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A study on the grid-to-rod fretting wear-induced fuel failure observed in the 16×16KOFA fuel
Nuclear Engineering and Design 240 (2010) 756–762
Contents lists available at ScienceDirect
Nuclear Engineering and Design journal homepage: www.elsevier.com/locate/nucengdes
A study on the grid-to-rod fretting wear-induced fuel failure observed in the 16×16KOFA fuel Kyu-Tae Kim ∗ Dongguk University, College of Energy & Environment, 707 Seokjang-Dong, Gyeongju, Gyeongbuk 780-714, Republic of Korea
a r t i c l e
i n f o
Article history: Received 10 September 2009 Received in revised form 13 November 2009 Accepted 10 December 2009
a b s t r a c t The burnup-dependent grid-to-rod gap combined with the fluid-induced vibration may generate grid-torod fretting wear-induced fuel failure for some fuel assemblies in a certain burnup range. The systematic grid-to-rod fretting wear-induced fuel failure occurred at the 16×16 Korean Optimized Fuel Assembly loaded in the 2-loop Westinghouse type plant in Korea. Prior to various tests and some measurements for investigating its root causes, they were assumed to be self-excited fuel assembly vibration caused by hydraulic-unbalanced mixing vane design, excessive cross-flow between fuel assemblies during the transition core, or relatively large grid-to-rod gap formation during in-reactor irradiation that may be caused by excessive initial spring force loss of fresh fuel during a fuel rod loading process and/or a fuel assembly transport to a plant and by excessive cladding creep-down. A wide spectrum of tests and some measurements were performed to find out root cause(s) of the grid-to-rod fretting wear-induced fuel failure. Based on these tests and measurements, it is concluded that the self-excited fuel assembly vibration is the primary root cause, while excessive initial spring force loss during the fuel rod loading process is the second major root cause. © 2009 Elsevier B.V. All rights reserved.
1. Introduction Nowadays a nuclear renaissance is under way due to global climate problem and limited energy source of fossil fuel. Considering that public perception of nuclear power as safe, efficient and reliable is more important than ever, nuclear fuel reliability is essential to sustain the nuclear renaissance. In the world various nuclear fuels have been designed and developed to be more competitive, but some fuels have had brief commercial lives because of designrelated systematic fuel failure. In the 1970s through 1980s, Korea had imported nuclear fuel assemblies since Korea did not have any fuel design and fabrication technology. Korea Nuclear Fuel (KNF) was established in 1982 to localize nuclear fuel technology, based on various technologies transferred from the former Siemens/KWU, the former Combustion Engineering, and Westinghouse Electric Company. KNF started to supply firstly localized PWR and PHWR fuels from 1989 and 1997, respectively. The firstly localized PWR fuels supplied by KNF are 14×14, 16×16 and 17×17 Korean Optimized Fuel Assemblies (KOFAs).
Abbreviations: KNF, Korea Nuclear Fuel; PWR, pressurized water reactor; PHWR, pressurized heavy water reactor; KOFA, Korean Optimized Fuel Assembly; GTRFW, grid-to-rod fretting wear; FIV, fluid-induced vibration; FR, fuel rod; FA, fuel assembly. ∗ Tel.: +82 11 9805 1447; fax: +82 54 770 2873. E-mail address: [email protected]. 0029-5493/$ – see front matter © 2009 Elsevier B.V. All rights reserved. doi:10.1016/j.nucengdes.2009.12.014
During the transition core with a mixed fuel loading having the imported fuel assembly and the 16×16KOFA in the 2-loop Westinghouse type plant in Korea, however, the systematic grid-to-rod fretting wear-induced fuel failure occurred at the 16×16KOFA (Kim, 1993), as shown in Fig. 1. The key reactor core operating conditions of the 2-loop plant and the 16×16KOFA design parameters are given in Table 1. From Fig. 1, it can be seen that the first leak signal was detected at a twice-burned fuel assembly (FA) at about 90 days after startup (∼22,000 MWD/MTU). Based on the ultrasonic tests, almost one hundred KOFA fuel rods (FRs) were found to be leaking due to the grid-to-rod fretting wear (GTRFW). The leak rods are located generally around the periphery of FA, as shown in Fig. 2. Fig. 3 shows perforated fretting wear configurations of failed KOFA fuel rods. From this figure, the spring and dimple marks on the fuel rod surfaces are clearly identified. To investigate root cause(s) of 16×16KOFA fuel failures, other fuel designs generating fretting wear-induced fuel failures (Kennard et al., 1995; Donovan, 2007) were reviewed in detail. The fretting wear-induced fuel failures over the world have occurred with various fuel designs that include 17×17 Vantage5H of Westinghouse Electric Company, 18×18 Konvoi of the former Siemens and 16×16 Guardian of the former Combustion Engineering, etc. It is reported that the GTRWF would occur mainly at the mid-grid positions and the first leak signals appear in a wide range of operating time covering the 1st through 3rd cycle (Kennard et al., 1995). Based on this evaluation of root causes of the GTRFW occurring in the world, probable root causes of the GTRFW observed in the
K.-T. Kim / Nuclear Engineering and Design 240 (2010) 756–762
Fig. 1. The ratio of Xe-133/Xe-135 versus reactor operation time for the 16×16KOFA. Table 1 Reactor core and fuel assembly design parameters for the 16×16KOFA. Design parameters
Values
Inlet coolant temperature Outlet coolant temperature Pressure Coolant flow rate Average LHGR Purification flow rate Fuel rod diameter Fuel rod pitch Number of spacer grids (materials) Mid-grid span length Grid-to-rod contact shape Initial spring force requirement Reload cycle length
287.7 ◦ C 324.7 ◦ C 155.1 bar 4.80 m/s 175.7 W/cm 75 GPM 9.50 mm 12.32 mm 8 (Inconel) 522 mm Point contact >12 N 360 days
16×16KOFA fuel were considered to be self-excited FA vibration caused by hydraulically unbalanced mixing vane design (Jang and Lu, 2001; Conner, 2001), excessive cross-flow during the transition core, or relatively large grid-to-rod gap formation during in-reactor irradiation that may be caused by excessive spring force loss of fresh
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Fig. 3. Perforated grid-to-rod fretting wear configurations of failed 16×16KOFA fuel rods.
