ML20040H223
| ML20040H223 | |
| Person / Time | |
|---|---|
| Site: | Comanche Peak  |
| Issue date: | 03/31/1978 |
| From: | Office of Nuclear Reactor Regulation |
| To: | |
| Shared Package | |
| ML18023A006 | List: |
| References | |
| NUDOCS 8202170461 | |
| Download: ML20040H223 (27) | |
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ANA6YSISBRANCH
~ DIVISION OF SYSTEMS SAFETY i 0FFIbE OF NUCLEAR REACTOR ~ REGULATION WCAP-8762
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SAFETY EVALUATION REPORT l
9.N.
WRB-1 CRITICAL HEAT FLUX CORRELATION e
MARCH 1978
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8202170461 820$12 E"
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WRB-1 SER Introduction The -staff has reviewed the WRB-1 'tritical Heat Flux (CHF) correlation as 1
presented in topical report WCAP-8762.
The staff evaluation of the WRB-1 CHF correlation is' presented in this report.
The Westinghouse Rod Bundle, critical heat flux correlation '(WRB-1) is an empirical correlation which specifies the critical heat flux (i.e., the heat flux at which departure from nucleate boiling occurs) as a function of local conditions in a rod bundle.
The local conditions in the rod bundle were calculated with the standard Westingh'ouse thermal hydraulic design code THINC.
This correlation is based on 24 test series with a total of 1147 i
l 4ata points.
Each of the 24 test series was conducted on an electrically s
i heated rod bundle containing from 9 to 25 rods.
Each test series correspends to a different rod bundle geometry except for one repeatability test series.
The 24 test series include variations in heated length, rod size and configuration spacer grid design, grid spacing, and axial heat flux distribution.
The range of coolant condition tested corresponds to the proposed range of coolant conditions for application to PWRs.
Westinghouse has indicated that the WRS-1 correlation may be used to replace the W-3 correlation in' the core thermal hydraulic design for both the 15x15 fuel assembly design and the 17xl7 fuel asspmbly design.
The Westinghouse topical report WCAP-8752 concludes that the WRS-1 correlation is significantly more accurate than the W-3 correlation in predicting departure frem nucleate bciling.
On the basis f t.e im:r:s5p c:rri.1:icr. Ic:uracy the :r:cosse value of a linimum departure fr:m nucleate boiling ratio (DNER) to meet t.5e reac
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design criterion of a 95t probability of not experiencing departure from i
nucleate boiling on a limiting rod at a 95% confidence level is 1.17. The
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comparable value'for the. W-3 correlation is approximately 1.3.
Scope of Review The staff rer,iew of WCAP-8762 has included an independent audit of the subchannel calculation perfonned to detennine the local coolant conditions in the rod bundle. Approximately 300 calculations have been perfomed with COBRA-IV and the results have been compared,with the published results from the Westinghouse THINC calculations. The staff review has also included an independent audit of the statisti. cal calculations used in establishing the DNBR design limit for the correlation.
The staff has reviewed the assumptions made in generating the CHF correlation and the DNBR design limit as well as the assumptions necessary in order to apply the correlation to the PWR geccetry and coolant conditions. The WRS-1 correlation was also compared to other' PWR CHF correlations to identify any anamalous behavior.
During the review, the staff requested and. received additional infomation in several areas.
Review Sumary Results of Audit Calculations The results of the staff audit calculations are presented in Tables I, I:
and III. Table 1 presents a comparison'of calculated local quality for THINC and COBRA-III and IV.
Both the THINC code and the COBRA code have been reviewed and c:moared to ex:e-hental data; and we c:ncluded that either code could be used to establish the local c:ncitions recuired for the develocment of the CHF correlation.
The c:mparison indicates l
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a very good agreement between the two methods.
Since the THINC and the COBRA-III and IV calculations.;are both subchannel calculations the term
" local conditions" refers to average conditions in a subchannel. Each
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subchannel in,the red bundle is represented in both codes and the effects of flow redistribution are included.
Table II presents the results of a typical COBRA-IV calculation for one of the CHF tests. These calculations
' indicate that the enthalpy distribution in the test bundle is nearly uniform.
It appears that the rod poWtrs and the peripheral channel e
hydraulic diameters were chosen, in inost cases, to produce a nearly equal enthalpy rise in each channel.
This minimizes the effects of
. crossflow and interchannel mixing since fluid exchange,between adjacent channels does 'not result in any significant energy exchange. The analysis of the local coolant conditions is therefore relatively insensitive to changes -in themal diffusion coefficient, spacer grid flow resistance, crossflow resistance and other modeling assumptions. This is a sound approach to CHF testing since it minimizes the possible effects of
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calculational bias and calculaticnal uncertainties en the resulting CHF correlation.
During the initial phase of the staff review of this correlation only COBRA-III was available. Subsequently COBRA-IV became available allcwing T
comparisons among the three subchannel codes.
4 Table III presents a ecmparison of the Measured CHF/ Predicted CHF for THINC The mean value' of the Measured CHF/ Predicted CHF and the and COBRA-IV.
s stancard ceviations are in gocd agree ent.
Table :" ;resen:s_a similac ccmparison bet.veen COBRA-III and COSRA-IV for test series A-1 (thrcughcu l
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this report the test series will be designated by the identification given in Table 3-1 of WCAP-8762). 'A standard, statistical analysis of variance test', the F-test, was performed on the THINC, COBRA-IV results for tests series A-1.
T[ie purpose of this test was to detennine if differences in the me.thod of calculating the local coolant condition were responsible for the observed variation in the CHF data. The results of this F-test clearly indicate (i.e., at a 99%' certainty) that the variation,among the mean values of Measured CHF/ Predicted CHF predicted by THINC, COBRA-III and COBRA-IV is much smaller than would be expected from a random sample with the same variation among the individual data points.
This implies that it is unlikely that the variation among the data points is the result of,the method of solving for the subchannel coolant condition.
I The good agreement between THINC and CCSRA-III and IV in terms of local conditions and Measured CHF/ Predicted CHF; and the relatively simple nature of the subchannel calculations indicate that the uncertainty in the local subchannel condition has only a small contribution to the overall uncertainty in the correlation. The difference between the COBRA-III and COBRA-IV individual cases was generally less than 1%; an'd the difference between the COSRA-III and COBRA-IV values for the mean of the Measured CHF/ Predicted CHF was less than 0.1%.. The differences between th'e THINC and COBRA-IV individual cases were generally on the order of 2% or less; and the differences between the THINC and COBRA-IV values for the mean of the s
Measured CHF/ Predicted CHF were generally less than li.
However, in the case
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of tes: series A-5 a censisten. ofas cf grea se than is fcund :et.sein - e THINC results ard the CCBRA results.
This difference was traced to
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the cosine axial power distribution used by Westinghouse' which was not fully normalized (i.e., average axial power = 1.008 rather than 1.00).
This, difference results in a conservative bias in the Westinghouse results for the A-5 test series.
In addition, it should be noted 1
that many of the test series which form the basis for the WRB-1 correlation were also used to support previous Westinghouse DNBR correlations, and that many.of these test series were previously subjected to review and comparisons to COBRA calculations.
Correlation Assumations The assumed form of the correlation and the modeling of average subchannel coolant conditions rather than actual local coolant conditions are important considerations in determining the acceptability of the
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correlation. Additional assumptions related to the application of the 4
correlation to PWRs will be discussed later.
The form of the WRB-1 correlation is as follows:
l-q CHF = PF + Al + B3 x Flow = B4 x Flow x Quality where:
PF (Performance f actor) is a function of mixing vane design l
Al is a function of Pressure, Flow, Heated Length to the CHF location, Grid Spacing and Distance from the. last grid 83 is a constant B4 is a function of Pressure, Flow and' Heated Length.
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.,A Correlations using a similar form have been used successfully in the past and the general trend df decreasing CHF with increasing fluid quality has been observed in many experiments. The correlation also depends on Pressure, local flow, heated length, grid spacing, distance from the last grid, hydraulic diameter, and mixing vane design. The form of the correlation relative to these parameters and the correlation coefficients are highly empirical.
In some instances the The physical basis for the individual correlation terms is not apparent.
inclusion of a term for the heated length appears to result in a more accurate correlation but the physical basis for this term is unclear.
In a correlation which,uses actual local conditions, a heated length" term would appear to be Another example df an apparently non-physical representation in unnecessary.
the WRS-1 correlation is the separation of the mixing vane design effect,
from the grid spacing effect.
The mixing vane grids have the effect of promoting downstream turbulent mixing and an upstream flow reduction due to the flow resistance.
It is reasonable to expei:t that the effect on CHF would be a function of the grid spacing. A large grid spacing, for example, could result in' the downstream turbulent mixing effect washing out before the next. mixing vane grid, could reestablish the turbulent mixing pattern.
Previously published Westinghouse data (WCAP-7411-1-P A) on various mixing t
vane designs and grid spacing appear to she'w a dependence of the mixing vane design effect and or. the grid spacing.
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. -7 These are two examples of areas in which the correlation may not be correctly
.a represent:ing the physical phenomena. This has two important implications; first, a non-physical representation of phenomena in the correlation could be a contributing factor to the scatter in the comparison of Measured CHF and Predicted CHF for a given series of tests. This effect is inherently accounted for in the statistical analysis of the data for that given test series. The second implication of having non-physical represe~ntation of important '
phenomena is that the extrapolation of the WRB-1 correlation to other geometries of coolant conditions may not be valid.
In fact the practice of combining data from different geometries in calculatino an overall correlation uncertainty may not be valid. This subject will be discussed more extensively in a later section of this report.
An important new feature of the WRB-1 CHF correlation is the inclusion,
.of a term which accounts for the distance (DG) betweery the CHF location and the last upstream mixing vane grid. This term reflects the observed behavior in which CHF occurred predominantly in locations just upstream of a mixing vane grid. WCAP-8762 indicates that the WRB-1 correlation, including the DG term, has an overall 81.7% accuracy in p'redicting the location of CHF. The method of establishing this accuracy is discussed on pace 3-6 of WCAP-8762.
