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Potential causes of failures associated with power changes in LWR's
Journal of Nuclear Materials 87 (1979) 251-258 0 North-Holland ~blishing Company
POTENTIAL CAUSES OF FAILURES ASSOCIATED WITH POWER CHANGES IN LWR’s * P.
3OUFFIOUX, J, VAN VLIET, P. DE~~AIX
and M. LIPPENS
Belgonucleaire, rue du Champs de Mars 25, B-1050 Brussels, Belgium Received 30 April 1979
In modern fuel rod designs, the failures are mainly correlated with power changes: they are believed to be induced by a chemically assisted mechanical process, i.e. stress corrosion cracking. The present SCC understanding relies on the concept of threshold stress together with fission product ava~ability. Both topics are closely related to the clad and fuel material behaviour, which are modelled in the COMETHE code. The latter is used for design of fuel rods, and can provide a quantitative design and operation policy in order to minimize PC1 induced clad failure occurrence. The examination and assessment of new rod designs and code calibration has been and is still being carried out in the BR 3 PWR and in the BR 2 MTR, both located at the CEN/Mol, together with reloads in other PWR and BWR’s.
or both of the following mechanisms: - mechanical overload of the Zircaloy clad subjected to tensile stress concentrations; - stress corrosion cracking of the Zircaloy clad by aggressive fission products such as iodine, bromine and cesium. Up to now, the last mechanism is recognized as being the predominant PC1 failure mechanism. Extensive efforts are being made to provide practical solution to minimize the PC1 problems in the power plants. Operating restrictions have been recommended to rn~irn~e and hopefully prevent PC1 failures, while design improvements are developed to make the fuel more resistant to PCI’s and to eventually eliminate the need for operating restrictions. The approach adopted by Belgonucleaire to cope with the PCI-induced clad failure avoidance mainly comprises: - improving the understanding of PCI-induced failure mech~isms; - the development and the validation of design computer codes (COMETHE code); - analytical evaluations based on data from tests performed in experimental reactors or commercial reactors; - the institution of some operating restrictions together with some fuel rod design improvements,
1. Introduction Fuel reliability has increased and failure rate has been reduced to an acceptable level in the past ten years. The major fuel failure mechanism remaining today is the pellet clad mechanical interaction with failures associated to fission product stress corrosion (P(3 PCI-induced clad failures can occur during power ramp, when returning to full power after a refueling shutdown (in that case handling of irradiated fuel assemblies could lead to relocation of pellet fragments), when the power level of the reloaded assemblies is substantially increased from that during the previous cycle, or when returning to full power after a period (a few weeks) preconditioning the fuel at a lower power level (i.e. during control rod withdrawal pattern change in a BWR). Power cyclings as daily load follow-up operation might cause cumulative damage to the cladding resulting in a fatigue failure of the cladding. PCI-induced clad failure can be attributed to one
* Paper presented at the meeting on “Ramping and Load Following Behaviour of Reactor Fuel”, Petten, 30th Nov./ 1st Dec. 1978. 251
P. Bouffioux et al. /Potential
252
causes of failures associated with power changes in L WRs
2. PC&induced failure mechanism
2.2. Fission product availability at clad inner surface
The SCC problem has been outlined elsewhere [l-l 91. Nevertheless, in the frame of this presentation, the mechanisms involved in SCC are briefly reviewed. The following conditions are required for SCC to occur: - a critical stress (threshold stress) has to be exceeded; - a critical concentration of fission products has to be available at the cladding inner surface.
The nature of the fission products involved in the stress corrosion cracking of Zircaloy clad is not definitively established. Nevertheless, iodine and bromine have been identified as the most aggressive agents in Zircaloy SCC. Cesium is also recognized to participate in Zircaloy SCC. Very low corrosive fission product concentration seems to be required to accelerate the PC1 induced failure. Out-of-pile experiments performed in iodine environment have shown that a iodine concentration as low as 3 X lob3 to 7 X 10e3 mg/cm* is sufficient
2. I. Threshold stress
[5,171. The time to failure has been shown to be strongly correlated to the stress level above a so-called threshold stress. The threshold stress and the time to failure depend on the metallurgical state and the texture of the Zircaloy, on the irradiation exposure and on the aggressive fission product concentration at the inner clad surface. Irradiation exposure has a detrimental effect on the resistance of Zircaloy clad to stress corrosion cracking as it decreases the magnitude of the threshold stress [3,4]. Table 1 gives threshold stresses observed on irradiated and unirradiated Zircaloy materials [4,5]. The threshold stress for crack formation is predicted to be a low fraction of the yield stress of irradiated Zircaloy (0.3 to 0.5) [ 11,5]. The effect of the texture on stress corrosion cracking susceptibility is not well established. However, it would seem that Zircaloy clad with the basal poles close to the radial direction would be more resistant to see.
