ML20011F498
| ML20011F498 | |
| Person / Time | |
|---|---|
| Site: | Saint Lucie  |
| Issue date: | 11/30/1989 |
| From: | ABB COMBUSTION ENGINEERING NUCLEAR FUEL (FORMERLY |
| To: | |
| Shared Package | |
| ML17223A533 | List: |
| References | |
| CEN-396-(L)-NP, NUDOCS 9003060141 | |
| Download: ML20011F498 (90) | |
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- 0 VERIFICATION OF THE ACCEPTABILITY OF ,
%a A 1-PIN BURNUP LIMIT OF'60 MWD /KG .# .g,, 1 o -- ws s 4: 4 FOR ST. LUCIE UNIT 2
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, ~ ;t A; A;M.] Combustion Engineering, Inc. l Nuclear Power Businesses fi.M % 2d _ 1000 Prospect Hill Road g Windsor, Connecticut 06095 f:1;;y
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LEGAL NOTICE THIS REPORT WAS PREPARED AS AN ACCOUNT OF WORK SPONSORED - BY CotWUSTION ENGINEERING, INC. NEITHER COMBUSTION ENGINEERING l i NOR ANY PERSON ACTING ON ITS 88HALP: )
- A. MAKES ANY WARRANTY OR REPRESENTATION, EXPRESS OR llWLIED INCLUDING THE WARRANTIES OP PITNESS POR A PARTICULAR PURPOSE OR MERCHANTABluTY, WITH RESPECT TO - THE ACCURACY, ;
CotrLETENESS, OR USEPULNESS OF THE INPORMATION CONTAINED IN THIS i REPORT, OR THAT THE 128 0F ANY INPORMATION, APPARATUS, METH00, i
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OR PROCESS Die nesn IN THIS REPORT MAY NOT INFRINGE PRIVATELY OWNED RIGHT8;0R
- 5. ASSUMES ANY LIABILITIES WITH RESPECT TO THE USE OP.OR FOR DAMAGES RESULTING PROM THE USE OF, ANY INFORMATION, APPARATUS, METH00 OR PROCESS DISCLOSED IN TH88 REPORT.
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t I i ABSTRACT Several Utilities using Combustion Engineering 16 S fuel assembly designs have implemented programs to extend their fuel cycle L 7ths from 12 to 18 ; months'and beyond, ine maximum 1-pin burnup predicted for these extended burnup cycles exceeds the 52 MWD /kg limit presented in the existing C-E . Extended Burnup Operation topical report. This report verifies the adequate modelling of the St. Lucie 2 reload fuel design pins to 60 MWD /kg (the new limit required by the implementation of longer fuel cycles) by supplementing the existing topical report with additional data and discussions. The conclusions of this report regarding fuel assembly length change and shoulder gap change are applicable to the St. Lucie 2 reload fuel assembly design employing [ ). ; b I l ii L
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- A85 TRACT. 11 TABLE 0F CONTENTS iii t
p LIST OF TABLES iv- t w L' LIST OF FIGURES y *
- . . INTRODUCTION l' +
L - DISCUSSION. 2 , L' 3.3.6.a . Cl add i ng C o11 ap s e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3
- h 4.1.1.a- Fatigue.............................................. 4 4.1.2.a- Cladding Corrosion................................... 5 .