fuel during a fuel rod (FR) loading process and/or a FA transport to the plant and by excessive cladding creep-down. On the other hand, the GTRFW rate models for PWR fuel were proposed to predict grid-to-rod fretting wear rates for various spacer grid designs (Kim, 2009). In addition, the fluid-induced vibration characteristics of a fuel rod supported by spacer grids were investigated to understand vibration behaviors within the reactor due to coolant flow (Choi et al., 2003). In this study, therefore, a wide spectrum of tests and some measurements that include FR loading tests, FA transport tests, FA vibration tests and cladding creep-down measurements were performed to find out root cause(s) of the GTRFW-induced fuel failure observed in the 16×16KOFA. Besides these tests and measurements, the extent of cross-flow between FAs was reviewed to evaluate the impact of the cross-flow on the GTRFW of 16×16KOFA. 2. Tests for verifying root cause(s) The GTRFW-induced fuel failure might occur due to FA designrelated excessive FA vibration usually coupled with insufficient grid-to-rod supporting conditions such as relatively a large gridto-rod gap size and a small grid-to-rod contact. In addition, the FA vibration might be accelerated by additional vibration sources originated from reactor core designs that include upper and lower reactor internal configurations, FA-to-baffle gap size, FA-to-FA gap size, reactor coolant axial velocity, etc. (Kim and Suh, 2009). Considering that the imported 16×16 FA has survived the GTRFW-induced fuel failure in the 2-loop Westinghouse type plant in Korea but the 16×16KOFA fuel did not survive in that reactor, however, the reactor core design-induced vibration may not be large enough to generate the GTRFW-induced fuel failure regardless of FA designs. In this study, therefore, a wide spectrum of tests and some measurements were performed only to evaluate the impact of the 16×16KOFA design-related excessive FA vibration and the insufficient grid-to-rod supporting conditions on the GTRFW-induced fuel failure. The tests and measurements performed in this work are summarized in the following: - FA vibration tests versus flow rate (flow sweep tests). - Initial spring force loss measurements during FR loading process. - Additional spring force loss measurements during FA transport to the plant. - Cladding creep-down measurements for two kinds of cladding materials.
Fig. 2. Locations of grid-to-rod fretting wear-induced failed FRs in the 16×16KOFA (Kim and Suh, 2009).
The extent of FA vibration versus flow rate may indicate whether or not the 16×16KOFA will generate the self-excited FA vibration
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K.-T. Kim / Nuclear Engineering and Design 240 (2010) 756–762
Fig. 4. Fuel assembly vibration versus reactor core flow rate.
Fig. 5. Initial spring force distributions of the 16×16KOFA before and after the FR loading with a FR loading speed of 0.33 m/s.
in the operating flow range, while the data obtained from the other three tests may show whether or not excessive grid-to-rod gaps will form in a certain burnup range. It is noteworthy that the FR vibration caused by the grid-to-rod gap formation will become more vigorous when combined with the FA vibration. On the other hand, the extent of cross-flow between FAs was reviewed to evaluate the impact of the cross-flow on the GTRFW of 16×16KOFA.
2.1. FA vibration tests versus flow rate It is known that the self-excited FA vibration is generated by asymmetric mixing vane pattern across the spacer grid assembly, which generates rotational force on the FA. In order to investigate the impact of FA design on the FA vibration, flow sweep tests were performed for the 16×16KOFA and the PLUS7 fuel, respectively. The PLUS7 fuel has been used at the OPR1000s operating in Korea and has not generated any GTRFW-induced fuel failure yet (Kim and Suh, 2008; Kim et al., 2008). Fig. 4 shows the extent of FA vibration versus reactor core flow rate. From this figure, it can be seen that the 16×16KOFA generates a sharp increase in FA vibration in the operating flow range, i.e., self-excited FA vibration, whereas the PLUS7 fuel generates a random FA vibration only, as expected from the in-reactor operation experience.
2.2. Initial spring force loss measurements during FR loading process In order to investigate the impact of the FR loading speed on the initial spring force loss during the FR loading into the FA skeleton, the initial spring forces of each spacer grid cell were measured before and after the FR loading with loading speeds of 0.18 and 33 m/s, respectively. The FR loading speed of 0.33 m/s was employed for fabricating the 16×16KOFA generating the GTRFWinduced fuel failure, while the FR loading speed of 0.18 m/s was tested to evaluate the impact of FR loading speed on the initial spring force loss. Figs. 5 and 6 show the initial spring force distribution before and after the FR loading with the loading speeds of 0.33 and 0.18 m/s, respectively. From these figures, it can be seen that with the FR loading speed of 0.33 m/s the initial spring forces of some spacer grid cells are less than 12 N, the minimum initial spring force requirement for preventing the GTRFW, while with the FR loading speed of 0.18 m/s the initial spring forces of all spacer grid cells is greater than 12 N.
Fig. 6. Initial spring force distributions of the 16×16KOFA before and after the FR loading with a FR loading speed of 0.18 m/s.
2.3. Additional spring force loss measurements during FA transport to the plant In order to find out the additional spring force loss after FR loading during the FA transport to the 2-loop Westinghouse type plant in Korea, the spring forces of each spacer grid cell were measured after the FA transport to the Kori-2 unit. Table 2 shows the additional spring force loss during the FA transport to the Kori-2 unit. Table 2 Additional initial spring force loss during the 16×16KOFA transport to the plant. Sample number
Spring force loss (N)
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16
0.0 0.7 0.5 0.4 0.5 0.1 0.3 0.4 0.0 0.0 0.7 0.2 0.0 0.1 0.1 0.2
Mean value
0.263
Standard deviation
0.245
K.-T. Kim / Nuclear Engineering and Design 240 (2010) 756–762
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Table 3 Cross-flow data in transition cores of various plants. Plant
Difference in pressure loss coefficient (%)
Maximum cross-flow (m/s)
2-loop 14×14 W type 2-loop 16×16 W type (16×16KOFA loaded) 3-loop 17×17 W type Angra-1
46 40 41 –
0.38 0.30 0.27 0.36
there is quite a large difference in the unirradiated thermal cladding creep-down between them. 2.5. Evaluation of cross-flow Fig. 7. Out-of-pile thermal creep behavior of cladding materials A and B (T = 400 ◦ C; P = 80 MPa).