Table 1 Threshold
stresses for SCC
Item
‘Jth (kg/mm* 1
Unirradiated at 320°C
stress-relieved
Unirradiated
annealed
Unirradiated at 360°C
stress-relieved
Irradiated
Post-irradiation examinations of defected LWR fuel rods revealed that most of iodine capable of Zircaloy chemical attack and SCC is supplied from CsI which deposited on the inner clad surface. CsI is a stable halide which can be dissociated in Cs t I in the presence of radiation. Laboratory experiments demonstrate that CsI does not lead to SCC in the absence of radiation [ 13,141. Thermodynamic considerations lead to the conclusion that no free iodine or bromine could be expected in Zircaloy clad fuel rods [ 16,6]. Cesium as a fission product is produced in much larger concentration than iodine and bromine and these latter elements can be totally bonded to cesium. Nevertheless, the efficiency of the chemical reactions Cs t I -+ CsI and Cs t Br + CsBr strongly depend upon the oxygen potential in the fuel rod. The oxygen potential is governed by the stoichiometry (O/U ratio) of the fuel. The affmity of oxygen for cesium is higher than the affinity for iodine and bromine. Therefore, in hyperstoichiometric fuel, CsO will preferentially form and free iodine and bromine can be expected [5,15].
stress-relieved
at 320°C
28.5 30.5
Zircaloy-4
Zircaloy-4
of CsI versus temperature
Pvap (Torr)
Temperature
1 10 40 100 400 760
738 873 976 1055 1200 1280
(“C)
33.6
Zircaloy-2
Zircaloy-2
Table 2 Vapour pressure
at 360°C
20.4
-
P. Bouffioux et al. /Potentialcausesof failuresassociatedwithpower changesin L WR 's
Iodine and bromine are in gaseous phase at fuel operating temperatures. Therefore, it might be assumed that existing free iodine and bromine in the fuel matrix could be released by a process similar to fission gas release. The way by which chemical compounds like CsI can reach the clad inner surface is not yet clearly understood. One possible way would be a vapour transport mechanism. That mechanism is mainly governed by the vapour pressure of the considered species. For instance, table 2 gives the vapour pressure of CsI at various temperatures [ 181. 2.3. Power ramping and power cycling The initiation of a crack is the critical event in SCC failure. Once the crack has been initiated, and has reached a definite depth, propagation can proceed at stresses below the threshold stress. The crack propagation rate is a strong function of the stress intensity factor Kscc and increases with it according to the formula: dc/dt = .4Ktcc = A [f(u, c)]” , where c is the crack depth, A and n are constants [ 10, 53; n is approximately 9. The time required to initiate a crack has been estimated to be an important fraction (75%) of the time for failure [5]. The most important parameter is therefore the time at stresses above the threshold stress. This time depends on the primary creep properties of the clad material. During a ramp, if the tensile stresses generated by the strong pellet clad interaction relaxes significantly fast by primary creep so that the acting stress would fall below the threshold stress after a time lower than the time for failure, no SCC failure could occur and even incipient cracks could not be formed. A large number of short times excursion above the threshold stress could have the same impact than a long single time above the threshold value. Short excursions above threshold stress can occur during power cyclings.
3. Theoretical approach The theoretical approach towards the analysis of PCI-induced failure mechanisms is mainly based on
253
the Belgonucleaire integral fuel rod modelling code COMETHE. 3.1. Design criteria The design criteria as adopted by Belgonucleaire in its own design and limitative for PC1 are: Fuel temperature. At the present stage, the usual criterion is to prevent core melting and so to keep the peak fuel temperature below the UOa melting point (2850°C). Therefore, in the existing LWR, that criterion is always satisfied. Nevertheless, it is adviseable to keep the fuel temperature at a sufficiently low value to prevent an excessive release or migration of corrosive fission products as iodine, cesium iodides and to minimize the occurrence of clad failures by stress corrosion cracking. Clad stress range and SCC. Clad failure by SCC will not occur if: - The equivalent stress at clad inner surface in tensile conditions does not exceed the threshold stress for SCC. According to literature [2], the threshold stress for cold work stress relieved Zircaloy4 (material adopted by Belgonucleaire for its tubing) is 20 kg/ mm2 for irradiated material. To adopt that value as a threshold stress would be a good criterion. However, to take into account the local effects due to ridges, chipping of pellets and the fact that irradiation reduces that threshold and the time for failure, we presently adopt a more severe value, i.e. 14 kg/mm2 (135 MPa). - The amount of corrosive fission products which have been released from hot fuel and have deposited on the colder inner clad surface, is sufficiently low. At present state, it is assumed that the amount of corrosive fission products eventually capable of SCC is directly proportional to the fission gas release. On the basis of PIE results, we limit the fission gas release to a value not exceeding 1%. 3.2. Example of sensitivity assessment under steady state conditions
Various results of sensitivity analyses have been presented earlier: e.g. on fuel of the Maine Yankee :ore 1 type [20] and on fuel for type 15 X 15 PWR’s 211. The key results reported in [20] are represented I fig. 1. Since the proceedings of refs. [2 1,221 will ot be issued before 1979, we will indicate hereafter
P. Bouffioux et al. /Potential causes of failures associated with power changes in I, WRs
254
10
5
mllr
have been considered. Three typical histories selected for calculations (fig. 2) with mean core power of 226 W/cm represent respectively the worst conditions for pellet-clad mechanical interaction (hist. l), the max. LOCA related fuel temperature (hist. 2) and the max. fission gas release (hist. 3).