It aopears that this accuracy in pred'icting the correct location of CHF is a major factor in reducing the' scatter of the data within each test series. Since there is a physical basis to expect the CHF to decend on the distance fr:m t.'s last -4xing vane grid and since it acoears to make the correlation mere accura e :r.e use of tne DG term is acceptable.
n res:cr.se to a staff question, Westinghouse stated that in the THINC subchannel i
analysis of. the-CHF data, the axial node just belcw a mixing vane grid was j
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8-assumed to have a value of DG equal to the full grid spacing. This treatment is acceptable but requires that de same assumption be made in the application of the WRB-1 correlation to PWR calculations.
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The WRB-1 CHF correlation uses average subchannel coolant conditions ar input.
This technique inherently excludes consideration of the flow distribution, temperature distribution or void distribution (i.e., flow pattern) within a.
subchannel. Since the mixing vane gridsend unheated rods can affect the
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distribution of flow, temperature and voids in the subchannel, the correlation needs empirical tems to account for these effects. The use of a subchannel analysis in developing the correlation implies that a similar technique needs to be used in the application of the correlation.
Since there are important phenomena on a scale smaller than a subchannel, the use of a sub-channel analysis means that the correlation must include empirical tems.
As discussed previously, the empirical nature of the correlation may be one of the important contribut'ing factors to the overall correlation uncertainty and has important implications relative to extrapolation of the correlation'.
The staff positions relative to the extrapolation of the correlation and the establishment of the DNBR limit will be discussed in a later section of this report.
Statistical Review
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The WRB-1 correlation has been reviewed to detemine if there are any residual trends (i.e., systematic variations in the Measured CHF/ Predicted CHF) as a function of coolant conditions. When the data from the 24 :sst series are viewed as a whole, there do not appear to be any residual trends with changes in coolant conditions. Figures 3-1, 3-5 and 3-6 in the topical i
were cresented to demcnstrate.the lack of such trends., The data have aisc l
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been reviewed to detemine if there are any systematic trends with changes
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.in geometry. This can be tested y comparing the variance in the data from one test series with the variance in the data from a second test series. When looking for systematic variations among three or more test series the variation' in the test series means can be compared to the variation of the, data within the test series.
In both cases a standard statistical analysis of variance test can be applied to detemine if the, differences found are statistically significant.
In fact, the presence of systematic variations can be identified with any degree of confidence desired. The analysis of variance, F-test, has been used to identify systematic variations among the test-series at a 99%
confidence level. The results of these tests are sumarized 1n Table-V, which presents the calculated values of F and the theoretical range of F values at a 99'.' confidence level for truly random data. Where:
F=S 3
B W
2 S
= variance of the means of the test series B
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= variance of the data within the test series y
The theoretical range given in Table V shows that if the data were truly randem there is a 99% probability that the calculated value of F would fall within that range. A value of F below the range indic'ates that there is too-little variation among.the means and a value"of F above the range indicates that there is too much variation among the means.
Sixteen individual F-tests were ::erformed to determine if systematic trends exist ancnq the test series.
As incicated in Table-V cnly three of -he 16 F ests results in values within the expected range.
For truly randem errors all 16 wculd have been ex;:ected to fall. within the range.
In a few instances ee F-value is beicw
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In,other we'rds the correlation was chosen to minimize the variations in the Measured CHF/ Predicted CHF.
In these cases the correlation was chosen in a manner which eliminated some of the riatural variations between the tes,t series.
For mor.c of the comparisons given in Table-V the calculated value of F significantly excesded the maximum expected value. The implication of these systematic trends is that the change in geometry ficm one test series to another introduces another component of variance. The following example is presented to illustrate the potential problems associated with combining all of the test series results to establish a DNBR limit without accounting.for a component of variance due to gecmetry changes. As indicitted in Table 3-1 of WCAP-8762 the mean value of the Measured CHF/ Predicted CHF for all 1147 data points is 1.0043 and the correspondir'ig standa:
deviation is 0.0873. This.mean value and this standard deviation were used to t
establish the proposed DNBR limit of 1.17;.where the intent is that a calculated value of DNBR of 1.17 corresponds to 95% probability (at a 95% cenfidence level) of not experiencing DNB.
In dealing with the CHF data the DNBR corresponds to the Predicted CHF/ Measured CHF.
Therefore the DNER limit corresponds to a Measured CHF/ Predicted CHF value of 1/1.17 or 0.855..The ex::ected percentage of data points with a; Measured CHF/ Predicted CHF below 0.855 would therefore be approximately 5%. Test series A-20 includes 36 data points. We wculd I
therefore expect that fewer than two data points would have values of Measured CHF/Predic:ed CHF less -han 0.355. The ac aal data i-dica e that six data points were below 0.955, which is core than three times the number which the correlation implies.
Similarly, the mean value o' l.0C43 for all 112 da a
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poir)ts implies.g50% of_the data _have. values above.l.0043 and.550% have....
- values below 1.0043 'For test siries A-20 none of the, data points has a t
value of' Measured CHF/ Predicted CHF above'l.0043, in fact none has a value above.9402.
Clearly, the mean value and standard deviation of all 1147 data points have little or no meaning when we are dealing with the test geometry for series A-20.
Similar problems, generally of smaller magnitude, exist when the mean value and standard deviations based on all 1147 data points are
' compared to other of the test series.
Combining all the data together in this mariner, in establishing a DNBR limit, is only valid if there are no systematic trends among the results from the various test series.
The WRB-1 correlation appears to predict the CHF tiest data reasonably well; and the mean value and standard deviations of Measured CHF/ Predicted CHF for each individual test series (presented in Table 3-1 of WCAP-8762) appear to be valid indicatiens of the correlation accuracy relative to that test series. None of the 24 test series used a gecmetry which is exactly the same as one of the Westinghouse fuel assembly dt. signs.' Some of the test series j
are very close to the fuel assembly designs while these include geometry variations (8 foot and 14 foot heat'ed lengths for example) which bound the actual values used in the fuel designs. The test series most representative of the Westinghouse 17x17, 14-foot fuel design are:
A.-1, A-2, A-5, A-18, the test series mest representative of the 17x17.12-foot fuel design are: A-1 A-2, A-3, A-4 A-5, A-18, A-19; the test series most representative of the Westinghouse 15x15, R grid,12-foot fuel design are A-6, A-7, A-8, A-9, A-10, 2 '11, 2-12, A-13, A-12, A-15, A-15, A-17 : -.e tes: ieries : s: re:resen a-ive of the Westinghouse 14x14 L grid and 15x15 L grid,12-fcot fuel design are:
A-20, A-21, A-22, A-23, A-21 An analysis of variance was perfer ec to deterni:e '
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' L significant components of variance among these four groups of test series.
The results of this analysis ind,icate that the R-grid data can be treated as a single data set where the variance of the data is represented by two components, -
one component to account for variance within a given test series and a second component to account for the observed variance among the test series.
Similarly, the L-grid data can be treated as a single data set with its own values, of variance within the test series and variance.among the test series. For both groupings of data, the total variance can be ' calculated as follows:
2 2
' Total 2 * "W
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'A where:
2 = variance within the test series about the mean value eg 2
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= variance among the test series means.
g The combined degrees of freedom for these two groupings (i.e., R-grid and L-grid) have been calculated on the basis of a weighted harmonic mean of the degrees of freedom for the data within the test series and for the data accng the test series. The following fornul; wasused:
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combined degrees of freedom e
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variance among the test series means l
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degrees of frei:cm for data amor.g the test series e3ns.
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Table V1 summarizes the results of these calculations. The differences. in j
e and e shown in Table VI for,.the R-grid and the L-grid data demonstrate y
A that a single value of e or e
f r all the data would not be appropriate.
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The cause of the increased variation among the L-grid test series means relative to the R-grid test series means is not fully known.
One possible explanation is that the heated length effect is different for L-grids than for R-grids and that the heated length effect in the WRB-1 correlation is not appropriate for L-grid designs. This speculation is supported by the fact that, for the L-grid tests, the three lowest test series mean values of Measured CHF/ Predicted CHF were all associated with the.eight foot tests and the three highest test series mean values of Measured CHF/ Predicted CHF were associated with the 14-foet tests. This can be seen by comparing the test series means on Figure 2.
In addition, direct comparisons can be made between test series which are i
geometrically identical except for the grid design. There are four cases' in' which these direct ccmparisons can be made.
For the two 8-foot c mparisons, (i.e., test series A-21 vs. A-9 and A-21 vs. A-14) the' L-Grid means were lower than the R-grid means (1.59% and 2.77% respectively). -For the two 14-foot ecmparisons', (i.e., A-22 vs. A-7 and new data frem tests W-206 to
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216 vs. A-6) the L-grid means were higher than the R-grid means (5.58% and 4.47% respectively).. These comparisons a1so. indicate,that the length effect may be different for L-grid tests and R-grid,, tests.
If there is a real difference in the length e'ffect of L-grids, then the length effect in the WRB-1 correlation is conservative since it produces a higher CNBR limit.
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14 Assumotions for PWR Aeolication Having reviewed the WRB-1 correla' ion relative to the CHF data, what remains
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t to be detennined'is the ' applicability to PWR design. Three areas have been I
reviewed relative to PWR application of the WRB.1: first, limitation,s of the I
correlation related to geometry and coolant conditions; second, uncertainties introduced by CHF test atypicalities; and third, establishment of an appropriate DNBR correlation limit.
l The proposed range of' coolant conditions is based on the range of all the f
data.
Based on the empirical nature of the correlation, the range of application should be based on the range of the data.
In tenns of geometry, the correlation should not be applied to any PWR geometry which has not been specifically -
tested or which has not been bracketed by the test data. The important.
parameters to which this applies are: rod size, rod pitch, heated length, mixing vane design and grid spacing.
t The following differences between the CHF test and the Westirighouse PWR fuel designs have been reviewed:
the inclusion of simp.le support grids between the mixing vane grids in the CHF test but not in the actual' fuel design; heated l'ength; number of r6ds, grid spacing, and use stainless stee1~ rods.
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the heated length and grid spacing in the CHF tests.is not the same as is used in Westinghouse PWR designs the values tested were sufficiently close to the
' design values; and the CHF tests also included ranges of heated length and grid spacing whichi cover the range of design values. The question of the use of simple spacer grids is more compic than the other atypicalities.