These additional spring force loss data were measured by the following three steps: Firstly, spring force loss of each spacer grid for two FAs during the fuel rod loading process was obtained by measuring before and after the FR loading, as explained in Section 2.2. It should be noted that one FA generates eight sets of spring force data from eight spacer grid assemblies. Therefore, 16 sets of residual spring force data just before the FA transportation to the Kori-2 unit. Secondly, two FAs were transported from the manufacturing facility to the Kori-2 unit and then came back to the manufacturing facility. Finally, data of additional spring force occurring during the transportation for 16 spacer grid assemblies were obtained by measuring residual spring forces after unloading FRs from each spacer grid assembly. From Table 2, it can be seen that the additional spring force loss occurred during the FA transport is considered very small, comparing with the initial spring force loss during the FR loading process. 2.4. Cladding creep-down measurements It is known that the larger creep-down generates the larger gridto-rod gap, which will subsequently cause the FR vibration to be more vigorous. In order to investigate the impact of cladding materials on the cladding creep-down, cladding creep-down data were measured for two kinds of Zry-4 cladding materials under unirradiated and irradiated conditions. Fig. 7 shows thermal cladding creep data under unirradiated conditions and Fig. 8 shows the axial cladding creep-down profiles for twice-burned FAs with cladding materials A and B, respectively. From these figures, it can be seen that there exists only a slight difference in the irradiated cladding creep-down data between cladding materials A and B, even though
Fig. 8. Axial profile of cladding creep-down for twice-burned KOFA fuel (Bu = 32,000 MWD/MTU).
Table 3 summarizes differences in pressure loss coefficient and cross-flow velocity during transition cores for various plants. The 2-loop Westinghouse type plant experiencing the GTRFW-induced fuel failure is generating a comparable pressure loss coefficient and cross-flow velocity, comparing with the other plants in Table 3. It is concluded that the cross-flow acting on the 16×16KOFA during the transition core is not large enough to generate excessive fuel rod vibration. 3. Results and discussion Fig. 4 shows the FA vibration test results that were generated by the flow sweep tests with a full-scale fuel assembly. From this figure, it is found that the 16×16KOFA generates very sharp increase in the FA vibration in the operating range, which is called as self-excited FA vibration, while the PLUS7 FA generates a random FA vibration. In detail, the 16×16KOFA generates self-excited FA vibration with the peak amplitude of about 20 × 10−5 m with the frequency of between 1 and 20 Hz in its operating flow range. This self-excited FA vibration is caused by hydraulic-unbalanced mixing vane pattern across the spacer grid assembly, which can be supported by nearly hydraulic-balanced mixing vane design of PLUS7 showing no GTRFW (Kim et al., 2004; Kim and Kim, 2002). In other words, the PLUS7 FA with nearly hydraulic-balanced mixing vane pattern (see Fig. 9) generates a random FA vibration only with the peak amplitude of about 3 × 10−5 m with the frequency of below 5 Hz in its operating range. It should be noted that the mixing vane pattern across the spacer grid assembly for the 16×16KOFA fuel does not have a perfect symmetry, i.e., perfectly hydraulicbalanced mixing vanes, since the mixing vane pattern at the center
Fig. 9. Flow paths generated by the hydraulic-balanced mixing vane pattern of the PLUS7 fuel (Kim and Suh, 2009).
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Table 4 Statistical information on initial spring force of the 16×16KOFA. Phase
Mean spring force (standard deviation)
Mean spring force loss (standard deviation)
One-sided 95/95% lower limit
Initial spring force before FR loading Spring force loss during FR loading Spring force loss during FA transport
16.750 N (0.840 N) 14.405 N 14.142 N
– 2.345 N (1.972 N) 0.263 N (0.245 N)
14.630 N 9.430 N 8.700 N
grid strip for the 16×16 geometry cannot be made to be perfectly hydraulic-balanced and furthermore orientations of mixing vanes around the guide thimble tubes were optimized to generate better mixing. Therefore, the flow paths shown in Fig. 9 are not perfectly symmetric. The flow directions drawn in Fig. 9 were determined by orientations of mixing vanes of each spacer grid strip. Naturally, one can say that the GTRFW-induced fuel failure of the 16×16KOFA is caused by the self-excited FA vibration with hydraulic-unbalanced mixing vane pattern across the spacer grid assembly. It is reported that the V5H FA also has experienced the GTRFW-induced fuel failure in the 3-loop Westinghouse type plants in the U.S. because of the self-excited FA vibration with hydraulic-unbalanced mixing vane pattern across the spacer grid assembly (Washington, 1993). However, it is noteworthy that the V5H FA without mixing vanes still generated the GTRFW-induced fuel failure in the Fort Calhoun plant in the U.S., which had not been expected from the viewpoints of the self-excited FA vibration. Therefore, it may be said that the GTRFW-induced fuel failure can occur partly due to the insufficient grid-to-rod supporting conditions coupled with a random FA vibration, even though there is no self-excited FA vibration. In order to investigate the impact of the 16×16KOFA design and a certain manufacturing process on the grid-to-rod supporting conditions, the various tests were performed such as initial spring force loss measurements during the FR loading process, additional spring force loss measurements during FA transport to the plant and cladding creep-down measurements for two kinds of cladding materials. From Figs. 5 and 6, it is found that with the FR loading speed of 0.33 m/s the initial spring forces of some spacer grid cells are less than 12 N, the minimum initial spring force requirement for preventing the GTRFW, while with the FR loading speed of 0.18 m/s the initial spring forces of all spacer grid cells is greater than 12 N. The statistical values for the spring force loss during the FR loading process and during the FA transport are summarized in Table 4. From this table, it can be seen that the mean value and standard deviation of the spring force loss during the FR loading process are 2.345 and 1.972 N, respectively, while those of the additional spring force loss during the FA transport are 0.263 and 0.245 N, respectively. From Fig. 8, on the other hand, it can be seen that there exists only a slight difference in the in-reactor cladding creep-down between cladding materials A and B, even though there is quite a large difference in the cladding thermal creep-down data between them (see Fig. 7). Considering that irradiation-induced cladding creep is much more dominant in the reactor operation than thermal-induced one, therefore, it is concluded that the cladding materials have no significant impact on the unexpected large grid-to-rod gap, if it would occur at the 16×16KOFA grid cells. Based on the statistical values of the initial spring force loss given in Table 4 and Fig. 5, three kinds of initial spring force are considered to have 8, 12 and 18 N to investigate the impact of initial spring force on the onset time of grid-to-gap appearance. The values of 18 and 12 N represent the mean initial spring force and the minimum one required in the design specification, respectively, while that of 8 N represents the one-sided 95/95% lower limit measured for fuel assemblies manufactured by the FR loading process with 0.33 m/s and transported to the plant. The GRIDFORCE program (Kim and Kim, 1997) was used to predict residual elastic spring deflection and force with relevant fuel
design data as a function of burnup. A calculation method used in the GRIDFORCE program can be briefly summarized as follows; A residual spring force at t = t at an operating temperature can be given by the Hooke’s law. fres (t) = COT ıres (t)
(1)
where fres (t) is the residual spring force at t = t, COT the spring constant at an operating temperature, and ıres (t) the residual elastic spring deflection at t = t. Then, the residual elastic deflection at t = t, ıres (t), can be derived by the following formula: ıres (t) = ıo − ıP (t) − ıT (t) − ıCR (t) − ıIR (t) − ırex (t)
(2)
where ıo is the initial elastic spring deflection at room temperature, ıP (t) the accumulated elastic cladding deflection caused by the coolant overpressure from t = 0 to t = t, ıT (t) the accumulated thermal expansion difference between spacer grid and cladding from t = 0 to t = t, ıCR (t) the accumulated cladding creep-down from t = 0 to t = t, ıIR (t) the accumulated spacer grid irradiation growth from t = 0 to t = t, and ırex (t) is the accumulated elastic spring deflection loss due to spring force relaxation from t = 0 to t = t. The value of ıo is determined by unstrained spring height, fuel rod pitch and initial cladding diameter, while the values of ıP (t), ıT (t), ıCR (t), ıIR (t) and ırex (t) are calculated with the use of reactor operating conditions and relevant fuel design data. A schematic diagram of elastic spring deflection as a function of burnup may be drawn, as shown in Fig. 10. As an illustration for grid-to-rod gap formation, changes in the residual elastic spring deflections, which were calculated by the GRIDFORCE program, are shown in Figs. 11 and 12. From these figures, it can be said that the initial spring force of 12 N does not generate a grid-to-rod gap at any spacer grids at the FA burnup of 22 MWD/kgU, whereas that of 8 N generates grid-to-rod gaps at the 3rd, 4th, 5th and 6th spacer grids. With the use of Eq. (1), changes in the residual spring forces were calculated for initial spring forces of 8, 12 and 18 N as a function of burnup, as shown in Fig. 13. From this figure, it is found that the grid-to-rod gap starts to form at 18, 24 and 28 MWD/kgU for the initial spring forces of 8, 12 and 18 N, respectively. It is obvious
Fig. 10. A schematic diagram of elastic spring deflection as a function of burnup.