%
l
OIHSICIC’W
I 200
I 100
JJm
0.3
Fig. 1, Fission gas release as a function of radial gap.
the assumptions ma~ta~ed constant for most sensitivity calculations (fuel rod length, plenum chamber volume, cladding thickness and anisotropy, pelletclad diametral gap, pellet density, densification behaviour, grain size, pre-pressurization, power rating histories and power ramping). As a justification of the speci~cation of 10.75 mm outer diameter Zr 4 clad fuel rod typical for the 14 X 14 and 15 X 15 PWR’s in operation in Belgium, three types of fuel rod designs were selected (table 3). On the basis of core configurations and most likely assembly reshuffling patterns, various possible rod histories
A power ramping parametric study has been performed for a 17 X 17 type BR 3 fuel rod the characteristics of which are listed in table 4. The fuel rod has been assumed to be irradiated in low power rated core zone during the two first cycles and shuffled in a higher power rated core zone for the third cycle irradiation. The reference power history is plotted in fig. 3a. Figs. 3b and 3c show respectively the evolution of the hoop permanent strain E,; the contact pressure PC; the equivalent stress u,~. These results are predicted by the computer code COMETHE by assuming no restriction in starting-up of cycle 3 and by considering a ridging factor equal to 2. The fuel-cladding interaction starts at the beg~n~g of cycle 2. Once the contact pressure and the inner gas pressure exceeds the outer coolant pressure, tensile stresses are generated in the cladding so that the cladding creep down stops. As the power is ramped at the beginning of cycle 3 (BOC3), the strong interaction between the expanding fuel and the Zircaloy cladding induces very high tensile stress. The equivalent stress at BOC3 (28 kg/mm2) exceeds the threshold stress limit for SCC
Table 3 PWR fuel characteristics Rod type Clad external diameter (mm) Clad internal diameter (mm) Clad thickness (mm) Diametral gap km) Fuel density (% TD) Active fuel length (mm) Plenum length (mm)
A
c
B ----10.75
10.75
10.72
9.30
9.30
9.48
0.715
0.715
0.621
230
190
184
94
94
94
2438.4
2438.4
2438.4
130.7
130.7
130.7
ClCLE
1
I
I
I
0.5
CVCW
2
CYCLE 3
CVCLE 4
Fig. 2. Possible power histories considered in sensitivity assessment.
255
P. Bouffioux et al. /Potential causes of failures associated with power changes in L WR‘s Table 5 Irradiation conditions -._.
Table 4 Fuel rod characteristics
Linear power before ramp (W/cm)
17 x 17 9.5 8.36 165 93.5
Array Clad outer diameter (mm) Clad inner diameter (mm) Diametral gap km) Fuel bulk density (% TD) ____^_.~.
Burn-up before ramp (MWd~tm) -~
adopted as design criteria (14 kg/mm2). Such a situation may not be tolerated as the integrity of the fuel rod is endangered. Therefore, the power increase rate during starting-up has to be limited to allow progressive creep of the clad and so to reduce the stress level. The impact of different power ramps on the stressstrain cladding response has been investigated at starting of cycle 3. The irradiation conditions before and after ramp are summarized in table 5.
I50
a 0
170
340
510
680
650
,450
1020
30
i Y P w 300
20
$
:
c
z
1”
$it t 750
10
$ w
2 6
: In
0 0 0
170
340
510 680 TIME (Days)
850
1020
Fig. 3. Power ramping parametric study - reference case.
Ramped linear power ~~crn~ Fission gas release before ramp (%)
.~~-_._Peak Average
258 201
Peak Average
34650 27300
Peak Average
340 272 0.27
The different power ramps considered in that study and their characteristics are given in fig. 4. It is not the purpose of the present paper to detail the calculation results for each ramp. An ~port~t conclusion is all power ramps might be tolerated as the threshold stress for SCC is not exceeded in any case except for ramp 1 which is too severe and so may not be accepted. It is generally believed that steady state power periods are essential to enable stress relaxation by clad creep. The comparison of calculation results obtained for ramps 5 and 6 demonstrates that steady state power periods are not required and can even be detrimental. Fig. 5 compares the calculated contact pressure and equiv~ent stress for ramp 5 and 6: in both cases, the threshold stress for SCC is not exceeded. It is interesting to notice that, despite the stress relaxation during the steady power period, the contact pressure and equivalent stress predicted for ramp 5 at 100% nominal power is practically the same as those predicted for ramp 6. An ~termediate hold-down period is not significantly better than a slow ramp at 0.5% for instance. During the steady period, burn-up is accumulated and the fuel swells. Although being very low, this swelling is significant enough to have an effect on the pelletclad interaction. In addition, the gain in energy by considering the ramp 6 instead of ramp 5 is more than .50%,which is a decisive consideration. A limit on stress levels should only be applied if the corrosive fission products concentration is above the critical level. The total amounts of cesium, iodine and bromine produced during the cycles 1 and 2 have
P. Bouffioux
et al. /Potential
causes
of failures associated with power changes in L WR‘s
OTlME(days)
TIMEhlayd
Fig. 4. Power ramping parametric study - starting-up at BOC3. Lharacteristics of the considered power ramps.
been calculated and are presented in fig. 6. One can notice that the effect of the reshuffling period during which the unstable isotopes of iodine disappears by /3- decay is not significant. Cesium is more abundant (more than a factor 10) than iodine and bromine and therefore, in stoichiometric fuels, the iodine and bromine can be totally bonded to cesium. However, let us assume that all iodine and bromine produced are free and can therefore be released by a process similar to fission gas release. The total quantity of iodine and bromine produced at BOC3 is 14.4 cm3 STP. To be conservative, we assume that the fractional release attributable to all fission gases (Xe, Kr, I, Br) is the same that the fractional release calculated for xenon and krypton alone, i.e. 0.27%. That leads to 0.038 cm3 STP and 0.012 STP of respectively released iodine and bromine. It corresponds to an average concentration of 3.3 X lo-’ mg/cm’ of iodine and 6.5 X 10e4 mg/cm2 of bromine. We notice that the concentration of iodine is in the range of critical concentration for SCC observed in out-of-pile tests [5,17]. That concentration was obtained assum-
ing that all iodine and bromine were free, which is not the case in the reality, a significant amount of iodine and bromine being bonded to cesium.