4.1.3.a Cladding Creep...................................... 11 . 4.1.4.a' Cl add i ng C o11 a p s e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13 1 4.1.5.a1 Ductility of Fuel C1 adding.......................... 16 4.1.6.a Fission Gas Release................................. 28 . 4.1.7.a1 Fuel Thermal Conductivity............................ 35 4.1.8 a Fuel Mel ting Temperature . . . . . . . . . . . . . . . . . . . . . . . . . . . . 36 - 4.l.9.a. ; Fuel Swe111ng....................................... 37 ' 4.1.10.a Fuel Rod Bow........................................ 38
-4.1.11.a Fretting Wear....................................... 39 4.1.12.a Pellet / Cladding Interaction.................. ... .. 40,
~4.1.13.a' Cladding Deformation and Rupture.................... 42 4.1.14.a Fuel Rod Growth..................................... 43
-4.2.1.a ' Guide ~ Tube Wear................. ...... .........,.. 47
, 4.2.2.a ~ Fuel Assembly Length Change: and Shoulder Gap Change. 48 4.2.3.a Fue l As sembly Ho1 ddown . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 62 4.2.4.a Grid Irrad i ati on - Growth. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 63 - i
~4.2.5.a Spacer Grid Rel axati on. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 64 4.2.6.a Corrosion of the Fuel Assembly Structure............ 65 4.2.7.a Burnable Poison Rod Behavior........................ 66 j CONCLUSION 75 REFERENCES 76 ,
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-f TABLES -l M h 1 4.1.6'a-1~
. Fission-Gas Release-Data from Fort Calhoun Fuel Rods.. 32 4.1.6.a-2 Fission-Gas Release Data from Zion 1 Fuel Rods........ 33 -
.4.1.6.a-3 FATES 3B Predictions of Gas Release from High Burnup,.. 34 i Low Power Test Rods j 4- 4.2.2.a-1 Analytical Models for [ )...............53' 4.2.7.a-1 Burnable Poison Rod Detail s. . . . . . . . . . . . . . . . . . . . . . . . . . . 73 b
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7 i r FIGURES Figure ELqt 4.1.2'.a 1 Oxide vs. Burnup ...................................... 8 4.1.2.a-2 Cladding Peak Oxide. Thickness as a Function of ........ 9 3 Average Burnup 4.. l . 2. a-3 Hydrogen Uptake as a Function of 0xide Thickness ..... 10 for Zircaloy 4 Cladding in PWRs
... (
4.1.3.a-1 Diameteral Strain of High Burnup Rods Irradiated ..... 12 in' Fort Calhoun and Calvert Cliffs-1 . 4.1.5.a-1 Yield Strength as a Function of Fluence for . .. .. . . . . . 22 'P [ .) Irradiation Temperature 500 to 650*F, Elevated Temperature Test 4.1.5.a-2 Ultimate Tensile Strength of Shog-Trangverse ........ 23 Specimens Irradiated to 4.3 x 10 n/cm (E>1Mev) 4.1.5.a-3 Uniform Elongation as a function of Fluence for ...... 24 [ . ] Zircaloy, Irradiation temps. 560 - 610*F 4.1.5.a-4 Percent Reduction of Area for Shgt-Trayerse ......... 25 Specimens Irradiated .to 4.3 x 10 n/cm (E>lMev)
'4.1.5.a-5 Effect of Hydrogen Concentration on the 8 eduction .... 26-of Area for Zircaloy-2 Irradiated to 10 2 n/cm2 4.1.5.a-6. Fluence Dependence of Strain for Irradiated .......... 27 Zircaloy-4 4.1.12.a-1 PCI Test Results on Standard C-E and KWU Rodlets ..... 41 4.1.14.a-1 Fuel Rod Growth Measurements Compared to C-E . . . . . . . . . 46 Zircaloy Fuel Rod Growth Model 4.2.2.a-1 Typical Probability Histogram for Fuel Assembly . . . . . . 54 -
Length Change 4.2.2.a-2 Comparison of 16 x 16 ( ) Guide Tube Length ........ 55 Change to SIGREEP Predictions b- 4.2.2.a-3 Comparison of ANO-2 Batch D Shoulder Gap Changes ..... 56
- to SIGREEP Predictions v
ir FIGURES Fioure ELqg 4.2.2.a 4 ~ Comparison of Shoulder Gap Change to SIGREEP ........., 57 Predictions for 16x16 fuel Assemblies with [ ] Guide Tubes 4.2.2.a-5 < Comparison of Shoulder Gap Change to SIGREEP . . . . . . . . . 58 Predictions for St.-Lucie 2 Fuel Assemblies with ( ) Guide Tubes 4.2.7.a Swel l i ng o f. Al 0 ' 0 C . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7 4 - 23 4 ,
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INTRODUCTION Several Utilities using Combustion Engineering 16x16 fuel assembly designs have implemented programs-to extend their fuel cycle lengths from 12 to 18 months and beyond. The predicted maximum 1-pin burnup for these extended burnup cycles exceeds the current C-E l-pin burnup limit presented in , Reference 1, 52 MWD /kg. This report justifies a 1-pin burnup limit of 60 MWD /kg for the St. Lucie 2 reload fuel assembly design by supplementing Reference I with data and discussions covering the additional burnup range c required by the implementation of longer cycles, 52 MWD /kg'to 60 MWD /kg. Reference 1 also specified a limit on batch average discharge burnup. However, a review of the various burnup dependent fuel performance topics discussed in Reference 1 indicated no explicit dependence on batch average-burnup. Therefore, the-C-E batch average discharge burnup limit of Reference , I has been deleted. Reference 1 presented data and/or discussions on 21 fuel performance topics that.were judged to be burnup dependent and/or important in determining the behavior of fuel at extended burnup. The existing data and discussions presented in Reference 1 support the acceptability of a 1-pin burnup limit of 60 MWD /kg for the following 8 fuel performance topics: fatigue of the fuel rod, fuel rod bowing, fuel rod fretting wear, cladding deformation and rupture, guide tube wear, fuel assembly holddown, grid irradiation growth and spacer grid relaxation. Consequently, only a short discussion is provided for
-each of these topics.