In the CHF tests the rods are electrically heated and the magnet'ic forces resulting from the electric current could cause the rods to bow in a manner which is not typical of PWR conditions. in order to reduce or eliminate this potential red bowing, the CHF tests include simple support grids to prevent rod motion. These simple support grids were designed to minimize their effect on the local fluid conditions. The flow resistance (K) factors for the simple support grids are :approximately"one third the7talUes
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for the mixing vane'gridsp In addition, previous Westinghouse CHF tests.
(reported in WCAP-7411-1-P-A) were reviewed to deter nine the effect of the simple support grids on the CHF results. The effects of' simple support grids, "T+H mixing vane grids", " mixing vane grids", and " scoop type mixing vane grids" were studied in an eight foot h'eated test section with 10", 20" and 26" grid spacing.
The effects'of simple, minimum resistance grids have also been studied by Babcock & Wilcox (References 6, 7) and Battelle pacifi.c Northwest Laboratories (Reference 8). These reports generally agree that the turbulen' downstream effects of simple grids is. eliminated (' asEed out) within a distance of w
. estinghouse fuel designs use mixing vane grid spacings 8 to 10 inches.
W of greater than 20 inches. The CHF tests ics: closely simulating -he nel
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design grid spacing were done with 20,- 22 and 26-inch grid spacing.
For these tests'the inclusion
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l of simple support grids half-way between the mixing vane grids results in at least a 10-inch distance between the simple support grid and the CHF location which was generally just u> stream of a mixing vane grid. Although there is no directidata on the effect of including simple support grids between l
the mixing vane grids, the existing data on simple support grids are 9
sufficient to indicate that the effect is small or negligible. Therefore no penalty or uncertainty is required ~ in the correlation to account for this atypicality.
The effect of the number of rods in the CHF tests has been reviewed. The CHF tests indicate that CHF occurs on the interior rods. The subchannel analysis of the CHF tests indicates that the enthalpy gradient across the bundle is very small and the potential for CHF being effected by bundle edge effects is therefore minimal.
In addition to these observations, it should be noted that the WRB-1 correlation fits the 3x3 bundle data, the 4x4 bundle data and the 5x5 bundle data reasonably well without the need for any empirical terms related to the, number of rods in the bundle.
Staff positions l
The s'taff has reviewed the subchannel calculations and the other assumptions used in the development of the WRB-1 correlation as well as the atypicalities in the CHF. We conclude that the inclusion of uncertainties for these areas would not significantly1 alter the WRS-1 correTation or the final DNBR correlation limits. Uncertainties in these areas are therefore not required.
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The WRB-1 CHF correlation predicts the CHF test data reasonably well; and the mean value and standard devia ion of the Measured C F/ Predicted CHF for
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each individual test: series are acceptable measures of the accuracy of the correlation. We co'nclude that there is a significant component of variance associated with changes from one test series to another and that, this component of variance must be accounted for in the establishment of a DNBR limit.
An acetJtable method of accounting for this component of variance was previously di.scussed. Table VI pre,sented the total variance and degrees of freedoms based on that method. The DNBR limits presented in Table VII were calculated using the same technique for establishing the DNBR limit
,as proposed by Westinghouse. However, the input values of variance, degrees of freedom and correlatfor, mean were taken from Table VI.
The unusually high DNBR limit of 1.37 for the application of the WRS-1 correlation to L-grid fuel is, in part, due to the limite'd number of L-grid geometries tested.
It may also be the result of a difference l
in the heated length effect for L-grids relative to R-grids, as l
previously discussed. The staff acknowledges that additional data or additional. analytical work on the WRS-1 correlation for L ' grid application could substantially improve the results.
Because of the empirical nature of the correlation and because of the 1
l additional compenent of variance among the test series, as previously discussed, we conclude that the mean v'alue, standard deviation and DNBR limit oresen ed in Table 'l:: are accrooria e easures of the correlati:n l
accuracy; and that the mean value, standited deviation and CNBR limit, preposed by Wes'tinghouse, based en ali 1147 data points are not a:preoriate r
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representations of the correlation accuracy. We find that the WRB-1 correlation is acceptable for use'in PWR thennal-hydraulic design with the DNBR correlation limits specified in Table VII.
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References 1.
Motley, F. E., Hill, K. W., Cadeh, F. F., Shefcheck, J.,
"New Westinghouse. Correlation WRB-1 For Predicting Critical Heat Flux in Rod Bundles with Mixing Vane Grids" WCAP-8762 2.
Rowe, D.
S., " COBRA IIIC: A Digital Computer Program for Steady State and Transient' Thermal-Hydraulic Analysis of Rod Bundle Nuclear Fuel Elements, BNWL-1695, March 1973 3.
Wheeler, C. L., et al, " COBRA IV-I: An Interim Version of COBRA for Thermal-Hydraulic Analysis of Rod Bundle Nuclear Fuel Elements and Cores", BNWL-1962 March 1976.
o 4.
Hochreiter, L. E., and Chelemer, H.. " Application of the THINC-IV Program 1
to PWR Design", WCAP-8054 September 1973.
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Rosal E. R.. Cermak, J. 0., Tong, L.
S., " Rod Bundle Axial Non-Uniform Heat Flux DNB Tests and Data" WCAP-7411-1-P-A, January 1975 6.
Zielhe, L. A., Wilson, R. H., " Transient Critical Heat Flux and Spacer Grid Studies", Nuclear Technology, Volume 24. October 1974 7.
Morgan, C. D., " Correlation of Critical Heat Flux in 'a Bundle by Pressurized Water" BAW-10036, January 1972.
8.
Alberman, R. T., and Rowe, D. S., "An Experimental Study of Turbulent Flow in a Model Babcock & Wilcox Fuel Bundle" PNL Report.?l2 B00763.
o of O
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,C.
TABLE 1 COMPARISON OF LOCAL COOLANT CONDITIONS i~~
Test Difference in Calculated % Ouality*
No.
THINC-COBRA III THINC-COBRA IV W-1796 i
-0.6
-0.6
^
W-17979
-1.0
-1. 0 W-1798
-0.4
-0.4 W-1799
-0.4
-0.3 W-1800
-0.1
+0'.1
~
W-1801
+0.1
+0.1 W-1802
+0.4,
+0.4 W-1803
-0.1
+0.1 W-1804
-0.4 60.4 W-1305
-0.8
-0.8 W~1806
- 0.1
-0.2 W-1807
-0.6
-0.6 W-1808 0.0
-0.5 L
W-1809
-0.7
-0.7 W-1810
-0.3
-0.3 W-1811
-0.3
-0.2 W-1812
-0.2
-0.1 W-1813
-0.2
-0.2
- Difference in % ' Quality =.THINC Quality in % - COS?A Quality in %
l
TABLE 'Il
~
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TYPICAL, LOCAL COOLANT CONDITIONS AT CHF LOCATION FRoM COBRA -
t I
~
MASS Flow Mtn/M FT2 SUDcilt.flilEL ENDdLPY BTU /LB QUALITY %
.a 1
794.8 19.0 2.755 2
793.7 18.8 2.766 3
793.1 18.6 2.773 g.
4 1
792.8 18.5 2.219 5
792.5' 18.4 2.217 o
~
6 792.4 18.4 1.961 A>
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. Typical Comearison of the Test Series Mean Values of Measured CHF/ Predicted CHF Using THINC and COBRA-IV Table III
.1 Predicted Test Measured CHF CHF Standard Deviation Series
.9964
.9910
.0655
.0640 A-3 73 1.0502 1.0360
.1020,
.0932
- Table IV
, Typical Comparison of the Test Series Mean Values of Measured CHF/ Predicted CHF Using COBRA III and COBRA IV Measured CHF Test Predicted CHF Standard Deviation Series
- Pts COBRA III
' COBRA IV COBRA III COBRA IV A-1 71
.9912
.9910,
.0640
.0640 g
- Test gecmetry for each. series is identifie,,d-in Table 3-1 of WCAP-8762 e
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Table V Analysis of Variance Test for CHF Data Theoretical F
~
F Calculated Range at 99%
Description Test Series Values Confidence All Data A-1 A-24 11.2
.6 - 1.8 All R grid A-1 A-19 6.2
.5 - 2.0 All L grid
'A-20 A-24 46.4
.3'- 3.3 14' UNIF A-1. A-2, A-15 15.4
.2 - 4.7 14' COSINE A-5 A-18
.1.05'
.2 - 6.9 14' USINU, A-9 A-10 A-11 R Grid A-12, A-14, A-16 A-17 6.9
.2 - 2.8 14'.USINU L Grid A-21,.A-23, A-24 16.3
.2 - 4.7 14' USINU A-9, A-10, A-11 R&L Grid A-12, A-14 A-16, -
A-17, A-21, A-23 A-24 8.0
.4 - 2.4 8' UNIF A-3, A-19, A-4 2.6
.2 - 4.6 14', 5x5 TYP A-1, A-2, A-5
.2
.2 - 4.6 l
14', 22" GSP A-1, A-5, A-18
.1
.2 - 4.6 14', 4x4, TYP, A-9, A-10, A-11 R-Grid A-12, A-14, A-15 15.7
.3 - 3.0 14', 5x5 A-1 A-2, A-5
.13
.3 - 3.8 14', 5x5, UNIF, TYP A-1, A-2
.4
.2 - 6.9 1
8', 5x5, l
UNIF, TYP A-3.A-4
.7
.2 - 6.9 l
l.
5x5, UNIF, TYP A-1, A-2, A-3, A-4 5.5
.3 - 3.8 I
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ACCEPTABLE PWR FUEL DESIGN, DNBR LIMITS ' : !i
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^ PLIC^
L FUEL DESIGN DNBR LIMIT TS ERIES 6
17 x 17 R GRID A-1, A-2, A-5, A-18
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14'
'LENGTI-I a
17 x 17 R GRID A-1, A-2, A-5, A-18 12' LENGTI-I A-3, A-4, A-19
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1.17 s
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16 x 15 B GRID A-6, A-7, A-8, A-9, 12' LENGTI-I A-10, A-11', A-12, A-13, A-14, A-15, A-16, A-17 j
o 16 x 15 L GRID A-20, A-21, A-22, A-23,
'l 12' LENGTI-I A-24
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1.37 h
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9
WRB-1
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R GRID DATA
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0 O.84 0.88 0.92 0.96 1.00 1.04 1.08 1.12 AVERAGE M/P
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L G RID DATA c! v.