4. Belgonucleaire power ramping and cycling data acquisition 4.1. BR 3 power ramping on SS clad PC1 leading to fuel failure has been observed for the first time in some stainless steel clad U02 rods irradiated in the VULCAIN and 2bis cores of the PWR BR3. Since some reactors like Chooz and Trino are still operating with SS clad fuel, Belgonucleaire has assessed these data, utilized them as design basis and complemented them by irradiations in Br2. 4.2. BR 2 power cycling on SS clad Two stainless steel clad mixed oxide fuel rods (one pelletized and one vibrocompacted) irradiated in the
P. Bouffioux et al. f Potential causes of failures associated with power changes in L WR ‘s
251
cores Vulcain and 2bis of the BR 3 have been re-irradiated in the BR 2 reactor under power cycling conditions. 4.3. BR 2 power cycling on Zr clad Various UO? and UOZ-PuOZ fuel rods fabricated for the BR 3 and Dodewaard reloads are presently irradiated in the Br 2 by means of baskets, boiling and pressurized water capsules. The irradiation programme includes a precycling phase and a cycling period with intermediate examinations, to get data on power ramping behaviour of fuel operating in a power cycling regime.
5. Conclusions
0
0
i
t
i
b 1 l?LAPSED TIME
b PROM
I b EOC 3 (
D.,:,
rb
II
Fig. 5. Power ramping parameteric study - starting-up at BOC3. Comparison of the effect of two different ramps.
The potential causes of failures associated with power changes in LWR’s have been reviewed. They were shown to be mainly attributed to PC1 defects, with a fission product enhancement through SCC. However, the mechanisms involving in SCC failure are not yet well understood. The modelling approach of the problem is therefore based on conservative design criteria adopted by Belgonucleaire. In order to increase the power plant capability (in the case of power ramp and load follow), provisions have to be taken to avoid detrimental PCI. Those provisions include intensive modelling work coupled with power ramp and cycling experiments performed in test reactors like BR 2.
References [l] J. Kenton, EPRI J. 5 (1978) 22. [ 21 J.T.A. Roberts and F.E. Gelhaus, 4th Intern. Conf. on
i% .27300 IRRADIATION
400 TIME
500 (D&/s)
MWd/tM 600
700
Fig. 6. Total production of cesium, bromine and iodine in a 17 X 17 fuel rod.
Zirconium in the Nuclear Industry, Stratford-upon-Avon, UK, June 1978. [3] J.T.A. Roberts and H. Ocken, EPRI J. 8 (1978) 75. [4] J.T.A. Roberts, R.L. Jones, E. Smith, D. Cubicciotti, A.K. Miller, H.F. Wachob and F.L. Yaggee, 4th Intern. Conf. on Zirconium in the Nuclear Industry, Stratfordupon-Avon, UK, June 1978. [S] D.O. Pickman, J.A. Gittus, R.A. Shaw, F.W. Trowse, P.D. Parsons, A. Garlick and R.A. Murgatroyd, Review of the Papers Presented at the 4th Intern. Conf. on Zirconium in the Nuclear Industry, Stratford-upon-Avon, 1978 and Prague, 1978.
258
P. Bouffioux
et al. /Potential
causes of failures associated with power changes in L WR S
[ 61 A. Garlick, ANS Topical Meeting on Water Reactor
Fuel Performance, St. Charles, IL, June 1977. [ 71 H. Hoppe, personal communication. [ 81 J.T.A. Roberts, personal communication. [9] M. Peehs, H. Stehle and E. Steinberg, 4th Intern. Conf. on Zirconium in the Nuclear Industry, Stratford-uponAvon, 1978. [lo] K. Videm and L. Lunde, IAEA Specialists Meeting on PC1 in LWR, Vienna, 1977. [ 1 l] E. Smith and A.K. Miller, IAEA Specialists Meeting, Blackpool, 1978. [ 121 W.J. Penn, R.K. Lo and J.C. Wood, Nucl. Technol. 34 (1977) 249. [ 131 J.T.A. Roberts, E. Smith, N. Fuhrman and D. Cubicciotti, Nucl. Technol. 35 (1977) 131. [ 141 J.C. Wood, ANS Topical Meeting, St. Charles, IL, 1977. [ 151 Th. M. Besmann and T.B. Lindemer, Nucl. Technol. 40 (1978) 297.
[ 161 F. GarzarolIi, R. Manzel, M. Peehs and H. Stehle, Kerntechnik 20 (1978) 27. [ 171 Katsumi Une (Toshiba), J. Nucl. Sci. Technol. 14 (1977) 433. [ 181 J.C. Bailar, Jr., Comprehensive Inorganic Chemistry (Pergamon, 1973). [ 191 K. Videm and L. Lunde, 4th Intern. Conf. on Zirconium in the Nuclear Industry, Stratford-upon-Avon, 1978. [20] H. Hoppe, 4th SMIRT Intern. Conf., San Francisco, 1977. [ 211 E. Demeulmeester and P. Deramaix, Symp. on Characterization and Quality Control of Nuclear Fuels, Karlsruhe, 1978; J. Nucl. Mater. 81 (1979) 161. 1221 H. Bairiot, P. Bouffioux, M. Gaube and C. VandenBerg, Intern. Symp. on LWR Fuel Element Fabrication with Special Emphasis on its Effect on Fuel Performance, Prague, 1978.