K.-T. Kim / Nuclear Engineering and Design 240 (2010) 756–762
Fig. 11. Residual spring deflection at FA average burnup of 22 MWD/kgU for an initial spring force of 12 N.
Fig. 12. Residual spring deflection at FA average burnup of 22 MWD/kgU for an initial spring force of 8 N.
that the lower initial spring force generates the larger grid-to-rod gap at the same burnup or the earlier grid-to-rod gap. Therefore, one can say that some grid cells with the initial spring force less than 12 N, the minimum initial spring force required in the design specification, may be susceptible to the GTRFW-induced fuel fail-
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ure. Furthermore, the probability of the GTRFW-induced fuel failure goes up sharply when combined with the self-excited FA vibration. Considering that the initial spring loss of the 16×61KOFA is mainly due to the FR loading process with 0.33 m/s, however, the impact of the additional initial spring force loss and the cladding creep-down on the unexpected early grid-to-rod gap formation is considered relatively small. On the other hand, it should be noted that the initial spring force loss during the FR loading process can be reduced drastically if the loading speed is changed from 0.33 to 0.18 m/s, as shown in Fig. 6. Accordingly, the GTRFW-induced fuel failure could be eliminated if one would develop the 16×16KOFA with hydraulic-balanced mixing vanes and adopt the FR loading speed of less than 0.18 m/s. Based on the wide spectrum of test results obtained in this study, guidelines against the GTRFW-induced fuel failure are proposed as follows. Firstly, to minimize FA-induced vibration, one should develop spacer grid design with hydraulic-balanced mixing vane pattern across the spacer grid assembly. The spacer grid with mixing vanes considered needs to be verified against the selfexcited fuel assembly-induced vibration through the flow sweep tests of the full-scale fuel assembly covering operating range of nuclear power plants considered. Secondly, to minimize the gridto-rod fretting wear caused by insufficient grid-to-rod supporting conditions, one should maximize the grid-to-rod contact area but minimize the grid-to-rod gap size during the fuel lifetime. It should be kept in mind that the lower initial spring loss during the FR loading process, the smaller cladding creep-down and the smaller spacer grid width growth will generate the smaller grid-to-rod gap. Finally, with the use of the full-scale FA containing oxidized and non-oxidized FRs with various grid-to-rod gap sizes, the FA endurance tests need to be performed to verify the effect of the FR supporting conditions and the FA vibration on the fretting wear rate (Lu, 2001). 4. Conclusions A wide spectrum of various tests and some measurements were performed to investigate root causes of the grid-to-rod fretting wear-induced fuel failure observed in the 16×16 Korean Optimized Fuel Assemblies. The primary root cause is considered to be the selfexcited fuel assembly vibration occurring in the operating range with the peak amplitude of about 20 × 10−5 m with the frequency of between 1 and 20 Hz. The second effective root cause might be the initial spring force loss during the fuel rod loading process with 0.33 m/s loading speed. The grid-to-rod gap appears at much lower burnup with the one-sided 95% lower limit of initial spring force than expected with minimum initial spring force required in the design specification. On the other hand, guidelines against the GTRFW-induced fuel failure are proposed, indicating that the GTRFW-induced fuel failure might be eliminated if one develops the 16×16KOFA with hydraulically balanced mixing vanes and adopts the FR loading speed of less than 0.18 m/s. Acknowledgements This research was supported by Basic Atomic Energy Research Institute Program through the National Research Foundation of Korea (NRF) funded by the Ministry of Education, Science and Technology (No. 2009-0075907). References
Fig. 13. Residual spring forces as a function of FA average burnup.
Choi, N., Kang, H., Yoon, K., Song, K., Jung, Y., 2003. Vibration analysis of a dummy fuel rod continuously supported by spacer grids. In: Transactions of the 17th International Conference on Structural Mechanics in Reactor Technology (SMiRT 17), Prague, Czech Republic, August 17–22, 2003.
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Conner, M., 2001. VISTA High Frequency Vibration Test Report-16×16 KAFD midgrid designs, KAFD-01D-77. Donovan, K., 2007. Fuel integrity initiative overview. In: ANS LWR Fuel Performance Conference, San Francisco, USA, September 30–October 3, 2007. Jang, Y., Lu, R., 2001. FACTS Fuel Assembly Vibration Test Report-16×16 KAFD, KAFD01D-116. Kennard, M.W., Sunderland, D.J., Harbottle, J.E., 1995. A Study of Grid-to-Rod Fretting Wear in PWR Fuel Assembly, vols. 1 and 2. S.M. Stroller Co. Kim, H., Kim, K., 1997. Development of a methodology for in-reactor supporting condition prediction. In: The 14th International Conference on Structural Mechanics in Reactor Technology, Lyon, France, August 17–22, pp. 105–112. Kim, K., 1993. Fuel Failure Analysis of Kori-2 Cycle 8, KAERI/TR-339/93. Kim, K., 2009. The study on grid-to-rod fretting wear models for PWR fuel. Nuclear Eng. Design 239, 2820–2824. Kim, K., Jang, Y., Choi, J., Lee, S., 2004. A study of flow-induced grid-to-rod fretting wear in PWR fuel assemblies. In: Proceedings of ASME PVP 2004, San Diego, CA, USA, July 25–29.