The remaining 13 fuel performance topics are discussed within update sections that present the additional data and/or discussions needed to support the - acceptability of a 1-pin burnup limit of 60 MWD /kg. h The conclusions of this report regarding fuel assembly length change and shoulder gap change are applicable to Combustion Engineering fuel assembly design for St. Lucie 2 employing [ 1 l 1 i- 1 l 1 Discussion The contents of the following update sections generally follow the format of their respective section in Chapter 3 or 4 of Reference 1. Each (sub)section l 1s numbered identically to its respective (sub)section-in Reference I with the
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addition of ".a". Each section has an introduction which specifies how the j succeeding subsections should be treated, i.e., whether they append or replace 1 their respective. subsection. The figures, tables and references of each section- are numbered sequentially in the following form, "section #" " sequence #", e.g., 4.1.3.a-1, with the exception of Reference I which'is a general reference that applies to all sections of this report. 4 e I W 2-
1 i 3.3.6.a Claddino co11anse !
)
.This section replaces Section 3.3.6 of Reference 1.
Collapse is the term applied to a condition where a slightly oval cladding ; tube will " flatten" into a significant axial gap in its fuel or poison pellet ! column. The conditions leading to collapse are long term phenomena since collapse occurs only after the cladding has crept into an oval shape from its nearly circular shape at beginning of life. The driving force for this creep is ' supplied by the differential pressure across the fuel or poison rod , cladding. C-E design characteristics which mitigate cladding collapse are: o Fuel and poison rods are prepressurized with helium which offsets the effects of external pressure to the extent that cladding long term creep and cladding ovalization kre reduced. ' o "Non-densifying" or stable fuel pellets are used to prevent the formation I-of significant axial gaps within the fuel column. This allows the fm.1 pellets to support the cladding later in life when the fuel-cladding ,,4p closes. o Poison rods behave in a similar fashion to fuel rods except the pellets are not subject to densification. The cladding collapse model is discussed in Section 4.1.4.a. l-b 1
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E' 1-iE 4.1.1.4 .Faticue The discussion provided in Section 4.1.1 of Reference I applies to the
- proposed increase in the 1-pin burnup limit to 60 MWD /kg. The metho'd used to calculate fatigue damage is applicable to extended burnup operation since the other sections of this report show that the individual components of the method (e.g.. cladding creep and fuel swelling) are adequately modeled and the cladding has adequate ductility. .
(I. 5 9 e W T-7 , i l 4.1.2.a. C1addina Corrosion . The following- subsections append the corresponding subsections of ' Reference'l. 4.1.2.1.a corrosion Behavior i
,; 0xide thickness data from- three C-E PWRs, Calvert Cliffs-1, Ft. Calhoun, .
and ANO 2, for rod average burnups of up to [ ] MWD /kg have recently - become 'available (References 4.1.2.a-1, 4.1.2.a-2, 4.1.2.a-3, 4.1.2.a-4).
- The maximam burnup rod for which oxide' thickness data are available for the -
14x14 design is approximately [ ] MWD /kg (Reference 4.1.2.a-4) and for the igx 16 design (Reference 4.1.2.a-3) is approximately 58 MWD /kg. The recent high-burnup oxide thickness data are presented together with the data of Figure 4-3 of Reference 1 in Figure 4.1.2.a-1. -The U 235 enrichment level for these high-burnup rods was between 3.03 and 4.00%. The U enrichment 235 for future fuel batches is expected to increase,. but the burnups are not expected to exceed 60 MWD /kg. The available oxide thickness data on i irradiated fuel cladding approximately covers the maximum burnup level of future high burnup rods. As.a first approximation, the oxide thickness at a burnup of 60 MWD /kg was estimated from a regression fit to the 16x16 l (ANO-2) oxide thickness data. Regression analysis of the 16x16 (ANO-2) oxide data resulted in a best estimate oxide thickness of [ ] microns at-60 MWD /kg and an upper bound (x.+ 3a) oxide thickness of [ ] microns at 60 MWD /kg. A similar fit to the Calvert Cliffs-1 14x14 data yields a l best-estimate thickness of [ ] microns and an upper bound of [ ] microns. Recently published high-burnup corrosion data from other PWRs (References
- l. -4.1.2.a-7 to 4.1.2.a-10) are presented together with the data from Figure 4 4 of Reference 1 in Figure 4.1.2.a-2. It is worthwhile to note that the corrosion data presented in Figure 4.1.2.a-2 refers to fuel rods with fuel l enrichments lower than 4% U 235 and several irradiation cycles of the order i of 12 months duration. The heat rates of these fuel rods are generally L lower than those expected for future high-burnup fuol rods.