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8' HEATED LENGTHS 14' HEATED LENGTHS c
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0.84 0.88 0.92 0.96 1.00 1.04 1.08 1.12 Ii AVERAGE M/P
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FE3 9079 Mr. Thc:as M. Anderson, Ilanager
!!uclear Safety Depart:nent
'lestinghouse Electric Corporation P. O. Box 355 Pittsburgn, Pennsylvania 15230
Dear Mr. Anderson:
SUBJECT:
SAFETT EVALUATION OF WCAP-8720 c"
The fluclear Regulatory Comission staff has completed its review of Westinghousa Electric Corporation topical reports WCAP-8720 (Proprietary) and 1! CAP-8785 (Non-proprietary) entitled "In: proved Analytical Models Used in Westinghouse Fuel Rod Desi5n Cccputations." Our safety evaluatim is enclosed.
1 CAP-8720 describes endifications to the Westinghouse PAD c:xsputer code, which is used to calculate fuel perfomance for steady state reactor operation and to detemine' initial fuel red gonditions for postulated transients and accidents, including the lossedf coolant accident. The modified code, designated PAD-3.3, supersedes the previous venion, cesignated PAD-3.1.
As a result of our review of WCAP-8720 and supplemental information provided during the review, we have concluded that FAD-3.3 is an acceptable code for use in safety analyses of license appitcations provided the folicting conditions are siet.
1.
The use of the empirical ecuation for fuel pellet-to-cladding gap conductance (Ecuation 9 in WCAP-87EC) is restricted to analyses of fuel perfomance for steady state and slow transient operations (startup, load following, and shut down) and is fu to gas conductivities greater than 0.03 Stu/hr-ft fJ1er restricted F.
I 2.
The use of the revised fission gas release mdel is restricted to analyses of fuel perfor=ance for steady stata and slow transient j
operations.
l 3.
The use of Equatiens 10 and 12 in WCAP-8720 for gap conductance i
is're x 10'gtricted to effectiv,g surface rougnness in the range of 14.4 feet to 22.4 x 10 feet.
l 4
Uncertainties in the cladding creep.model art accounted for by a
-l 30t reductim in computed c:scoressive creep values and by a 20:
' 2 :::: "
- ::: ::.= 7: :.:G,2..;.
l Se'% D
.VAC PCRM M8 IMS).MRCM 0240 1lt u. a. seva-mam==m emse. son.ese eae t
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Mr. Thorr.as M. Anderson FE3 9 ;;;g Cecause the previous venion of the code, PAD-3.1, contains neither the improvements nor the restrictions discussad in this safety evaluation, future fuel perfomance analyses must be done with the revised version of the code, PAD-3.3. We understand you c:ay propose that some of the restrictions be dropped fract our approval on the basis of new infonr.a-tion.- If such a proposal is submitted as an amend:nent to WCAP-8720 and WCAP-8785, and if it is judged appropriate for safety analyses, we will document such conclastons in a future supolement to the enclosed safety evaluation.
Fuel matarial properties used in the code am reported in WCAP-9179, which is currently under review. Restrictions or changes recuired by our approval of WCAP-9179 will also apply to VCAP-8720 and WCAP-8735, where appropriate. The use of the code to calculate initial conditions for transients and accidents will be reviewed for each tyce of transient or accident to assure that code input values result in acceptable conservatism.
Tepical reports WCAP-8720~ and WCAP-8785 must be revised to include the MRC acceptance letter, the enclosed safety evaluation, and supplemental infomation resulting frem the review. The revised version should incorporate referinces 2, 4 and 7 of our enclosed safety evaluation in order.to have a complete description of the PAD-3.3 Code. The revised versions of WCAP-8720 and WCAP-8.785 will be acceptable for reference in license applications.
In accordance with established procedure, it is requested that Westinghouse issue the revised versions of these reports within three months of receipt of this letter.
'a We do not intend to repeat our review of these recorts when they appear as references in a particuhr license application axcept to assure that the material presented in these reports is applicable to the specific plant involved.
Should Nuclear Regulatory Commission criteria or re'gulations change, such that cu'r conclusions concaming these reports are invalidated, you will be notified and given an opportunity to revise and resut:mit your topical reports, should you so desire.
Sincerely, I
Crigi sl :ig=ed. ty k h: T.Stol:
John F. Stolz, Chief Light Water Reactors Branch No. 1 Division of Project Management
. -. +
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c
'C ENCLOSURE-l SAFETY EVALUATION 0F THE 7-WESTINGHOUSE ELECTRIC CORPORATION TOPICAL REPORT WCAP-8720 IMEROVED ANALYTICAL MODELS USED IN WESTINGHOUSE FUEL ROD DESIGN COMPUTATIONS psREGy
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.s DECEMBER 1978 CORE PERFORMANCE BRANCH UNITED STATES NUCLEAR REGULAT",AY COMMISSION WASHINGTP '
J. C. 21lE38 e
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CONTENTS I
Section Page 1.
Introduction........................
...................... 1 2.
The rmal Performance Model Revi sions........................ 3 2.1. Helium Thermal Conductivity......:.....................
3 3[
2.2 Accommodati on Coe f fici ent........ :..................... 4
~..
n 2.3 Gap Conductance........................................ 5
?s -':.
r 2.4 Helium Solubility...................................... 10
~
- .2-3 2.5 Fi s si on Ga s Rel ea se.................................... 11
.i 2.6 Swelling and Densification.............................
14
- ?I" 2.6.1 Soli d Swelling and Densi fication................15 T,
2.6.2.F i s sion Gas Swe11 i ng............................ 17 2.7 Mate ri al P rope rti es.................................... 18 3.
Majo r Con siderations i n the Revi ew.......................... 19 3.1 The rmal Perfo rmance Predi ctions........................ 19 3.2 Mechanical Perfo rmance Predictions..................... 21 3.2.1 Fuel Rod Internal Gas Pressure.................. 22 3.2.2 Cladding Creep Correlation...................... 22 3.3 Previous and Revised Model Comparison.................. 27 3.4 Model. Applications.....................................
28
- 4. Conclusions................................................. 31 i
j
- 5. ' References.................................................. 33 c
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2.3 Gap Conductance A variety of gap conductance relationships is used in the Westing-house fuel performance code. Three of these relationships are proposed for use in safety analyses, each to become effectivEunder different conditions.
The other relationships are used only in fuel design studies and are not intended for use in safety analyses. The cquations* used are (1) an i
empirical correlation used for open gaps and identified as Equation 9; (2) a simple annular gap model identified as Equation 10, which is also used for open gaps; and (3) an equation used when the pellet is in contact with the cladding and identified as Equation 12.
If the pellet-to-cladding gap is closed, Equation 12 is used to calculate gap conductance.
If the pellet-to-cladding gap is open, the gap conductance is calculated with both Equations 9 and 10; the greater of the two resulting values is used in the final analysis. Some restrictions on the use of all of these equations are needed in order to assure predictions that we judge to be acceptable. These restrictions are described below.
The empirical gap conductance correlation, Equation 9, was derived from thermocouple and melt radius data for use with finite gaps. The equation implicitly assumes some measure of pellet relocation and is similar to the gap conductance model proposed by Broughton and MacDonald l
(H). We have noted two anomalies resulting from the use of this i
ertuation.
?
t
- Equations 9, 10, and 12 in WCAP-8720.
(
(
The first anomaly arises when a rod is filled with a low-thernal-conductivity gas, such as xenon. In that case the Westinghouse model predicts greater gap conductance just before fuel and cladding contact occurs (Equation 9) than just after contact occurs (Equation 12); this is an anomalous prediction. Several remedies for this situation have been considered. Modification or elimination of this empirical ecuation would require rederivation of the fission gas release and other models in PAD-3.3, because Equation 9 was used to generate these models. West-inghouse has, therefore, agreed to limit the application of Equation 9 to a restricted range of gas conductivity. Within this concuctivity range, the gap conductance is consistent with the other gap conductar.ce models and exhibits no obvious anomalies in behavior. The resulting predictions are also consistent with the octa presented by Westinghouse.
The use of Equation 9 is therefore acceptable for use in safety analyses when limited to values of gas conductivity greater than 0.03 Btu /hr-f t*F.
The second anonaly results from the nonlinear decendence of gap conductivity on gap size, which in turn is typical of a gap conductance model that implicitly includes the effects of relocation. In this case gap conductance values do not go to zero as they should when gap sizes become very large.
(
Over a restricted range of gap sizes, Equation 9 is a reasonable relation. The data used to develop this equation were taken fron WAPD-228 (25). These data were derived from in-pile irradiation of 11 capsules, including both helium and fission gas-filled rods with a large range of e.
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cold gap sizes. We believe thai; the analysis of the data is valid and that a suitable form of the gap conductance function has been incorporated into Equation 9.
However, the constants in this equation were estimated from trial fits of the data and calculated hot gap size, and given the same data base, it would be difficult to independently reproduce these constants. Equation 9 has also been used to predict centerline tempera-ii tures of other experiments (H, 2],).
These data show reasonable agreement I
with measured and predicted values of temperature over a range of fuel rod parameters. We note further that several other fuel vendors have proposed j
the use of annular gap conductance models in conjunction with independent relocation models rather than t'he combination of both effects proposed by Westinghouse.
The Westinghouse empirical gap conductance model compares f avorably with several non-proprietary methods (24,18,, g). We thus 8
conclude that the use of Equation 9 is acceptable over a restricted range of gap sizes.
i To accommodate this second ancmaly of Equation 9, two restrictions were considered and rejected. These restrictions were (1) the elimination I
of Equation 9 and total reliance on Equation 10, and (2) the application l
i of limits on gap size dimensions. Since Equation 9 had been utilized in the derivation of fission gas release and other models in PAD-3.3, the use of either of these restrictions would necessitate the rederi-vation of a number of models.
As an alternative, Westinghouse has shown that Equation.9 is used only at the rod ends where a large gap, low conductivity condition exists t
(
(
at high burnups. Equation 10, the acceptable (see below) annular gap model, is used for the middle of the rod. It is the center section of the rod, which yields the highest value of, stored energy and fuel tempera-tures, that is limiting for LOCA analysis. The,use of Equation 9 is there-fore acceptable for gap sizes anticipated with the present Westinghouse fuel design during nomal operation because it is only indirectly used in safety analyses. Our approval includes application for steady-state and maneuvering conditions (i.e., start-up, load-following, and shut-down).