Kim, K., Jang, Y., Kim, J., 2008. In-reactor performance of an advanced PWR fuel, PLUS7, for OPR1000s in Korea. J. Nuclear Sci. Tech. 45, 836–849. Kim, K., Suh, J., 2008. Development of an advanced PWR fuel for OPR1000s in Korea. Nuclear Eng. Design 238, 2606–2613. Kim, K., Suh, J., 2009. Impact of nuclear fuel assembly design on grid-to-rod fretting wear. J. Nuclear Sci. Tech. 46, 149–157. Kim, Y., Kim, K., 2002. Advanced spacer grid design for the PLUS7 fuel assembly. In: The 3rd Korea-Japan Symposium on Nuclear Thermal Hydraulics and Safety, Gyeongju, Korea, pp. 493–499. Lu, Y., 2001. VIPER 16×16 KAFD Fuel Assembly and KSNP Fuel Assembly Long-term Wear Test Report, KAFD-01D-115. Washington, 1993. Fuel failures at two plants force Westinghouse to redesign VANTAGE5H. Nuclear Fuel, 18.
Contents lists available at ScienceDirect
Nuclear Engineering and Design journal homepage: www.elsevier.com/locate/nucengdes
A study on the grid-to-rod fretting wear-induced fuel failure observed in the 16×16KOFA fuel Kyu-Tae Kim ∗ Dongguk University, College of Energy & Environment, 707 Seokjang-Dong, Gyeongju, Gyeongbuk 780-714, Republic of Korea
a r t i c l e
i n f o
Article history: Received 10 September 2009 Received in revised form 13 November 2009 Accepted 10 December 2009
a b s t r a c t The burnup-dependent grid-to-rod gap combined with the fluid-induced vibration may generate grid-torod fretting wear-induced fuel failure for some fuel assemblies in a certain burnup range. The systematic grid-to-rod fretting wear-induced fuel failure occurred at the 16×16 Korean Optimized Fuel Assembly loaded in the 2-loop Westinghouse type plant in Korea. Prior to various tests and some measurements for investigating its root causes, they were assumed to be self-excited fuel assembly vibration caused by hydraulic-unbalanced mixing vane design, excessive cross-flow between fuel assemblies during the transition core, or relatively large grid-to-rod gap formation during in-reactor irradiation that may be caused by excessive initial spring force loss of fresh fuel during a fuel rod loading process and/or a fuel assembly transport to a plant and by excessive cladding creep-down. A wide spectrum of tests and some measurements were performed to find out root cause(s) of the grid-to-rod fretting wear-induced fuel failure. Based on these tests and measurements, it is concluded that the self-excited fuel assembly vibration is the primary root cause, while excessive initial spring force loss during the fuel rod loading process is the second major root cause. © 2009 Elsevier B.V. All rights reserved.
1. Introduction Nowadays a nuclear renaissance is under way due to global climate problem and limited energy source of fossil fuel. Considering that public perception of nuclear power as safe, efficient and reliable is more important than ever, nuclear fuel reliability is essential to sustain the nuclear renaissance. In the world various nuclear fuels have been designed and developed to be more competitive, but some fuels have had brief commercial lives because of designrelated systematic fuel failure. In the 1970s through 1980s, Korea had imported nuclear fuel assemblies since Korea did not have any fuel design and fabrication technology. Korea Nuclear Fuel (KNF) was established in 1982 to localize nuclear fuel technology, based on various technologies transferred from the former Siemens/KWU, the former Combustion Engineering, and Westinghouse Electric Company. KNF started to supply firstly localized PWR and PHWR fuels from 1989 and 1997, respectively. The firstly localized PWR fuels supplied by KNF are 14×14, 16×16 and 17×17 Korean Optimized Fuel Assemblies (KOFAs).
Abbreviations: KNF, Korea Nuclear Fuel; PWR, pressurized water reactor; PHWR, pressurized heavy water reactor; KOFA, Korean Optimized Fuel Assembly; GTRFW, grid-to-rod fretting wear; FIV, fluid-induced vibration; FR, fuel rod; FA, fuel assembly. ∗ Tel.: +82 11 9805 1447; fax: +82 54 770 2873. E-mail address: [email protected]. 0029-5493/$ – see front matter © 2009 Elsevier B.V. All rights reserved. doi:10.1016/j.nucengdes.2009.12.014
During the transition core with a mixed fuel loading having the imported fuel assembly and the 16×16KOFA in the 2-loop Westinghouse type plant in Korea, however, the systematic grid-to-rod fretting wear-induced fuel failure occurred at the 16×16KOFA (Kim, 1993), as shown in Fig. 1. The key reactor core operating conditions of the 2-loop plant and the 16×16KOFA design parameters are given in Table 1. From Fig. 1, it can be seen that the first leak signal was detected at a twice-burned fuel assembly (FA) at about 90 days after startup (∼22,000 MWD/MTU). Based on the ultrasonic tests, almost one hundred KOFA fuel rods (FRs) were found to be leaking due to the grid-to-rod fretting wear (GTRFW). The leak rods are located generally around the periphery of FA, as shown in Fig. 2. Fig. 3 shows perforated fretting wear configurations of failed KOFA fuel rods. From this figure, the spring and dimple marks on the fuel rod surfaces are clearly identified. To investigate root cause(s) of 16×16KOFA fuel failures, other fuel designs generating fretting wear-induced fuel failures (Kennard et al., 1995; Donovan, 2007) were reviewed in detail. The fretting wear-induced fuel failures over the world have occurred with various fuel designs that include 17×17 Vantage5H of Westinghouse Electric Company, 18×18 Konvoi of the former Siemens and 16×16 Guardian of the former Combustion Engineering, etc. It is reported that the GTRWF would occur mainly at the mid-grid positions and the first leak signals appear in a wide range of operating time covering the 1st through 3rd cycle (Kennard et al., 1995). Based on this evaluation of root causes of the GTRFW occurring in the world, probable root causes of the GTRFW observed in the
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Fig. 1. The ratio of Xe-133/Xe-135 versus reactor operation time for the 16×16KOFA. Table 1 Reactor core and fuel assembly design parameters for the 16×16KOFA. Design parameters
Values
Inlet coolant temperature Outlet coolant temperature Pressure Coolant flow rate Average LHGR Purification flow rate Fuel rod diameter Fuel rod pitch Number of spacer grids (materials) Mid-grid span length Grid-to-rod contact shape Initial spring force requirement Reload cycle length
287.7 ◦ C 324.7 ◦ C 155.1 bar 4.80 m/s 175.7 W/cm 75 GPM 9.50 mm 12.32 mm 8 (Inconel) 522 mm Point contact >12 N 360 days
16×16KOFA fuel were considered to be self-excited FA vibration caused by hydraulically unbalanced mixing vane design (Jang and Lu, 2001; Conner, 2001), excessive cross-flow during the transition core, or relatively large grid-to-rod gap formation during in-reactor irradiation that may be caused by excessive spring force loss of fresh
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Fig. 3. Perforated grid-to-rod fretting wear configurations of failed 16×16KOFA fuel rods.
