gage sections of the specimens were oriented at the top and bottom of the half cylinder of the grip, such that a constant specimen curvature can be maintained during a deformation. A photograph of the jig and the grip is shown in Fig. 7. The interface was lubricated with a graphite-containing vacuum grease lubricant at the beginning of each test to minimize a loss of the applied load. The ring tensile tests were performed in a hot cell with the Instron Servohydraulic System, Model 8562. The tests were performed at 25, 135, 200, 300, 350, 400, 600, and 800 ◦ C, and the initial strain rate was maintained at 0.01/s.
Fig. 3. Photograph of the defueling machine.
The hydride morphologies were observed, which are shown in Fig. 8. As seen in the figure, a hydride rim was formed around the metal substrate/oxide interface. This hydride rim is believed to cause a decrease of the ductility of the cladding tube (Daum et al., 2002).
Fig. 2. A schematic illustration for the procedure of ring tensile tests.
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Fig. 4. Photograph of the machine for the ring tensile specimen fabrication.
Fig. 7. Photograph of the jig and the grip for the ring tensile specimen.
Fig. 5. Dimension of the ring tensile specimen.
3. Results and discussion 3.1. Effects of the lubricant and the gage section position The friction between the inner surface of the ring specimen and the half-cylinder die inserts causes a loss of the applied load. Accordingly, it is desirable to diminish the loss of the applied load by
Fig. 6. Photograph of the ring tensile specimen.
a friction. In this study, the effects of a lubrication were investigated to determine the lubricant for this ring tensile test. The ring tensile tests were performed with Teflon tape and graphite vacuum grease as the lubricant for as-received Zircaloy-4. The load–displacement correlation is shown in Fig. 9. The load–displacement behavior and elastic modulus of the Teflon tape in the elastic region coincides with those of the graphite vacuum grease. On the contrary, the load–displacement behavior is not in accordance with each other in the plastic region. The ultimate tensile strength for the Teflon tape was approximately 2.1 kN, while that of the graphite vacuum grease was approximately 1.8 kN. This difference in the ultimate tensile strength is caused by the difference in their friction coefficients. From a point of view regarding a friction, Teflon tape is more desirable than graphite vacuum grease. But, the maximum allowable temperature of the Teflon tape is approximately 260 ◦ C, so graphite vacuum grease whose maximum allowable temperature is approximately 1000 ◦ C was adopted as a lubricant in this experiment. The load–displacement behavior as a positioning direction of the gauge section was also investigated. As shown in Fig. 10, the ultimate tensile strength and displacement at the fracture showed a different behavior in the up and down positioning direction and the left and right positioning direction. The ultimate tensile strength for the up and down positioning direction was evaluated to be a little higher than that for the left and right positioning direction. And the displacement at fracture direction was evaluated to be a little higher than that for the left and right positioning direction too. When the gauge section of the specimen is located towards the left and right positioning direction in the two half-cylinder inserts system, the elongation of the gauge section may be flattened because there is no circumferential elongation by the central dog-bone insert. Because the specimen grip part in this experiment does not contain the central dog-bone insert, the gauge section of the specimen was located towards the up and down positioning
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257
Fig. 10. Load–displacement behavior with the lubricant type.
To obtain the mechanical strength, the 0.2% offset yield strength (YS) and ultimate tensile strength (UTS) were evaluated, and the
uniform elongation (UE) and total elongation (TE) were also evaluated for the ductility. The hoop stress–strain curves at various temperatures are shown in Fig. 11. The evaluation results of the 0.2% offset YS and the UTS are shown in Figs. 12 and 13. From the figure, it is confirmed that the 0.2% offset YS and the UTS abruptly decrease with an increasing temperature. The UTS was evaluated to be 942.70 MPa at RT, 678.83 MPa at 400 ◦ C, but, it is abruptly diminished to 282.64 MPa at 600 ◦ C, which is achievable in the RIA condition. Especially, it decreases to 58.30 MPa at 800 ◦ C, an extreme condition, which corresponds to 6% of the UTS at RT. This means that the mechanical strength of the high burn-up Zircaloy-4 cladding sharply decreases in the RIA-relevant temperature ranges. The evaluation results of the UE and TE are shown in Figs. 14 and 15. The results show that both the UE and TE increase with an increasing temperature. Especially, they abruptly increase at 600 ◦ C, but become lower beyond this temperature. This peculiar behavior was also observed in the PROMETRA program (Averty et al., 2003) which is a mechanical property relevant test program in conjunction with the CABRI program simulating a RIA. The results of hoop-directional mechanical strength and ductility of irradiated Zircaloy-4 with comparison with PROMETRA database are shown in Figs. 16 and 17, respectively. As shown in the figure, the behavior of this study is consistent with the PROMETRA data. It is believed that this behavior is caused by the elongation minimum phenomenon by
Fig. 9. Load–displacement behavior with the lubricant type.
Fig. 11. Hoop stress–strain curves at various test temperatures.