In the case of some accidents, such as the large-break LOCA, the fuel-to-cladding gap increases and the use af Equation 9 may be inappropriate in predicting gap conductance. Therefore, we do not approve of the use of this equation for the analysis of transients and accidents where the fuel,
to-cladding gap is predicted to increase. In this regard, we have re-examined the gap conductance models used in LOCA analysis (30,, 31) and i
find that Equation 9 is not used.
I Equation 9, as used in PAD-3.3, is essentially the same as the empirical gap conductance equation used in the previous version of the code, PAD-3.1.
Because the previous version does not contain all of the restrictions described in this safety evaluation, we no longer consider PAD-3.1 acceptable for reference in future license applications.
f The annular gap conductance correlation, Equation 10, is used to calculate open gap conductance as a function of the effective gap and themal conductivity of the fjll gas. In its limit, this relation has f
a pressure-dependent contact conductance tem as given by Equation 12.
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The later relation is based on the work of Dean (32) and Ross and Stoute I
(H). Gap conductance relations similar to Equations 10 and 12 are widely used in nuclear fuel performance analyses. The use of these equations in safety analyses is therefore acceptable provideFthat appropriate values of (1) gas conductivity, (2) surface roughness, and (3) gap width or contact pressure (when the gap is closed) are used.
Modifications to the Westinghouse relationship for item (1), gas conductivity, were reviewed in Sections 2.1 and 2.2 of this report and found to be acceptable. The values for item (2), surface roughness, are based largely on experimental data. Where no data exist, it is appro-priate to restrict the roughness values to realistic, non-zero values.
Westinghouse has agreed with this staff conclusion. They have stated that a value of 22.4x10~0 feet is currently used in all safety analyses l
for the' effective surface roughness of the fuel and cladding.,It is thus agreed (and required) that the roughness value will be bounded by ah
-6 upper limit of 22.4x10 feet and a lower limit of 14.4x10'0 feet for all i
safety analyses.
.i In considering item (3), gap width or contact pressure (when the I
gap is closed), we have nr.*.ed that no cracking and relocation model is present in the Westinghouse analysis. As a result, the values of gap width or contact pressure will be conservatively estimated. There fore, we have concluded that the use of the annular-gap (Equation 10) and closed gap (Equation 12) models are appropriate for safety analyses, given the previously described restriction on surface roughness.
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We have examined the data and note that the revised solubility squation conforms to these data over a wide range of exposures. The equation extrapolates to a non-zero helium content in the fuel at zero exposure. The overall analysis, however, assumes no helium content in the fuel at the time the rod is backfilled and sealed. The model thus indicates that some helium is absorbed after filling but before loading. We have questioned the reality of this reduction in backfill pressure at zero exposure. Westinghouse responded with evidence of I
insignificant absorption of helium during prepressurization and has j
provided further calculations to show that the predicted reouction in e
the initial rod pressure is small and unimpartant. The use of the i
revised helium solubility equation is therefore acceptable for safety analysis.
2.5 Fission Gas Release The revised fission gas release model constitutes a major revision in the Westinghouse fuel thennal performance analysis. The previous fission gas release model (3_, 34) was developed using data obtained from low-burnup fuel. As more data from high burnup fuel became i
available, Westinghouse found that their previous model underpredicted the " measured f'ission gas release in high-burnup fuel rocs. Consequently, a new fission gas release model was developed. The revised model is based upon the observation that fission gas release increases as a parabolic function of burnup (3_5) in fuel rods operated at nearly 5
constant power and constant temperature..
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In support of the revised equation, Westinghouse submitted data covering a wide range of design parameters, including burnups in excess of 57,000 mwd /t, and measured fission gas release fr.:,ctions as high as 40".. The data were taken from fuel rods that are. similar to, but not always identical to, the Westinghouse standard product line. For example, supporting data were taken from plutonium-bearing Saxton II fuel rods, which are not a standard product. These data were utilized because high-burnup fission gas release data are difficult! to generate and do not exist in profusion. Although there has been some speculation that mixed-oxide fuel release more gas (3_6) than UO fuels, no major differences between 2
releasas from power reactor, test reactor, and mixed-oxide fuels have been demonstrated.
We therefore accept all of the data as being applicable.
We have also compared the Westinghouse gas release model with similar models proposed by the NRC (3]7,) and the ANS-5.4 working group (36). These models are base'd on both LWR (38) and LMFBR (39) data, which include plutonium-bearing fuels. Results indicate that the Westinghouse gas release modeT is applicable to the standard product line.
We have examined the derivation of the burnup-dependent fission
- gas release model described in Appendix 8 of.WCAP-8720. We find two limiting assumptions in its derivation. The first of these is that, for a fuel rod subjected to a varying power operation, the power history may be approximated as a series of steps, each at a constant power level. As smaller and smaller discrete time-steps are used (at greater and greater computational cost), the approximation of a continuoutly varying power
(~
b level becomes better. The discrete time-step approach to varying power operation is a comonly used method in fuel rod analysis. A limitation of this method is that inaccuracies result when very large time-steps f
are used. Westinghouse has shown that a reasonatrie number of time-steps j
can be accommodated in the analysis. Westinghouse has also chosen suffi-ciently small time steps in all calculations we have audited. We therefore conclude that adeouate attention has been paid to this consideration.
The second limiting assumption in the Westinghouse fission gas release model is that the fission gas released during any increnental burnup period during varying power operation is the same as the fission gas that would have been released during that period had the fuel been operating at the same power and temperature over the entire operating history. This assumption is conservative for periods of decending power operation, but it is nonconservative for periods of ascending power operation. The assumption is particularly troublesome for large variations in power history such as those occurring in transient analysis. For fuel rods operated at nearly con-I stant power and temperature, as was the case for the supporting data, the Westinghouse model appears to be adequate. We thus approve of the use of the revised fission gas release model for fuel performance analyses cnly i
during normal operation. Normal operation in this sense includes steady-l e.
state and maneuvering conditions (i.e., start-up, Ioad-following, ano shut-down) since the supporting data were based on these conditions. We l
do not approve of the use of this model for gas release predictions during l
1 transient and accident conditions where fuel temperatures increase rapidly
- i I
l
as a function of time. The fission gas release model is, of course, accep-table for use in analyzing pre-transient and pre-accident fuel conditions.
It is also acceptable for use in analyzing transient and accident conditions where fuel temperatures monotonically decrease'as a function of time.
i The limitations identified in the proposed fission gas release model (PAD-3.3) are also apparent in the previous version of the code (PAD-3.1).
We further note that the previously approved model does not consider the t
effects of burnup enhancement. Because the previous version of the gas release model contains neither the improvements nor the restrictions
{
discussed in this safety evaluation, we no longer consider PAD-3.1 acceptable for reference in future license applications.
l 2.6 Swelling and Densification Tiie dimensional changes of a fuel pellet during irradiation are primarily detemined by themal expansion, swelling, and densifi-cation characteristics of the fuel. These dimensional changes are subsequently used in detemining the heat transfer between the fuel and cladding, the linear heat genstration rate, and also for possible power spiking in the fuel column. The themal expansion model for the fuel pellet has been described previously (3,, 4).
Tha swelling and densification models (4_, 5, 9_) have been revised in WCAP-8720 to better match incore stack height measurements, mercury pycnometry and camna scan data presented by Westinghouse. The solid swelling and densifi-cation model, which describes fuel shrinkage and dispersed fission
' product swelling, is reviewed in the follcwing section.
. l
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2.6.1 Solid Swelling and Densification The Westinghouse model of time-dependent incore fuel densification is entirely empirical. No theoretical description of densification mechanisms has been utilized. From examination 1 f irradiated fuels, Westinghouse has concluded that the amount of fuel densification during irradiation decreases with increasing fabrication sintering temperature, and, consequently,- they have chosen sintering temperature as the key variable in the model..
The empirical correlation, Equation 18 in WCAP-8720, contains the sintering temperature, initial pellet density, burnup, and fuel swelling and densification rate constants. This correlation for the fuel column fractional length change (a!./L ) is a best fit to a large g
number of column length changes measured in Westinghouse pressurized fuel in' operating PWRs. The form of this equation remains unchanged I
l from the previously approved version of the model (9), but the values l
of the fuel. swelling and densification rate constants have been changed.
The revised constants are based on post-irradiation examination of Zorita fuel rods. All data were screened to exclude plastic flow and gaseous swelling effects in the fuel. The remaining data were used to generate fuel density and pore removal rate as a function of exposure. The later quantity, pore removal rate, was obtained by point-counting metallographic specimens of irradiated pellets. The solid swelling rate constant is determined from the difference between the pore removal and densification rates. The densification rate 3.-------m.y-.
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constant is then redetermined by using data presented with the previous densification model (9). For Westinghouse fuel, the pellet fabrication process has re.iained essentially unchanged for many years with the.ex-ception of increases in the sintering temperatufiand nominal pellet density.
Thus, data obtained from the previous programs are directly relevant to i
current safety analysis methods. The resulting solid swelling and f
densification model was described and compared with the Westinghouse measured in-core stack length change data. The results show that the revised model is generally conservative. Good agreement was also obtained when the revised model was applied to recently published EPRI data (40).
In considering early-in-life fuel temperatures, Westinghouse has shown the revised rate constants to be more conservative than the previous constants. WCAP-8720 shows the predicted change in fuel stack length using both the old and new rate constants. From beginning-of-life to a significantly high burnup value, the predicted fuel stack length change is greater for the revised constants than for the previous constants.
During the time period when maximum fuel temperature Occurs, th'e use of the revised rate constants result in a small additional amount of densification. This, in turn, results in a greater fuel-to-cladding gap and higher fuel temperatures. On this basis, we find the revised solid swelling and densification model acceptable.
i f
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C 2.6.2 Fission Gas Swelling Fuel densification begins on first rise to power and is essen-tially complete for Westinghouse fuel at a burnup of approximately 5,000 mwd /t. Athigherburnups,fuelswellingEicomesthedominant
{
mechanism in detemining the current dimensions of the fuel pellet.
The Westinghouse code assumes fuel swelling exists in essentially two forms. Solid swelling, which was reviewed in Section 2.6.1, describes a fine dispersion of gaseous and solid fission products within the fuel matrix. The second form of fuel swelling, called gaseous swelling, occurs when small gas bubbles coalesce to fom larger bubbles along grain boundaries. The second form, gaseous swelling, is now considered.