1 1 I
-Nevertheless, for a rod average burnup of - 60 MWD /kg,- the upper limit of expected oxide thickness from Figure 4.1.2.a-2 is about 100 microns. This is in reasonable agreement with the upper bound estimate presented above for the ANO-2 data [ -] and the Calvert Cliffs-1 data' [
]
Another important aspect of cladding corrosion is the extent of hydrogen
+ uptake by the cladding. A fraction of the amount-of hydrogen liberated by the Zircaloy corrosion reaction is absorbed by the cladding. As discussed . in Section 4.1.5.a. the absorbed hydrogen may reduce the ductility of the cladding. Hydrogen concentrations measured on cladding specimens from several' PWR fuel rods are presented in Figure 4.1.2.a-3. A detailed-analysis of the data (Reference 4.1.2.a-11) shows that a pickup fraction of 18% represents a reasonable upper limit on hydrogen absorption by cladding at high burnups. This pickup fraction translates to a cladding hydrogen level of about [ ). The relationship between hydrogen level and cladding ductility is further d sq issed in Section 4.1.5.2.a.
4.1.2.2.a Evaluation of Claddino Corrosion at Extended Burnuo Based on the limitations discussed in Section 4.1.2.1.a. the 3a upper bound' - oxide thickness is estimated to be about [ ] microns for fuel cladding at a rod average burnup of 60 MWD /kg. The cladding wastage due to this level of oxide layer thickness is insignificant with regard to cladding stresses. Although some tendency towards (oxide spalling] was l i reported at a level of oxide layer thickness of about [ ] microns j- (References 4.1.2.a-4 and 4.1.2.a-7), fuel rod integrity was not impaired. It is, therefore, concluded that cladding corrosion is not likely to impair the integrity of fuel rods irradiated to rod average burnups of 60 MWD /kg. An oxide -layer will, of course, increase the surface temperature of the
. cladding. For example, the maximum local temperature increase at the metal-oxide interface due to a 3a upper bound oxide layer (on the 16x16 or j l_ 14x14 design fuel rods), assuming a local fuel rod linear heat rate of
( ) kw/ft, is calculated to be about [ j. On a rod-average 1
]
1 l
1
, +1
-1 basis, the_ temperature increase at the metal-oxide interface will be considerably _less. The it.rgest impact of the insulating' oxide layer occurs J
, at end-of-life when the linear heat rate of the fuel rod is significantly -
lower than the linear heat _ rate..of .the- peak- rod in the, core.. Thus, it is i concluded that-the effect of oxide build-up on fuel temperature and stored ) energy is essentially - counteracted by the lower linear heat rates that l occur towards end-of-l i fe. Thus, corrosion, . based on observed l
- j. . oxide-thicknesses at 60 MWD /kg in operating reactors, will not be limiting.
l However, additional factors in the future must be considered. Specifically, the factors that need to be considered are EFPD l (corresponding average-linear heat rate) to achieve 60 MWD /kg and the , reactor coolant conditions (temperature and chemistry). If these -factors differ substantially from the data base, additional corrosion evaluations would be warranted. These factors will be monitored and corrosion evaluations performed as necessary, particularly if 1) the EFPD to achieve maximum burnup are considerably shorter, 2) the reactor inlet temperatures are considerably higher, or 3) the coolant lithium level is significantly higher than the ranges covered in the current data base from the operating reactors. In addition, the ic+ect of cladding changes that optimize composition and processing history to improve the in-reactor corrosion . resistance compared to that used in the current data base will also be included in the above evaluations.