POTENTIAL CAUSES OF FAILURES ASSOCIATED WITH POWER CHANGES IN LWR’s * P.
3OUFFIOUX, J, VAN VLIET, P. DE~~AIX
and M. LIPPENS
Belgonucleaire, rue du Champs de Mars 25, B-1050 Brussels, Belgium Received 30 April 1979
In modern fuel rod designs, the failures are mainly correlated with power changes: they are believed to be induced by a chemically assisted mechanical process, i.e. stress corrosion cracking. The present SCC understanding relies on the concept of threshold stress together with fission product ava~ability. Both topics are closely related to the clad and fuel material behaviour, which are modelled in the COMETHE code. The latter is used for design of fuel rods, and can provide a quantitative design and operation policy in order to minimize PC1 induced clad failure occurrence. The examination and assessment of new rod designs and code calibration has been and is still being carried out in the BR 3 PWR and in the BR 2 MTR, both located at the CEN/Mol, together with reloads in other PWR and BWR’s.
or both of the following mechanisms: - mechanical overload of the Zircaloy clad subjected to tensile stress concentrations; - stress corrosion cracking of the Zircaloy clad by aggressive fission products such as iodine, bromine and cesium. Up to now, the last mechanism is recognized as being the predominant PC1 failure mechanism. Extensive efforts are being made to provide practical solution to minimize the PC1 problems in the power plants. Operating restrictions have been recommended to rn~irn~e and hopefully prevent PC1 failures, while design improvements are developed to make the fuel more resistant to PCI’s and to eventually eliminate the need for operating restrictions. The approach adopted by Belgonucleaire to cope with the PCI-induced clad failure avoidance mainly comprises: - improving the understanding of PCI-induced failure mech~isms; - the development and the validation of design computer codes (COMETHE code); - analytical evaluations based on data from tests performed in experimental reactors or commercial reactors; - the institution of some operating restrictions together with some fuel rod design improvements,
1. Introduction Fuel reliability has increased and failure rate has been reduced to an acceptable level in the past ten years. The major fuel failure mechanism remaining today is the pellet clad mechanical interaction with failures associated to fission product stress corrosion (P(3 PCI-induced clad failures can occur during power ramp, when returning to full power after a refueling shutdown (in that case handling of irradiated fuel assemblies could lead to relocation of pellet fragments), when the power level of the reloaded assemblies is substantially increased from that during the previous cycle, or when returning to full power after a period (a few weeks) preconditioning the fuel at a lower power level (i.e. during control rod withdrawal pattern change in a BWR). Power cyclings as daily load follow-up operation might cause cumulative damage to the cladding resulting in a fatigue failure of the cladding. PCI-induced clad failure can be attributed to one
* Paper presented at the meeting on “Ramping and Load Following Behaviour of Reactor Fuel”, Petten, 30th Nov./ 1st Dec. 1978. 251
P. Bouffioux et al. /Potential
252
causes of failures associated with power changes in L WRs
2. PC&induced failure mechanism
2.2. Fission product availability at clad inner surface
The SCC problem has been outlined elsewhere [l-l 91. Nevertheless, in the frame of this presentation, the mechanisms involved in SCC are briefly reviewed. The following conditions are required for SCC to occur: - a critical stress (threshold stress) has to be exceeded; - a critical concentration of fission products has to be available at the cladding inner surface.
The nature of the fission products involved in the stress corrosion cracking of Zircaloy clad is not definitively established. Nevertheless, iodine and bromine have been identified as the most aggressive agents in Zircaloy SCC. Cesium is also recognized to participate in Zircaloy SCC. Very low corrosive fission product concentration seems to be required to accelerate the PC1 induced failure. Out-of-pile experiments performed in iodine environment have shown that a iodine concentration as low as 3 X lob3 to 7 X 10e3 mg/cm* is sufficient
2. I. Threshold stress
[5,171. The time to failure has been shown to be strongly correlated to the stress level above a so-called threshold stress. The threshold stress and the time to failure depend on the metallurgical state and the texture of the Zircaloy, on the irradiation exposure and on the aggressive fission product concentration at the inner clad surface. Irradiation exposure has a detrimental effect on the resistance of Zircaloy clad to stress corrosion cracking as it decreases the magnitude of the threshold stress [3,4]. Table 1 gives threshold stresses observed on irradiated and unirradiated Zircaloy materials [4,5]. The threshold stress for crack formation is predicted to be a low fraction of the yield stress of irradiated Zircaloy (0.3 to 0.5) [ 11,5]. The effect of the texture on stress corrosion cracking susceptibility is not well established. However, it would seem that Zircaloy clad with the basal poles close to the radial direction would be more resistant to see.