The previous fission gas swelling model assumed three regions within the fuel:
(1) a region below 800*C, (2) a region between 800*C and the restructuring temperature, and (3) a region above the restructuring temperature. Only solid swelling was allowed in the fuel region below 800*C. whereas both solid and gaseous swelling were accounted for in the fuel regions above 800*C.
In the revised model, only two fuel regions are considered; one above and one below the. restructuring temperature. Gaseous swelling is considered only in the fuel region above the restructuring temperature.
In reviewing the revised gaseous swelling model, we note that t
the exposure level where the gaseous swelling effect is saturated 17-r-
7
,-----7 r---
m m
I k.
(i.e., becomes insensitive to burnup) has been decreased from the previous model. However.,he dependence of gaseous swelling on temperature has also been reduced. As a consequence, swelling becomes significant at lower burnups, but is reduced in.megnitude. The revised model also allows for a reduction in the amount of gas bubble swelling at high temperatures. This was not the case with the previous version of the code. The cummulative effect is conservative with respect to the l
previous model. The revision is therefore approved for safety analyses.
2.7 Material Properties Several of the material properties contained in the Westinghouse fuel performance code have been revised. These include thermal expansion, Poisson's ratio', and Young's modulus for UO2 and stainless steel. We have noted some errors in the tabulation of these properties, which have been corrected by Westinghouse. The property.evisions do not res11t in significant changes 'to parameters of licensing interest. The changes are therefore acceptable for use in the Westinghouse safety analysis.
.We further note that the previous and revised documentation of PAD is not adequate for all material properties considered in the code. A more detailed description of material properties used in Westinghouse safety analysis can be found in WCAP-9179 (41). This topical report is currently under review.
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3.
Major Consideration in the Review In reviewing WCAP-8720, we have not attempted to reexamine all of the previous information (1-9) submitted in support of earlier versions of the Westinghouse fuel performance cdd'e. Our review has been limited to those changes described in WCAP-8720 and to those unchanged features of the code impacted by new data or other new analyses submitted by Westinghouse. Nevertheless, e ference 2, 4, e
and Attachments 1-14 of Reference 7 should be. included in the approved version of WCAP-8720 in order to consolidate the description of PAD-3.3.
3.1 Thermal Performance Predictions One of the major applications of the Westinghouse fuel performance code is to calculate the fuel temperatures or stored energy at the initiation of a LOCA. These calculations are required by 10 CFR 50, Appendix K, and the results of these thermal performance predictions are changed by the code revisions described in WCAP-8720.
I The overall capability of the Westinghouse thermal performance model I
~has been demonstrated by comparison of measured and predicted temperatures from test rods. The comparisons, which are shown in Appendix A of WCAP-8720, i
1 are based on thermocouple measurements from WAPD-228 (25), HPR-80 (42),
AE-318 (43), and IFA-226 (44). The results show that the predicted and measured fuel temperatures compare well with each other.
I We note that this comparative study was restricted to helium-filled i
i rods because they are more representative of the Westinghouse standard
(
(
product line. Although this is true for rods with low burnup, we believe that high-power, high-burnup rods of Westinghouse design are more closely represented by rods containing a mixture 'of fission gases rather than pure helium.
In the bounding sense, such c5mpar' sons should include xenon-filled rods. We requested such data (J6) and Westinghouse responded (13) with results o' f fission-gas-filled rods from WAPD-228 (25),
a WCAP-2923 (26), CVTR (27), and IFA-431 (45). These results show reasonable agreement between measured and predicted fuel temperatures for rods des-cribed in the first three reports, and the predictions are generally more conservative at higher power levels. The predicted value for gap conduc-tance in the IFA-431 case, however, is considerably larger than the published value. Westinghouse attributes this anomaly to a possible lack of relocation effects in the experiment. We agree with this possi-bility and find the Westinghouse thermal performance model predictions in agreement with the remaining helium-and fission-gas-filled-roa data.
f We also note that the Westinghouse predicted and measured centerline fuel temperature comparisons are restricted to low burnups.
In the case of HpR-80, data indicate a poorer agreement between prediction and experiment at 4300 mwd /t than at beginning of life.
In response to our concern about this discrepancy, Westinghouse pointed out a possible experimental bias in the HPR-80 data resulting from irradiation-induced decalibration of the thermoccuples. We agree l
with this explanation and further conclude that experimental qualifi-l cation of thermal performance predictions by direct methods may be l
C.
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impossible at high burnup. At high burnups, indirect indications (e.g., fission gas release) of fuel temperature must be used. At 4
lower burnups, where fuel temperatures are maximum, the Westinghouse thermal performance model has been adequately ver.ified.
i 3.2 Mechanical Performance Predictions The second major consideration in the review of WCAP-8720 is the ability of the revised Westinghouse model to predict the mechanical behavior of the fuel at high burnups. The predictive capability of the code at high exposure is of concern because the burnup levels now considered are well beyond those anticipated with the previous version of PAD. We note that Westinghouse has submitted data in support of WCAP-8720 with ' peak pellet burnups in excess of 57,000 mwd /t and has irradiated fuel rods to burnups in excess of 65,000 mwd /t (35).
The consequences of these high exposures on fuel performance are i
i not well understood because of the limited data available. We recognize, however, the need to make performance predictions at these burnup I
level s.
We have limited our review to two recently identified safety criteria that are effective at high exposures:
(1) maintaining the rod internal gas pressure below the coolant system pressure, and (2) requiring acceptable cladding behavior when the first condition is not met.
l l
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i 3.2.1 Fuel Rod Internal Gas Pressure The current NRC acceptance criteria (M) for the fuel system design basis states that " fuel... rod internal gas pressures should.
remain below the nominal system pressure during'iiormal operation unless otherwise justified." This criterion, which limits reactor operation at very high-burnup, is mainly dependent on the release of fission gas to the fuel-to-cladding gap. The Westinghouse fission gas release model was discussed in Section 2.5 of this report. The other parameter that is used to establish the fuel rod internal gas pressure is the available free volume within the rod for the released gases.
The process by which the free volume within'the rod is calculated remains largely unchanged from the previous version of the Westinghouse code. The calculational method for the initial internal rod volume has been re-examined and a number of questions have been raised on the manner in which the volume changes during reactor operation. We have reviewed the responses to these questions and examined the experimental data submitted in support of the Westinghouse analysis. We conclude that the fuel rod internal l
void volume calculations, and therefore the fuel rod internal gas pressure calculation, remain appropriate at high burnup.
3.2.2 Cladding Creep Correlation The NRC acceptance criterion (M) for fuel rod internal pressure was recently modified (4],) for Westinghouse fuel to allow normal i
reactor operation with fuel red internal pressure exceeding system Therevisedcriterion(g)statesthattheinternalpressure pressure.
of the lead rod in the reactor will be limited to a value below that which could cause (1) the diametral gap to increase due to outward t
cladding creep during steady-state operation and (2) extensive DNB I
propagation to occur. Compliance with the first condition -is deter-mined with a steady-state fuel performance code (i.e., with PAD-3.3),
whereas the second condition is determined with various transient
!(
analysis codes.
l' The process by which the gap size increases may be explained as follows: At high burnup, the released fission gas may cause fuel rod internal pressures to become higher than the reactor coolant pressure (approximately 2250 psi). After fuel-to-cladding contact, fuel swelling i
and fuel hot-pressing processes also contribute to internal pressures.
The total pressure differential results in a driving force that causes the cladding to deform plastically. As the cladding creeps outward, the gap increases. At the same time, however, the fuel continues to swell as a j
I function of burnup. Compliance with the revised rod pressure criterion l
becomes a matter of detemining whether the cladding creeps more rapidly I
than the fuel swells. We have discussed the Westinghouse fuel swelling models in Section 2.6 and the rod internal pressure calculations in Section 3.2.1 of this report. The remaining item of interest is the cladding creep behavior of Westinghouse fuel rods.
t l
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The correlation for modeling Zircaloy cledding creep in the PAD-3.3
[
code is identical to that originally approved for the PAD-3.1 code.
- However, Westinghouse has now requested that the range of.the correlation's,applica-bility be extended to include applications involving tensile creep.*
The Westinghouse cladding creep correlation is based on a theoretical model proposed by Nichols_(49). The correlation predicts a total creep t
rate equal to the algebraic sum of the individual creep rates for (1) thermal i
i creep, (2) compressive creep due to stress-free irradiation growth, and (3) irradiation-induced creep. Volume-average effective stress, which is used to calculate creep rate, is determined in a manner that is consistent i
with the von Mises criterion for isotropic material behavior in'multiaxial l
stress states.
The thermal creep component of the correlation is empirically derived from out-of-pile tensile creep tests in which the cladding was internally pressurized to effectivi-stress levels that rattged from 4 to 30 kpsi at temperatures between 600 to 775'F. There is reason to believe that most
.of the strain measured during these tests was due to primary creep.
Westinghouse compared the resulting thermal creep component to a limited amount of compressive creep strain data obtained from externally pressurized Zircaloy tubing. The comparison shows that compressive creep strains are somewhat overpredicted by the thermal component of l
l the Westinghouse correlation.
l
- For the purpose of this discussion, tensile creep will be cefined as creep-
. out, meaning that the cladding outer diameter increases; conversely, compressive creep implies a reduction in cladding outer diameter.
1
?
The second component, creep due to stress-free irradiation growth, is also empirically derived. Axial growth measurements from a arge number of Westinghouse commercial reactor fuel rods were considered.
The derivation of the irradiation growth componeYt was then completed by analyzing the crystallographic orientation of Westinghouse production cladding as detemined by inverse pole figures. We accept this component of the creep model as a best-estimate representation of a substantial and appropriate data base.
To determine the the third component, irradiation-induced creep, it was necessary to detemine the the total in-pile creep. Profilometry j
measurements from irradiated fuel rods were compared to profilometry measurements on.unirradiated cladding to detemine the total creep. Only
~~
measurements from fuel rods that were predicted to be below system pressure were included in this data base..The data were also screened to eliminate pellet-cladding-meceanical-interaction (PCMI) ef fects. Additional measure-ments from non-reference Westinghouse cladding, failed rods, rod ends, and rods with crud deposition layers were not considered. The final data, which include some 1200 rod measurements, were used to derive the irradiation-l induced creep component. The resultant overall creep correlation is a best-estimate function.