Fig. 8. Optical microscopy of the Zircaloy-4 from Ulchin Unit 2.
direction to avoid a linear elongation in some regions of the gauge section. 3.2. Evaluation of the mechanical strength and the ductility
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Fig. 12. Yield strength of the un-irradiated and high burn-up fuel cladding.
Fig. 15. Total elongation of the un-irradiated and high burn-up fuel cladding.
Fig. 16. Comparison of mechanical strength with the PROMETRA database.
the dynamic strain aging of the Zircaloy-4 cladding material beyond 600 ◦ C. Fig. 13. Ultimate tensile strength of the un-irradiated and high burn-up fuel cladding.
3.3. Fracture characteristics of the ring tensile tests The stereoscope photographs of the specimens after the ring tensile test are shown in Fig. 18. Fig. 18 represents the fracture patterns of the non-irradiated and high burn-up Zircaloy-4 cladding
Fig. 14. Uniform elongation of the un-irradiated and high burn-up fuel cladding.
Fig. 17. Comparison of mechanical ductility with the PROMETRA database.
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Fig. 18. Stereoscope photograph of the specimens after the ring tensile test.
specimens. As shown in the figure, four fracture patterns such as 45◦ shear type fracture, cup and cone type fracture, cup and cup type fracture and chisel edge type fracture were observed in the non-irradiated cladding.
The fracture type was found to depend strongly on the deformation temperature. At room temperature non-irradiated Zircaloy-4 cladding tends to be fractured by 45◦ shear type fracture or cup and cone type fracture. Cup and cone type fracture was dominant at 135
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Table 1 Fracture types for the ring tensile tests of the Zircaloy-4 fuel cladding.
4. Conclusions
Temperature
On the basis of the ring tensile tests for the high burn-up Zircalay-4 cladding from Ulchin Unit 2 and the as-received nonirradiated Zircalay-4 cladding, the following conclusions were drawn. Firstly, the mechanical properties are abruptly degraded beyond 600 ◦ C, which corresponds to a design-basis accident condition such as a RIA. Secondly, the un-irradiated fuel cladding showed ductile fracture behaviors such as 45◦ shear type fracture, cup and cone type fracture, cup and cup type fracture and chisel edge type fracture, while the high burn-up Zircalay-4 cladding showed a brittle fracture behavior even at the high temperatures (e.g. over 600 ◦ C) which are achievable during a RIA. Thirdly, in the case of the high burn-up Zircalay-4 cladding, the strength, ductility and the energy to break are strongly dependent on the material property itself which are degraded by oxidation and hydriding during an operation rather than the temperature. Fourthly, the hydride rim formation in the vicinity of the metal-oxide interface can play an important role in the degradation of the mechanical properties for the high burn-up fuel cladding.
Fracture type Non-irradiated cladding
Room temperature 135 ◦ C 200 ◦ C 300 ◦ C 350 ◦ C 400 ◦ C 600 ◦ C 800 ◦ C
45◦ shear type fracture, cup and cone type fracture Cup and cone type fracture Cup and cone type fracture Cup and cone type fracture, cup and cup type fracture Cup and cup type fracture Cup and cup type fracture Chisel edge type fracture Unidentified
High burn-up cladding
Brittle fracture
and 200 ◦ C. Both cup and cone type fracture and cup and cup type fracture was observed at 300 ◦ C. Only cup and cup type fracture was observed at 350 and 400 ◦ C. At 600 ◦ C chisel edge type fracture was observed unlike at lower temperatures with few fracture surface area. At 800 ◦ C the fracture type is unclear. These fracture types are summarized in Table 1. In conclusion, the fracture type of non-irradiated Zircaloy-4 cladding depends on the temperature, and the behavior is governed by a ductile fracture with a necking. On the contrary, the fracture patterns of the high burn-up Zircaloy-4 cladding showed completely different fracture patterns from the non-irradiated Zircaloy-4 cladding. The fracture type was observed to be vertical in the tensile direction without a necking at all of the test temperatures, which is convincing evidence of the brittle fracture behavior of high burn-up fuel cladding regardless of temperature. This means that even at a high temperature, 600 or 800 ◦ C, the fracture pattern showed a brittle fracture behavior. Accordingly, it was found that the high burn-up Zircaloy-4 cladding becomes very brittle even at the high temperatures achievable during a designbasis accident (Daum et al., 2002).
Acknowledgment This work was financially supported by the Ministry of Education, Science and Technology in Korea. References Meyer, R.O., et al., 1986. A regulatory assessment of test data for reactivity initiated accidents. Nuclear Safety 37 (4), 271–288. Averty, X., et al., 2003. Tensile tests on ring specimens machined in M5 cladding irradiated 6 cycles. IRSN 2003/50. Daum, R.S., et al., 2002. On the embrittlement of Zircaloy-4 under RIA-relevant conditions. In: Zirconium in the nuclear industry: thirteenth international symposium, ASTM STP 1423, PA, pp. 702–719.