Table 1 Threshold
stresses for SCC
Item
‘Jth (kg/mm* 1
Unirradiated at 320°C
stress-relieved
Unirradiated
annealed
Unirradiated at 360°C
stress-relieved
Irradiated
Post-irradiation examinations of defected LWR fuel rods revealed that most of iodine capable of Zircaloy chemical attack and SCC is supplied from CsI which deposited on the inner clad surface. CsI is a stable halide which can be dissociated in Cs t I in the presence of radiation. Laboratory experiments demonstrate that CsI does not lead to SCC in the absence of radiation [ 13,141. Thermodynamic considerations lead to the conclusion that no free iodine or bromine could be expected in Zircaloy clad fuel rods [ 16,6]. Cesium as a fission product is produced in much larger concentration than iodine and bromine and these latter elements can be totally bonded to cesium. Nevertheless, the efficiency of the chemical reactions Cs t I -+ CsI and Cs t Br + CsBr strongly depend upon the oxygen potential in the fuel rod. The oxygen potential is governed by the stoichiometry (O/U ratio) of the fuel. The affmity of oxygen for cesium is higher than the affinity for iodine and bromine. Therefore, in hyperstoichiometric fuel, CsO will preferentially form and free iodine and bromine can be expected [5,15].
stress-relieved
at 320°C
28.5 30.5
Zircaloy-4
Zircaloy-4
of CsI versus temperature
Pvap (Torr)
Temperature
1 10 40 100 400 760
738 873 976 1055 1200 1280
(“C)
33.6
Zircaloy-2
Zircaloy-2
Table 2 Vapour pressure
at 360°C
20.4
-
P. Bouffioux et al. /Potentialcausesof failuresassociatedwithpower changesin L WR 's
Iodine and bromine are in gaseous phase at fuel operating temperatures. Therefore, it might be assumed that existing free iodine and bromine in the fuel matrix could be released by a process similar to fission gas release. The way by which chemical compounds like CsI can reach the clad inner surface is not yet clearly understood. One possible way would be a vapour transport mechanism. That mechanism is mainly governed by the vapour pressure of the considered species. For instance, table 2 gives the vapour pressure of CsI at various temperatures [ 181. 2.3. Power ramping and power cycling The initiation of a crack is the critical event in SCC failure. Once the crack has been initiated, and has reached a definite depth, propagation can proceed at stresses below the threshold stress. The crack propagation rate is a strong function of the stress intensity factor Kscc and increases with it according to the formula: dc/dt = .4Ktcc = A [f(u, c)]” , where c is the crack depth, A and n are constants [ 10, 53; n is approximately 9. The time required to initiate a crack has been estimated to be an important fraction (75%) of the time for failure [5]. The most important parameter is therefore the time at stresses above the threshold stress. This time depends on the primary creep properties of the clad material. During a ramp, if the tensile stresses generated by the strong pellet clad interaction relaxes significantly fast by primary creep so that the acting stress would fall below the threshold stress after a time lower than the time for failure, no SCC failure could occur and even incipient cracks could not be formed. A large number of short times excursion above the threshold stress could have the same impact than a long single time above the threshold value. Short excursions above threshold stress can occur during power cyclings.
3. Theoretical approach The theoretical approach towards the analysis of PCI-induced failure mechanisms is mainly based on
253
the Belgonucleaire integral fuel rod modelling code COMETHE. 3.1. Design criteria The design criteria as adopted by Belgonucleaire in its own design and limitative for PC1 are: Fuel temperature. At the present stage, the usual criterion is to prevent core melting and so to keep the peak fuel temperature below the UOa melting point (2850°C). Therefore, in the existing LWR, that criterion is always satisfied. Nevertheless, it is adviseable to keep the fuel temperature at a sufficiently low value to prevent an excessive release or migration of corrosive fission products as iodine, cesium iodides and to minimize the occurrence of clad failures by stress corrosion cracking. Clad stress range and SCC. Clad failure by SCC will not occur if: - The equivalent stress at clad inner surface in tensile conditions does not exceed the threshold stress for SCC. According to literature [2], the threshold stress for cold work stress relieved Zircaloy4 (material adopted by Belgonucleaire for its tubing) is 20 kg/ mm2 for irradiated material. To adopt that value as a threshold stress would be a good criterion. However, to take into account the local effects due to ridges, chipping of pellets and the fact that irradiation reduces that threshold and the time for failure, we presently adopt a more severe value, i.e. 14 kg/mm2 (135 MPa). - The amount of corrosive fission products which have been released from hot fuel and have deposited on the colder inner clad surface, is sufficiently low. At present state, it is assumed that the amount of corrosive fission products eventually capable of SCC is directly proportional to the fission gas release. On the basis of PIE results, we limit the fission gas release to a value not exceeding 1%. 3.2. Example of sensitivity assessment under steady state conditions
Various results of sensitivity analyses have been presented earlier: e.g. on fuel of the Maine Yankee :ore 1 type [20] and on fuel for type 15 X 15 PWR’s 211. The key results reported in [20] are represented I fig. 1. Since the proceedings of refs. [2 1,221 will ot be issued before 1979, we will indicate hereafter
P. Bouffioux et al. /Potential causes of failures associated with power changes in I, WRs
254
10
5
mllr
have been considered. Three typical histories selected for calculations (fig. 2) with mean core power of 226 W/cm represent respectively the worst conditions for pellet-clad mechanical interaction (hist. l), the max. LOCA related fuel temperature (hist. 2) and the max. fission gas release (hist. 3).