In examining the creep correlation, we have raised questions on I
several areas not considered in the cricinal model description (2).
3 These questions concerned the cladding fabrication processes, cladding I
l material properties, and other phenomena that may affect creep behavior. '
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Strain hardening, strain aging, and anisotropic behavior such as Bauschinger or strength differential effects were considered.
In addition, we examined the Westinghouse creep correlation for agreement with recent data on cladding ovalization, wall 4bickening, and surface undulations (10,11,12,13,14), although the Westin;nouse creep correlation does not consider these effects.
To resolve our questions, audit calculations were performed to judge the predictive capability of the Westinghouse cladding creep model.
These audits were designed to compare the predicted Westinghouse creep rates, total strains, and related cladding behavior with the predictions b
of other proprietary and nonproprietary creep models. Based on these i
audit calculations, we conclude that the Westinghouse model is appropriate for calculating cladding dimensional changes in plant safety analyses so long as.an allowance is made to accommodate data scatter. In the previous version of PAD, where tensile creep was not considered, a 30% reduction in l
the predicted cladding creep rate was conservatively assumed in order to l
l compensate for uncertainties in the data base and the manner in which the
[
i correlation was applied. We believe that, at very high burnups, an increase i
in the fuel-to-cladding gap may lead to limiting values of fuel temperature and stored energy.
It is therefore important to conservatively establish the conditions wherein the gap criterion is violated. We conclude that l
compensation for uncertainties in the Westinghouse data base and correlation application should be applied during conditions of both creep-down and creep-s out. As a result of this conclusion, we have requested (and Westinghouse e.__
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has agreed) that the sense of the conservatism should be reversed if tensile t
rather than compressive creep is considered. That is, if the internal rod pressure exceeds nominal system pressure, the resulting outward cladding creep predictions should be increased, rather tiran decreased, by 30%. We find that this conservatism is the' only revision to the proposed creep correlation that is needed to assure that the Westinghouse creep model is
+
adequate for current safety analyses.
3.3 Previous and Revised Model Comparison At the time the revised Westinghouse fuel per'formance code was sub-mitted for review, the NRC recognized the importance of burnup-enhanced fission gas release in safety analyses. In hopes of providing a basis for interim approval of the revi, sed model, we requested (M) comparative cal-culations for the previous (PAD-3.1) and revised (PAD-3.3) fuel codes'.
Westinghouse responded (J1) to triis request with a set of fuel performance predictions for typical peak and average power pellet conditions. The results were generated for a range of fuel parameters, which included both 1
15x15 and 17x17 Westinghouse fuel designs.
We examined these comparative calculations in detail. In several 1
l are,as, such a.s rod internal gas pressure, the predictions of the revised code were conservative (i.e., higher) than those calculated j
i with the previous version of the code.
In other critical areas, such as gap conductance and fuel temperatures, the revisec code was g
shown to be less conservative than the previous version of tne code.
1 i i i
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' Because these calculations failed to provide a basis from which to judge.
the appropriateness of a reduction in conservative margin, interim approval of WCAP-8720 was not granted.
In spite of the fact that the original objhEtive of the comparative calculations was not met, the impact of these results on the subsequent review process should.not be underestimated. A number of conclusions have been drawn that are based, for the most part, on the comparative predictions of the previous and revised versions of PAD. Among these is our finding that the revised Westinghouse gas release model is inappropriate for transient and accident predictions (see Section 2.1).
The comparison of predictions made with the previous and revised PAD code revis' ions should therefore be considered as part of the formal documentation submitted in support of WCAP-8720.
3.4 Model Applications The Westinghouse PAD code is a computational tool for evaluating i
fuel rod perform.ance.
It is used for both fuel design and reactor
' safety (i.e., regulatory) calculations. Because of the special nature of the safety evaluations, conservatisms are generally included in the safety calculations that are not used for the fuel rod design calculations. These conservatisms may include (1) margins applied to the input of the code (2) margins applied to the output of the code and (3) margins within the models themselves.
It should be noted that not 8
all available conservatisms are applied in each safety analysis, nor i
I e j
should they be. A margin may be conservative in one particular appli-l cation, but ncnconservative in another.
l 4
Since the majority of safety analyses performed with the PAD 7.
i code are concerned with the thermal conditions, a number of model conservatisms are activated within the code for licensing applications where fuel temperatures are required. These include isotropic densi-
{
I t
fication, a subset of available gap conductance models, and other in-ternal margins or limits, Additional compensation for uncertainties (55) is also added to the output of the fuel performance code when the l'
temperatures are used for LOCA analyses. The result of these individual t
conservatisms is a conservatively high estimate of fuel temperatures and stored energy.
4 For the case of a fuel performance parameter such as gap conduc ~
i
~
,g tance, the sense of the conservatism is not clear. 'For power-increasing j
transients, it is conservative to assume a low Doppler coefficient for the fuel. This could be achieved with a high value of gap conductance.
For power-decreasing transients, on the other hand, it is conservative j
to assume a high Doppler coefficient for the. fuel. This could be l
l achieved with a low value of gap conductance. There are many competing l
mecnanisms in' action during a hypothetical transient or accident. Bounding conditions can only be established by means of sensitivity studies con-ducted for the particular event being considered. Although we will provide some guidance in this area, we believe that the application of the Westinghouse fuel performance code is beyond the scope of this review.
-- +, C.~
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The manner in which the code is used, tnat is, the validity of the input and use of the output values, should be considered in the review of specific transient or accident analysis.
1
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4.
Conclusions j
We have reviewed the Westinghouse report, " Improved Analytical Models used in Westinghouse Fuel Rod Design Computations " WCAP-8720.
Based on the information in that report and subsequent information 1
i.
provided by Westinghouse, we have arrived at the following conclusions:
i 1.
The empirical gap conductance equation (Equation 9 in WCAP-8720) should be restricted to analysis of fuel performance during normal operation and should further be limited to values of gas conductivity greater than 0.03 Btu /ft*F.
2.
The revised fission gas release model should be restricted to analysis of fuel performance during normal operation.
3.
The value of surface roughness for the fuel and cladding should be limited to values between 14.4x10-6 and 22.4x10-6 feet.
4.
A 30% reduction in compressive creep of the cladding and a 30t increase in the tensile creep of the cladding should be used in order to compensate for uncertainties l
in the Westinghouse data base and model application.
The restrictions required by our evaluation are the results of deficiencies appearing in PAD-3.3 predictions under transient and high burnup conditions. We further note that these deficiencies also appear in the previous version of the code. PAD-3.1.
Because
. l l
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t the previous version does not contain all of the restrictions described above, we no longer consider PAD-3.1 acceptable for reference in future license applications.
-r.
Based on our review, we conclude that, when the revised fuel perfomance g
model PAD-3.3 is amended as described above, WCAP-8720 will be accep-table for reference in license applications.
l l
9 9
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t 5.
REFERENCES 1.
Letter from R. Salvatori, Westinghouse, to D. Knuth AEC, Entitled: Creep Collapse Parametric Analysis, NS-SL-483, date,d November 6,1972.
2.
Letter from R. Salvatori, Westinghouse to D. Knuth, AEC, e
i Entitled: Clad Creep Model, NS-SL-507, dated December 8,1972.
9 3.
Letter from R. Salvatori, Westinghouse, to D. Knuth, AEC, NS-SL-518, dated December 22, 1972, with the following J
attachments:.
A.
Creep Constants B.
Fuel Rod Creep Data C.
Pressure Calculations and Void Volume l..
D.
Fission Gas Release and Data I,-
E.
Pellet Thermal Expansion F.
Gap Conductance
- i G.
UO2 Thermal Conductivity
~
H.
Radial Power Distribution I.
Axial Shape and Power History 4.
Letter from R. Salvatori, Westinghouse, to D. Knuth, AEC, NS-SL-521, dated January 4,1973, with the following attachments:
A.
Creep Constants B.
Fuel Rod Creep Data i
C.
Pressure Calculations and Void Volume j
D.
Fission Gas Release and Data E.
Pellet Thermal Expansion F.
Gap Conductance t
G.
U0g Thermal Conductivity H.
Radial Power Distribution 33-I l
--v~--
a
k.
I.
Axial Shape and Power History J.
Fuel Swelling K.
Effect of Clad Ovality on Pellet Fuel Swelling l
L.
Effect of Design Power History-n.
M.
Fuel Rod Plenum Gas Pressure t
N.
Helium Solubility During Radiation l
t t
O.
Release of Absorbed Gases P.
Core Coolant and Rod Surface Temperature g
e 5.
Letter from R. Salvatori, Westir.ghouse, to D. Knuth, AEC, NS-SL-524, dated January 4,1973, with the following attachments:
1.
Effect of Fuel Densification on Design Models 2.
Effect of Fuel Density Yariations p
3.
Supplemental Information - Development of Fq for Point Beach, Unit 2.
4.
Supplemental Information - LOCA Analysis for Point Beach, Unit 2.
5.
Power Spike Probability 6.
Letter from R. Salvatori, Westinghouse, to D. Knuth, AEC, i
NS-SL-527, dated January 2,1973 (non-proprietary version of NS-SL-518, NS-SL-521, and NS-SL-524).
l l
7.
Letter from R. Salvatori, Westinghouse, to D. Knuth, AEC, l
NS-SL-543, dated January 12, 1973, with attachments described in Table 2.
8.
Letter from R. Salvatori, Westinghouse, to D. Knuth, AEC, NS-SL-544, dated January 12, 1973 (non-preprietary version t
of NS-SL-543).
l 9.
J. M. Hellman, Ed., " Fuel Densification Experimental Results and
.Hodel for Reactor Application," Westinghouse Electric Corporation Report WCAP-8218-P-A, March 1975 (proprietary) and WCAP-8219-A, March 1975 (non-proprietary).
9 l l
l
i..
(
10.
J. V. Miller, Ed., " Improved Analytical Models Us'ed in Westing-house Fuel Rod Design Computations," Westinghouse Electric Corporation WCAP-8720, October 1976 (proprietcry) and WCAP-8785, October,1976 (non-Proprietary).
i
- 11. Letter from C. Eiche1dinger, Westinghouse, to J. F. Stolz, NRC, NS-CE-1441, dated May 19, 1977.