fuel during a fuel rod (FR) loading process and/or a FA transport to the plant and by excessive cladding creep-down. On the other hand, the GTRFW rate models for PWR fuel were proposed to predict grid-to-rod fretting wear rates for various spacer grid designs (Kim, 2009). In addition, the fluid-induced vibration characteristics of a fuel rod supported by spacer grids were investigated to understand vibration behaviors within the reactor due to coolant flow (Choi et al., 2003). In this study, therefore, a wide spectrum of tests and some measurements that include FR loading tests, FA transport tests, FA vibration tests and cladding creep-down measurements were performed to find out root cause(s) of the GTRFW-induced fuel failure observed in the 16×16KOFA. Besides these tests and measurements, the extent of cross-flow between FAs was reviewed to evaluate the impact of the cross-flow on the GTRFW of 16×16KOFA. 2. Tests for verifying root cause(s) The GTRFW-induced fuel failure might occur due to FA designrelated excessive FA vibration usually coupled with insufficient grid-to-rod supporting conditions such as relatively a large gridto-rod gap size and a small grid-to-rod contact. In addition, the FA vibration might be accelerated by additional vibration sources originated from reactor core designs that include upper and lower reactor internal configurations, FA-to-baffle gap size, FA-to-FA gap size, reactor coolant axial velocity, etc. (Kim and Suh, 2009). Considering that the imported 16×16 FA has survived the GTRFW-induced fuel failure in the 2-loop Westinghouse type plant in Korea but the 16×16KOFA fuel did not survive in that reactor, however, the reactor core design-induced vibration may not be large enough to generate the GTRFW-induced fuel failure regardless of FA designs. In this study, therefore, a wide spectrum of tests and some measurements were performed only to evaluate the impact of the 16×16KOFA design-related excessive FA vibration and the insufficient grid-to-rod supporting conditions on the GTRFW-induced fuel failure. The tests and measurements performed in this work are summarized in the following: - FA vibration tests versus flow rate (flow sweep tests). - Initial spring force loss measurements during FR loading process. - Additional spring force loss measurements during FA transport to the plant. - Cladding creep-down measurements for two kinds of cladding materials.
Fig. 2. Locations of grid-to-rod fretting wear-induced failed FRs in the 16×16KOFA (Kim and Suh, 2009).
The extent of FA vibration versus flow rate may indicate whether or not the 16×16KOFA will generate the self-excited FA vibration
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Fig. 4. Fuel assembly vibration versus reactor core flow rate.
Fig. 5. Initial spring force distributions of the 16×16KOFA before and after the FR loading with a FR loading speed of 0.33 m/s.
in the operating flow range, while the data obtained from the other three tests may show whether or not excessive grid-to-rod gaps will form in a certain burnup range. It is noteworthy that the FR vibration caused by the grid-to-rod gap formation will become more vigorous when combined with the FA vibration. On the other hand, the extent of cross-flow between FAs was reviewed to evaluate the impact of the cross-flow on the GTRFW of 16×16KOFA.
2.1. FA vibration tests versus flow rate It is known that the self-excited FA vibration is generated by asymmetric mixing vane pattern across the spacer grid assembly, which generates rotational force on the FA. In order to investigate the impact of FA design on the FA vibration, flow sweep tests were performed for the 16×16KOFA and the PLUS7 fuel, respectively. The PLUS7 fuel has been used at the OPR1000s operating in Korea and has not generated any GTRFW-induced fuel failure yet (Kim and Suh, 2008; Kim et al., 2008). Fig. 4 shows the extent of FA vibration versus reactor core flow rate. From this figure, it can be seen that the 16×16KOFA generates a sharp increase in FA vibration in the operating flow range, i.e., self-excited FA vibration, whereas the PLUS7 fuel generates a random FA vibration only, as expected from the in-reactor operation experience.
2.2. Initial spring force loss measurements during FR loading process In order to investigate the impact of the FR loading speed on the initial spring force loss during the FR loading into the FA skeleton, the initial spring forces of each spacer grid cell were measured before and after the FR loading with loading speeds of 0.18 and 33 m/s, respectively. The FR loading speed of 0.33 m/s was employed for fabricating the 16×16KOFA generating the GTRFWinduced fuel failure, while the FR loading speed of 0.18 m/s was tested to evaluate the impact of FR loading speed on the initial spring force loss. Figs. 5 and 6 show the initial spring force distribution before and after the FR loading with the loading speeds of 0.33 and 0.18 m/s, respectively. From these figures, it can be seen that with the FR loading speed of 0.33 m/s the initial spring forces of some spacer grid cells are less than 12 N, the minimum initial spring force requirement for preventing the GTRFW, while with the FR loading speed of 0.18 m/s the initial spring forces of all spacer grid cells is greater than 12 N.
Fig. 6. Initial spring force distributions of the 16×16KOFA before and after the FR loading with a FR loading speed of 0.18 m/s.
2.3. Additional spring force loss measurements during FA transport to the plant In order to find out the additional spring force loss after FR loading during the FA transport to the 2-loop Westinghouse type plant in Korea, the spring forces of each spacer grid cell were measured after the FA transport to the Kori-2 unit. Table 2 shows the additional spring force loss during the FA transport to the Kori-2 unit. Table 2 Additional initial spring force loss during the 16×16KOFA transport to the plant. Sample number
Spring force loss (N)
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16
0.0 0.7 0.5 0.4 0.5 0.1 0.3 0.4 0.0 0.0 0.7 0.2 0.0 0.1 0.1 0.2
Mean value
0.263
Standard deviation
0.245
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Table 3 Cross-flow data in transition cores of various plants. Plant
Difference in pressure loss coefficient (%)
Maximum cross-flow (m/s)
2-loop 14×14 W type 2-loop 16×16 W type (16×16KOFA loaded) 3-loop 17×17 W type Angra-1
46 40 41 –
0.38 0.30 0.27 0.36
there is quite a large difference in the unirradiated thermal cladding creep-down between them. 2.5. Evaluation of cross-flow Fig. 7. Out-of-pile thermal creep behavior of cladding materials A and B (T = 400 ◦ C; P = 80 MPa).