%
l
OIHSICIC’W
I 200
I 100
JJm
0.3
Fig. 1, Fission gas release as a function of radial gap.
the assumptions ma~ta~ed constant for most sensitivity calculations (fuel rod length, plenum chamber volume, cladding thickness and anisotropy, pelletclad diametral gap, pellet density, densification behaviour, grain size, pre-pressurization, power rating histories and power ramping). As a justification of the speci~cation of 10.75 mm outer diameter Zr 4 clad fuel rod typical for the 14 X 14 and 15 X 15 PWR’s in operation in Belgium, three types of fuel rod designs were selected (table 3). On the basis of core configurations and most likely assembly reshuffling patterns, various possible rod histories
A power ramping parametric study has been performed for a 17 X 17 type BR 3 fuel rod the characteristics of which are listed in table 4. The fuel rod has been assumed to be irradiated in low power rated core zone during the two first cycles and shuffled in a higher power rated core zone for the third cycle irradiation. The reference power history is plotted in fig. 3a. Figs. 3b and 3c show respectively the evolution of the hoop permanent strain E,; the contact pressure PC; the equivalent stress u,~. These results are predicted by the computer code COMETHE by assuming no restriction in starting-up of cycle 3 and by considering a ridging factor equal to 2. The fuel-cladding interaction starts at the beg~n~g of cycle 2. Once the contact pressure and the inner gas pressure exceeds the outer coolant pressure, tensile stresses are generated in the cladding so that the cladding creep down stops. As the power is ramped at the beginning of cycle 3 (BOC3), the strong interaction between the expanding fuel and the Zircaloy cladding induces very high tensile stress. The equivalent stress at BOC3 (28 kg/mm2) exceeds the threshold stress limit for SCC
Table 3 PWR fuel characteristics Rod type Clad external diameter (mm) Clad internal diameter (mm) Clad thickness (mm) Diametral gap km) Fuel density (% TD) Active fuel length (mm) Plenum length (mm)
A
c
B ----10.75
10.75
10.72
9.30
9.30
9.48
0.715
0.715
0.621
230
190
184
94
94
94
2438.4
2438.4
2438.4
130.7
130.7
130.7
ClCLE
1
I
I
I
0.5
CVCW
2
CYCLE 3
CVCLE 4
Fig. 2. Possible power histories considered in sensitivity assessment.
255
P. Bouffioux et al. /Potential causes of failures associated with power changes in L WR‘s Table 5 Irradiation conditions -._.
Table 4 Fuel rod characteristics
Linear power before ramp (W/cm)
17 x 17 9.5 8.36 165 93.5
Array Clad outer diameter (mm) Clad inner diameter (mm) Diametral gap km) Fuel bulk density (% TD) ____^_.~.
Burn-up before ramp (MWd~tm) -~
adopted as design criteria (14 kg/mm2). Such a situation may not be tolerated as the integrity of the fuel rod is endangered. Therefore, the power increase rate during starting-up has to be limited to allow progressive creep of the clad and so to reduce the stress level. The impact of different power ramps on the stressstrain cladding response has been investigated at starting of cycle 3. The irradiation conditions before and after ramp are summarized in table 5.
I50
a 0
170
340
510
680
650
,450
1020
30
i Y P w 300
20
$
:
c
z
1”
$it t 750
10
$ w
2 6
: In
0 0 0
170
340
510 680 TIME (Days)
850
1020
Fig. 3. Power ramping parametric study - reference case.
Ramped linear power ~~crn~ Fission gas release before ramp (%)
.~~-_._Peak Average
258 201
Peak Average
34650 27300
Peak Average
340 272 0.27
The different power ramps considered in that study and their characteristics are given in fig. 4. It is not the purpose of the present paper to detail the calculation results for each ramp. An ~port~t conclusion is all power ramps might be tolerated as the threshold stress for SCC is not exceeded in any case except for ramp 1 which is too severe and so may not be accepted. It is generally believed that steady state power periods are essential to enable stress relaxation by clad creep. The comparison of calculation results obtained for ramps 5 and 6 demonstrates that steady state power periods are not required and can even be detrimental. Fig. 5 compares the calculated contact pressure and equiv~ent stress for ramp 5 and 6: in both cases, the threshold stress for SCC is not exceeded. It is interesting to notice that, despite the stress relaxation during the steady power period, the contact pressure and equivalent stress predicted for ramp 5 at 100% nominal power is practically the same as those predicted for ramp 6. An ~termediate hold-down period is not significantly better than a slow ramp at 0.5% for instance. During the steady period, burn-up is accumulated and the fuel swells. Although being very low, this swelling is significant enough to have an effect on the pelletclad interaction. In addition, the gain in energy by considering the ramp 6 instead of ramp 5 is more than .50%,which is a decisive consideration. A limit on stress levels should only be applied if the corrosive fission products concentration is above the critical level. The total amounts of cesium, iodine and bromine produced during the cycles 1 and 2 have
P. Bouffioux
et al. /Potential
causes
of failures associated with power changes in L WR‘s
OTlME(days)
TIMEhlayd
Fig. 4. Power ramping parametric study - starting-up at BOC3. Lharacteristics of the considered power ramps.
been calculated and are presented in fig. 6. One can notice that the effect of the reshuffling period during which the unstable isotopes of iodine disappears by /3- decay is not significant. Cesium is more abundant (more than a factor 10) than iodine and bromine and therefore, in stoichiometric fuels, the iodine and bromine can be totally bonded to cesium. However, let us assume that all iodine and bromine produced are free and can therefore be released by a process similar to fission gas release. The total quantity of iodine and bromine produced at BOC3 is 14.4 cm3 STP. To be conservative, we assume that the fractional release attributable to all fission gases (Xe, Kr, I, Br) is the same that the fractional release calculated for xenon and krypton alone, i.e. 0.27%. That leads to 0.038 cm3 STP and 0.012 STP of respectively released iodine and bromine. It corresponds to an average concentration of 3.3 X lo-’ mg/cm’ of iodine and 6.5 X 10e4 mg/cm2 of bromine. We notice that the concentration of iodine is in the range of critical concentration for SCC observed in out-of-pile tests [5,17]. That concentration was obtained assum-
ing that all iodine and bromine were free, which is not the case in the reality, a significant amount of iodine and bromine being bonded to cesium.