- 12. Letter from C. Eicheldinger, Westinghouse, to J. F. Stolz, NRC,
.i NS-CE-1619, dated December 2, 1977.
~j
- 13. Letter from C. Eicheldinger, Westinghouse, to J. F. Stolz, NRC, NS-TMA-1895, dated August 14, 1978.
9 i
- 14. Letter from J. F. Stolz, NRC, to C. Eicheldinger, Westinghouse, i
dated May 2, 1977.
- 15. Letter from J. F. Stolz, NRC, to C. Eicheldinger, Westinghouse, dated November 4,1977.
- 16. Letter from J. F. Stolz, NRC, to C. Eicheldinger, Westinghouse, I
dated January 6,1978.
17.
J. M. Gandhi and S. C. Saxena, " Correlated Thermal Conductivity Data of Rare Gases and Their Binary Mixtures at Ordinary Pressures,"
Journal of Chemical and Engineering Data, Vol.13, No. 3 (July 1965).
18.
W. G. Kannuluik and E. H. Carman, "The Thermal Conductivity of Rare I
Gases," Proceedinns Physics Society London, Yol. 65, No. 701 (September 1952).
19.
N. V. Tsederberg, Themal Conductivity of Gases and Licuids, M.I.T.
Press, Cambridge, MA (1965).
- i ll 20.
L. S. Zaytseva, Zhurnal Tekhnicheskoi Fiziki, Vol. 29, No. 4, 497 (1952)
'i
- 21. Second Annual American-Geman Exchange Program on Fuel Performance Modeling, Idaho Falls, Idaho, May 16-20, 1977.
.2 2.
E. H. Kennard, Kinetic Theery of Gases, McGraw-Hill, New York, N.Y.,
(1938).
23.
N. K. Zimina, " Combined Allowance for the Effect of Pressure and Accomodation Coefficient on the Thermal Conductivity of Light Gases at Elevated Pressures," Heat Transfer-Soviet Research, Vol. 6, No. 6 (November-December,1974).
24.
J. M. Broughton and P.E. MacDonald, " Gap Heat Transfer (GAPHTR),"
MATPRO - VERSION 09, Appendix C, Idaho National Engineering Laboratory Report, TREE-NUREG-1005, December 1976.
i !
7- - -~
.t **.
t 25.
.I. Cohen, B. Lutsman, and J. O. Eichenberg', " Measurement of Themal Conductivity of Metal-Clad Uranium Oxide Rods During l
Irradiation," Westinghouse Electric Corporation Report WAPD-228, August 1960.
26.
M. G. Balfour, J. A. Christensen, and H. M. Ferrari, "In-Pile Measurement of UO2 Themal Conductivi7,y," Westinghouse Electric i
Corporation Report WCAP-2923, 1966.
- 27. i.. N. Duncan, " Rabbit Capsule Irradiation of UO2 CVTR Project,"
Carolinas-Virginias Nuclear Power Assoc., Inc. Report CVNA-142, j
June 1962.
28.
D. D. Lanning and C. R. Hann, " Review of Methods Applicable to the Calculation of Gap Conductance in Zircaloy-Clad UO2 Fuel Rods,"
Battelle Pacific Northwest Laboratories Report BNWL-1894, April 1975.
29.
C. E. Beyer, C. R. Hann, D. D. Lanning, F. E. Panisko and L. J.
Parchen, "GAPCON-THERMAL-2: A Computer Pro the Thermal Behavior of an Oxide Fuel Rod," gram for Calculating i
Battelle Pacific 3
Northwest Laboratories Report BNWL-1898, November 1975.
30.
F. M. Bordelon et al., "LOCTA-IV Program: Loss-of-Coolant Transient Analysis," Westinghouse Electric Corporation Report WCAP-8301, June 1974.
31.
F. M. Bordelon et al., SATAN VI Program: Comprehensive Space-Time Dependent Analysis of Loss-of-Coolant," Westinghouse Electric Corporation Report WCAP-8302, June 1974.
32.
R. A. Dean. "Thennal Contact Conductance Between UO2 and Zircaloy-2,"
Carolinas-Virginia Nuclear Power Assoc. Report CYNA-127, May 1972.
33.
A. M. Ross and R. L. Stoute, " Heat Transfer Coefficient Between UO2 and Zircaloy-2," Atomic Energy of Canada Limited Report AECL-1552, June 1962.
34.
J. Weisman, P. E. MacDonald, A. Miller and H. M. Ferrari, " Fission Gas Release from U02 Fuel Rods with Time Varying Power Histories,"
Trans. Am. Nuclear Soc., Vol. 12, 1969.
35.
E. Roberts, M. G. Bal four, G. W. Hopkins, W. R. Smalley and E. Doguidt " Fuel Modeling and Performance of High-Burnup Fuel Rods," ANS Topical Meeting on Water Reactor Fuel Performance, May 9-11, 1977, St. Charles, Ill.
- 36.. American Nuclear Society " Status Report: ANS-5.4 Fuel Plenum Gas Activity (N218) (Fission Product Release from UO2 Fuel)," ANS-5.4 l
Working Group, April 1977.
37.
R. O. Meyer, C. E. Beyer, and J. C. Voglewede, " Fission Gas Release from Fuel at High Burnup," USNRC Report NUREG-0418, March 1978.
1 i
---r- - -
--.n, mr
+
t.. ' *
(.
I
~
38.
C. E. Beyer and C. R. Hann, " Predictions of Fission Gas Release from UO2." Battelle Pacific Northwest Laboratories Report BNWL-1875, November 1974.
39.
D. S. Dutt and R. 8. Baker, "SIEX: A Correlation.for the Pre-diction of Liquid Metal Fast Breeder Reactor (LMFBR) Fuel Thermal Performance," Westinghouse'Hanford Report HEDL-THE 74-55, June 1975.
~
40.
D. W. Brite et. al., "EEI/EPRI Fuel Densification Project,"
Electric Power Research Report EPRI-131, March 1975.
,1 41.
P. J. Kuchirka, " Properties of Fuel and Core Component Materials,"
Westinghouse Report WCAP-9179 (transmitted by NS-CE-1584, dated October 25, 1977).
42.
G. Kjaerheim and E. Rolstad, "In-Pile Determination of UOp~ Themal Conductivity, Oensity Effects and Gap Conductance," Halden Research Project Report HRP-80, December 1967.
I 43.
I. Devold, "A Study of the Temperature Distribution in UO 2 Reactor Fuel Elements," Aktiebolaget Atomenergie AE-318, 1968.
l 44.
E. T. Laats and P. E. MacDonald, "NRC-0 ECD Halden Project Fuel Behavior Test Program-Experimental Data Report for Test Assemblies,"
IFA-226 and IFA-239," March 1975.
4 5.-
R. E. Williford and C. R. Hann, " Effects of Fill Gas Composition and Pellet Eccentricity,"> Battelle Pacific Northwest Laboratories Report BNWL-2285, July 1977.
l t
46.
U.S. Nuclear Regulatory Commission, " Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants - LWR Edition," USNRC Report HUREG-75/087, Section 4.2, " Fuel System Design," Revision 1.
47.
D. H. Risher, Ed., " Safety Analysis for the Revised Fuel Rod Internal Pressure Design Basis," Westinghouse Electric Corporation Report WCAP-8963, November 1976.
48.
U.S. Nuclear Regulatory Commission, " Safety Evaluation of Westing-house Topical Report WCAP-8963," Core Perfomance Branch, May 1978.
1 49.
F. A. Nichols, " Estimated Creep Properties of Zircaloy During Neutron Irradiation," Westinghouse Electric Corporation Report WAPD-TM-756, May 1968.
50.
D. O. Hobson, "Zircaloy Fuel Cladding. Collapse Test Plan," Oak Ridge National Laboratory Report ORNL/NUREG/TM-45, August 1976.
1...
c 51.
D. O. Hobson and C. V. Dodd, " Interim Report on the Creepdown of Zircaloy Fuel Cladding," Oak Ridge National Laboratory Report ORNL/NUREG/TM-103, May 1977.
52.
G. M. Bain and D. L. Baty, " Creep Collapse Testing of Zircaloy-4 Tubes," Babcock & Wilcox Lynchburg Research Center Topical Report, LRC-9065, August 1977.
'T-
- 53. " Hot Cell Examination of Creep Collapse and irradiation Growth Specimens - End of Cycle 1," EPRI/B&W Cooperative Program on PWR Fuel Rod Performance (RP-711-1), Key Phase Report NunDer 2, Babcock & Wilcox Lynchburg Research Center Topical Report LRC 4733-4, July 1977.
[
- 54. " Cladding Creepdown and Ovalization Measurements on Pressurized j
and Non-Pressurized Segments of PWR Sized Fuel Rods," OECD Halden Reactor Project Quarterly Progress Report, October to Decemoer j
i 1976, HPR-200.
- 55. Letter from C. Eicheldinger, Westinghouse, to D. Vassallo, NRC, NS-RS-133, dated February 5,1974.
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[
TABLE II ATTACHMENTS TO WESTINGHOUSE LETTER NS-SL-543' Attachment Subject
~
1 Supplemental information on the sensitivity of gap conductance to fission product release.
t-2 Supplemental infomation on fission gas release for Westinghouse i
design case.
3 Supplemental information on fuel design model predictions.
4 Supplemental information on clad film thickness.
y 5
Supplemental information on gap dimensions for fuel rods used to develop creep model.
~
6 Supplemental information on predicted versus observed clad flattening.
7 Supplemental information on two dimensional fuel pellet temperature calculation.
8 Supplemental information on clad temperature adjacent to a fuel gap..
9 Supplemental information on the methods used in the LOCA analysis.
f 10 Supplemental information on the treatment of pellet density variations in accident evaluations.
11 Supplemental information on the value of uncertainty applied to fuel rod average temperature for use in accident evaluations.
12 Supplemental information on the expected reactivity consequences of a rod-ejection accident as compared to cases analyzed in the Point Beach 2 submittal of a rod-ejection.
13 Supplemental information on fuel rod integrity following the control rod ejection and locked rotor accidents.
I 14 Supplemental information on collapse time reductions.
15 Supplemental information on fission gas release.
- j.