These additional spring force loss data were measured by the following three steps: Firstly, spring force loss of each spacer grid for two FAs during the fuel rod loading process was obtained by measuring before and after the FR loading, as explained in Section 2.2. It should be noted that one FA generates eight sets of spring force data from eight spacer grid assemblies. Therefore, 16 sets of residual spring force data just before the FA transportation to the Kori-2 unit. Secondly, two FAs were transported from the manufacturing facility to the Kori-2 unit and then came back to the manufacturing facility. Finally, data of additional spring force occurring during the transportation for 16 spacer grid assemblies were obtained by measuring residual spring forces after unloading FRs from each spacer grid assembly. From Table 2, it can be seen that the additional spring force loss occurred during the FA transport is considered very small, comparing with the initial spring force loss during the FR loading process. 2.4. Cladding creep-down measurements It is known that the larger creep-down generates the larger gridto-rod gap, which will subsequently cause the FR vibration to be more vigorous. In order to investigate the impact of cladding materials on the cladding creep-down, cladding creep-down data were measured for two kinds of Zry-4 cladding materials under unirradiated and irradiated conditions. Fig. 7 shows thermal cladding creep data under unirradiated conditions and Fig. 8 shows the axial cladding creep-down profiles for twice-burned FAs with cladding materials A and B, respectively. From these figures, it can be seen that there exists only a slight difference in the irradiated cladding creep-down data between cladding materials A and B, even though
Fig. 8. Axial profile of cladding creep-down for twice-burned KOFA fuel (Bu = 32,000 MWD/MTU).
Table 3 summarizes differences in pressure loss coefficient and cross-flow velocity during transition cores for various plants. The 2-loop Westinghouse type plant experiencing the GTRFW-induced fuel failure is generating a comparable pressure loss coefficient and cross-flow velocity, comparing with the other plants in Table 3. It is concluded that the cross-flow acting on the 16×16KOFA during the transition core is not large enough to generate excessive fuel rod vibration. 3. Results and discussion Fig. 4 shows the FA vibration test results that were generated by the flow sweep tests with a full-scale fuel assembly. From this figure, it is found that the 16×16KOFA generates very sharp increase in the FA vibration in the operating range, which is called as self-excited FA vibration, while the PLUS7 FA generates a random FA vibration. In detail, the 16×16KOFA generates self-excited FA vibration with the peak amplitude of about 20 × 10−5 m with the frequency of between 1 and 20 Hz in its operating flow range. This self-excited FA vibration is caused by hydraulic-unbalanced mixing vane pattern across the spacer grid assembly, which can be supported by nearly hydraulic-balanced mixing vane design of PLUS7 showing no GTRFW (Kim et al., 2004; Kim and Kim, 2002). In other words, the PLUS7 FA with nearly hydraulic-balanced mixing vane pattern (see Fig. 9) generates a random FA vibration only with the peak amplitude of about 3 × 10−5 m with the frequency of below 5 Hz in its operating range. It should be noted that the mixing vane pattern across the spacer grid assembly for the 16×16KOFA fuel does not have a perfect symmetry, i.e., perfectly hydraulicbalanced mixing vanes, since the mixing vane pattern at the center
Fig. 9. Flow paths generated by the hydraulic-balanced mixing vane pattern of the PLUS7 fuel (Kim and Suh, 2009).
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Table 4 Statistical information on initial spring force of the 16×16KOFA. Phase
Mean spring force (standard deviation)
Mean spring force loss (standard deviation)
One-sided 95/95% lower limit
Initial spring force before FR loading Spring force loss during FR loading Spring force loss during FA transport
16.750 N (0.840 N) 14.405 N 14.142 N
– 2.345 N (1.972 N) 0.263 N (0.245 N)
14.630 N 9.430 N 8.700 N
grid strip for the 16×16 geometry cannot be made to be perfectly hydraulic-balanced and furthermore orientations of mixing vanes around the guide thimble tubes were optimized to generate better mixing. Therefore, the flow paths shown in Fig. 9 are not perfectly symmetric. The flow directions drawn in Fig. 9 were determined by orientations of mixing vanes of each spacer grid strip. Naturally, one can say that the GTRFW-induced fuel failure of the 16×16KOFA is caused by the self-excited FA vibration with hydraulic-unbalanced mixing vane pattern across the spacer grid assembly. It is reported that the V5H FA also has experienced the GTRFW-induced fuel failure in the 3-loop Westinghouse type plants in the U.S. because of the self-excited FA vibration with hydraulic-unbalanced mixing vane pattern across the spacer grid assembly (Washington, 1993). However, it is noteworthy that the V5H FA without mixing vanes still generated the GTRFW-induced fuel failure in the Fort Calhoun plant in the U.S., which had not been expected from the viewpoints of the self-excited FA vibration. Therefore, it may be said that the GTRFW-induced fuel failure can occur partly due to the insufficient grid-to-rod supporting conditions coupled with a random FA vibration, even though there is no self-excited FA vibration. In order to investigate the impact of the 16×16KOFA design and a certain manufacturing process on the grid-to-rod supporting conditions, the various tests were performed such as initial spring force loss measurements during the FR loading process, additional spring force loss measurements during FA transport to the plant and cladding creep-down measurements for two kinds of cladding materials. From Figs. 5 and 6, it is found that with the FR loading speed of 0.33 m/s the initial spring forces of some spacer grid cells are less than 12 N, the minimum initial spring force requirement for preventing the GTRFW, while with the FR loading speed of 0.18 m/s the initial spring forces of all spacer grid cells is greater than 12 N. The statistical values for the spring force loss during the FR loading process and during the FA transport are summarized in Table 4. From this table, it can be seen that the mean value and standard deviation of the spring force loss during the FR loading process are 2.345 and 1.972 N, respectively, while those of the additional spring force loss during the FA transport are 0.263 and 0.245 N, respectively. From Fig. 8, on the other hand, it can be seen that there exists only a slight difference in the in-reactor cladding creep-down between cladding materials A and B, even though there is quite a large difference in the cladding thermal creep-down data between them (see Fig. 7). Considering that irradiation-induced cladding creep is much more dominant in the reactor operation than thermal-induced one, therefore, it is concluded that the cladding materials have no significant impact on the unexpected large grid-to-rod gap, if it would occur at the 16×16KOFA grid cells. Based on the statistical values of the initial spring force loss given in Table 4 and Fig. 5, three kinds of initial spring force are considered to have 8, 12 and 18 N to investigate the impact of initial spring force on the onset time of grid-to-gap appearance. The values of 18 and 12 N represent the mean initial spring force and the minimum one required in the design specification, respectively, while that of 8 N represents the one-sided 95/95% lower limit measured for fuel assemblies manufactured by the FR loading process with 0.33 m/s and transported to the plant. The GRIDFORCE program (Kim and Kim, 1997) was used to predict residual elastic spring deflection and force with relevant fuel