4. Belgonucleaire power ramping and cycling data acquisition 4.1. BR 3 power ramping on SS clad PC1 leading to fuel failure has been observed for the first time in some stainless steel clad U02 rods irradiated in the VULCAIN and 2bis cores of the PWR BR3. Since some reactors like Chooz and Trino are still operating with SS clad fuel, Belgonucleaire has assessed these data, utilized them as design basis and complemented them by irradiations in Br2. 4.2. BR 2 power cycling on SS clad Two stainless steel clad mixed oxide fuel rods (one pelletized and one vibrocompacted) irradiated in the
P. Bouffioux et al. f Potential causes of failures associated with power changes in L WR ‘s
251
cores Vulcain and 2bis of the BR 3 have been re-irradiated in the BR 2 reactor under power cycling conditions. 4.3. BR 2 power cycling on Zr clad Various UO? and UOZ-PuOZ fuel rods fabricated for the BR 3 and Dodewaard reloads are presently irradiated in the Br 2 by means of baskets, boiling and pressurized water capsules. The irradiation programme includes a precycling phase and a cycling period with intermediate examinations, to get data on power ramping behaviour of fuel operating in a power cycling regime.
5. Conclusions
0
0
i
t
i
b 1 l?LAPSED TIME
b PROM
I b EOC 3 (
D.,:,
rb
II
Fig. 5. Power ramping parameteric study - starting-up at BOC3. Comparison of the effect of two different ramps.
The potential causes of failures associated with power changes in LWR’s have been reviewed. They were shown to be mainly attributed to PC1 defects, with a fission product enhancement through SCC. However, the mechanisms involving in SCC failure are not yet well understood. The modelling approach of the problem is therefore based on conservative design criteria adopted by Belgonucleaire. In order to increase the power plant capability (in the case of power ramp and load follow), provisions have to be taken to avoid detrimental PCI. Those provisions include intensive modelling work coupled with power ramp and cycling experiments performed in test reactors like BR 2.
References [l] J. Kenton, EPRI J. 5 (1978) 22. [ 21 J.T.A. Roberts and F.E. Gelhaus, 4th Intern. Conf. on
i% .27300 IRRADIATION
400 TIME
500 (D&/s)
MWd/tM 600
700
Fig. 6. Total production of cesium, bromine and iodine in a 17 X 17 fuel rod.
Zirconium in the Nuclear Industry, Stratford-upon-Avon, UK, June 1978. [3] J.T.A. Roberts and H. Ocken, EPRI J. 8 (1978) 75. [4] J.T.A. Roberts, R.L. Jones, E. Smith, D. Cubicciotti, A.K. Miller, H.F. Wachob and F.L. Yaggee, 4th Intern. Conf. on Zirconium in the Nuclear Industry, Stratfordupon-Avon, UK, June 1978. [S] D.O. Pickman, J.A. Gittus, R.A. Shaw, F.W. Trowse, P.D. Parsons, A. Garlick and R.A. Murgatroyd, Review of the Papers Presented at the 4th Intern. Conf. on Zirconium in the Nuclear Industry, Stratford-upon-Avon, 1978 and Prague, 1978.
258
P. Bouffioux
et al. /Potential
causes of failures associated with power changes in L WR S
[ 61 A. Garlick, ANS Topical Meeting on Water Reactor
Fuel Performance, St. Charles, IL, June 1977. [ 71 H. Hoppe, personal communication. [ 81 J.T.A. Roberts, personal communication. [9] M. Peehs, H. Stehle and E. Steinberg, 4th Intern. Conf. on Zirconium in the Nuclear Industry, Stratford-uponAvon, 1978. [lo] K. Videm and L. Lunde, IAEA Specialists Meeting on PC1 in LWR, Vienna, 1977. [ 1 l] E. Smith and A.K. Miller, IAEA Specialists Meeting, Blackpool, 1978. [ 121 W.J. Penn, R.K. Lo and J.C. Wood, Nucl. Technol. 34 (1977) 249. [ 131 J.T.A. Roberts, E. Smith, N. Fuhrman and D. Cubicciotti, Nucl. Technol. 35 (1977) 131. [ 141 J.C. Wood, ANS Topical Meeting, St. Charles, IL, 1977. [ 151 Th. M. Besmann and T.B. Lindemer, Nucl. Technol. 40 (1978) 297.
[ 161 F. GarzarolIi, R. Manzel, M. Peehs and H. Stehle, Kerntechnik 20 (1978) 27. [ 171 Katsumi Une (Toshiba), J. Nucl. Sci. Technol. 14 (1977) 433. [ 181 J.C. Bailar, Jr., Comprehensive Inorganic Chemistry (Pergamon, 1973). [ 191 K. Videm and L. Lunde, 4th Intern. Conf. on Zirconium in the Nuclear Industry, Stratford-upon-Avon, 1978. [20] H. Hoppe, 4th SMIRT Intern. Conf., San Francisco, 1977. [ 211 E. Demeulmeester and P. Deramaix, Symp. on Characterization and Quality Control of Nuclear Fuels, Karlsruhe, 1978; J. Nucl. Mater. 81 (1979) 161. 1221 H. Bairiot, P. Bouffioux, M. Gaube and C. VandenBerg, Intern. Symp. on LWR Fuel Element Fabrication with Special Emphasis on its Effect on Fuel Performance, Prague, 1978.