design data as a function of burnup. A calculation method used in the GRIDFORCE program can be briefly summarized as follows; A residual spring force at t = t at an operating temperature can be given by the Hooke’s law. fres (t) = COT ıres (t)
(1)
where fres (t) is the residual spring force at t = t, COT the spring constant at an operating temperature, and ıres (t) the residual elastic spring deflection at t = t. Then, the residual elastic deflection at t = t, ıres (t), can be derived by the following formula: ıres (t) = ıo − ıP (t) − ıT (t) − ıCR (t) − ıIR (t) − ırex (t)
(2)
where ıo is the initial elastic spring deflection at room temperature, ıP (t) the accumulated elastic cladding deflection caused by the coolant overpressure from t = 0 to t = t, ıT (t) the accumulated thermal expansion difference between spacer grid and cladding from t = 0 to t = t, ıCR (t) the accumulated cladding creep-down from t = 0 to t = t, ıIR (t) the accumulated spacer grid irradiation growth from t = 0 to t = t, and ırex (t) is the accumulated elastic spring deflection loss due to spring force relaxation from t = 0 to t = t. The value of ıo is determined by unstrained spring height, fuel rod pitch and initial cladding diameter, while the values of ıP (t), ıT (t), ıCR (t), ıIR (t) and ırex (t) are calculated with the use of reactor operating conditions and relevant fuel design data. A schematic diagram of elastic spring deflection as a function of burnup may be drawn, as shown in Fig. 10. As an illustration for grid-to-rod gap formation, changes in the residual elastic spring deflections, which were calculated by the GRIDFORCE program, are shown in Figs. 11 and 12. From these figures, it can be said that the initial spring force of 12 N does not generate a grid-to-rod gap at any spacer grids at the FA burnup of 22 MWD/kgU, whereas that of 8 N generates grid-to-rod gaps at the 3rd, 4th, 5th and 6th spacer grids. With the use of Eq. (1), changes in the residual spring forces were calculated for initial spring forces of 8, 12 and 18 N as a function of burnup, as shown in Fig. 13. From this figure, it is found that the grid-to-rod gap starts to form at 18, 24 and 28 MWD/kgU for the initial spring forces of 8, 12 and 18 N, respectively. It is obvious
Fig. 10. A schematic diagram of elastic spring deflection as a function of burnup.
K.-T. Kim / Nuclear Engineering and Design 240 (2010) 756–762
Fig. 11. Residual spring deflection at FA average burnup of 22 MWD/kgU for an initial spring force of 12 N.
Fig. 12. Residual spring deflection at FA average burnup of 22 MWD/kgU for an initial spring force of 8 N.
that the lower initial spring force generates the larger grid-to-rod gap at the same burnup or the earlier grid-to-rod gap. Therefore, one can say that some grid cells with the initial spring force less than 12 N, the minimum initial spring force required in the design specification, may be susceptible to the GTRFW-induced fuel fail-
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ure. Furthermore, the probability of the GTRFW-induced fuel failure goes up sharply when combined with the self-excited FA vibration. Considering that the initial spring loss of the 16×61KOFA is mainly due to the FR loading process with 0.33 m/s, however, the impact of the additional initial spring force loss and the cladding creep-down on the unexpected early grid-to-rod gap formation is considered relatively small. On the other hand, it should be noted that the initial spring force loss during the FR loading process can be reduced drastically if the loading speed is changed from 0.33 to 0.18 m/s, as shown in Fig. 6. Accordingly, the GTRFW-induced fuel failure could be eliminated if one would develop the 16×16KOFA with hydraulic-balanced mixing vanes and adopt the FR loading speed of less than 0.18 m/s. Based on the wide spectrum of test results obtained in this study, guidelines against the GTRFW-induced fuel failure are proposed as follows. Firstly, to minimize FA-induced vibration, one should develop spacer grid design with hydraulic-balanced mixing vane pattern across the spacer grid assembly. The spacer grid with mixing vanes considered needs to be verified against the selfexcited fuel assembly-induced vibration through the flow sweep tests of the full-scale fuel assembly covering operating range of nuclear power plants considered. Secondly, to minimize the gridto-rod fretting wear caused by insufficient grid-to-rod supporting conditions, one should maximize the grid-to-rod contact area but minimize the grid-to-rod gap size during the fuel lifetime. It should be kept in mind that the lower initial spring loss during the FR loading process, the smaller cladding creep-down and the smaller spacer grid width growth will generate the smaller grid-to-rod gap. Finally, with the use of the full-scale FA containing oxidized and non-oxidized FRs with various grid-to-rod gap sizes, the FA endurance tests need to be performed to verify the effect of the FR supporting conditions and the FA vibration on the fretting wear rate (Lu, 2001). 4. Conclusions A wide spectrum of various tests and some measurements were performed to investigate root causes of the grid-to-rod fretting wear-induced fuel failure observed in the 16×16 Korean Optimized Fuel Assemblies. The primary root cause is considered to be the selfexcited fuel assembly vibration occurring in the operating range with the peak amplitude of about 20 × 10−5 m with the frequency of between 1 and 20 Hz. The second effective root cause might be the initial spring force loss during the fuel rod loading process with 0.33 m/s loading speed. The grid-to-rod gap appears at much lower burnup with the one-sided 95% lower limit of initial spring force than expected with minimum initial spring force required in the design specification. On the other hand, guidelines against the GTRFW-induced fuel failure are proposed, indicating that the GTRFW-induced fuel failure might be eliminated if one develops the 16×16KOFA with hydraulically balanced mixing vanes and adopts the FR loading speed of less than 0.18 m/s. Acknowledgements This research was supported by Basic Atomic Energy Research Institute Program through the National Research Foundation of Korea (NRF) funded by the Ministry of Education, Science and Technology (No. 2009-0075907). References
Fig. 13. Residual spring forces as a function of FA average burnup.
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