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= Cladding. Peak Oxide Thickness- as a Function of Average.Burnup :
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Figure 4.1.2.a-3 Hydregen Uptake as a Function of 0xide Thickness for Zircaloy-4 Cladding in PWRs I s -j l l 4.1.3.a Claddina Creen I Append Subsection 4.1.3.3 of Reference 1 with the following material: l
- 4.1.3.3.a Evaluat-ion of Creen 9 .
Diametral creep measurements are now available for several high burnup fuel rods irradiated in Calvert Cliffs-1 and Fort Calhoun (References 4.1.3.a-1_ and 4.1.3.a-2) . These data, corrected for the presence of oxide and converted to resulting diametral strain, are presented in~ Figure 4.1.3.a 1. The rod average burnups of these rods ~ are [
] Due to the contact between the fuel pellet and cladding at these high burnup levels, the fuel rod diametral strain is strongly influenced by the _ fuel pellet's swelling behavior. The data presented in Figure 4.1.3.a-1 show that the diametral behavior of the fuel rod is.a continuous _
function to rod average burnups of 60 MWD /kg and that the model discussed in Reference 1 is adequate for 1-pi.n burnups of up to 60 MWD /kg. , The diametral strain data presented in Figure 4.1.3.a-1 show that the fuel , L 1 rod diameter does not change significantly during extended burnup
- operation.- Early-in-life, prior to the establishment of fuel-cladding l contact, cladding creepdown occurred due to. coolant pressure. [
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. . Figure 4.I.3.a-1 Diameteral Strain of High Burnep Rods -Irradiated '
in Fort Calhoun and Calvert Cllffs-l i i 2G E9 6- x
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h 4.1.4.a Claddina Co11anse
'I This section replaces Section 4.1.4 of Reference 1. ';
Cladding tubes generally have a minor degree of variation from a perfectly. ; circular cross section with uniform wall thickness. When subjected to a net external pressure in the reactor, bending stresses are produced as a [
- result of the slightly imperfect geometry. Under the high temperature and neutron flux conditions in the reactor, the Zircaloy cladding creeps in L._ response to the bending stresses. The resulting creep strain increases the deviation from the circular shape,-thereby increasing the bending stresses.
This process continues at an increasing rate until contact is made with the pellets, or if a significant axial gap exists in the pellet column, until L an unstable condition is reached and the cladding " collapses" into a distorted shape. Observations indicate that no significant axial gaps form in the fuel L pellet column during the operation of Combustion Engineering's modern l design fuel, which has prepressurized fuel rods and stable "nondensifying" fuel- pellets. Such gaps would be evidenced by unusual local evalities of
' the fuel rod cladding, a distinct region of atypical crud deposition around -
the cladding circumference, or atypical signals during gamma scanning. None- of these indications- have been observed during the extensive post-irradiation examination programs conducted on both the 14x14 and 16x16 fuel designs. It can be inferred from these post irradiation examinations of modern design C-E fuel that during-hot full power cperation the axial gaps in a fuel column are usually only a fraction of the length of a pellet. The {' T gaps are measured in the cold condition. The largest cold gap measured in modern C-E fuel was 0.9 inches. It was calculated that thermal expansion of the fuel column during reactor startup reduces this cold gap to $ 0.3 l inches. Thus, the largest hot gap inferred from all post irradiation examinations of modern C-E fuel was 0.3 inches. This conclusion is supported by the corrosion patterns observed during visual examinations.
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x L - e . a 4.1.4.1.a Modelina of Claddina Collaosu The current methods of evaluating resistance to cladding collapse are described in Reference ~4-17 of Reference 1, and Reference .~ 4.1.4.a-1. Reference 4-17 of Reference 1 describes a method which utilizes the CEPAN computer code to predict creep deformation and collapse time of Zircaloy , cladding containing an initial ovality. Although large hot gaps have not
' been inferred for modern design C-E fuel, this method assumes that a gap in a the pellet column exists at the most unfavorable elevation in the fuel rod.
- No credit.is taken for the support offered by the pellets at the-edges of 4 the gap. This original method of selecting input to CEPAN resulted-in a deterministic combination of the worst case cladding as-built dimensions and worst case operating conditions during the fuel lifetime. The NRC concluded that CEPAN provides an acetptable analytical procedure for determining the minimum time to collapse for C-E Iircaloy clad- fuel. If this . minimum collapse time exceeds the fuel lifetime, then collapse.
resistance has been demonstrated. [ A modification of the above method is described in Ref. 4.1.4.a-1. This modification is applied to the normal CEPAN results to account for the ' support provided to the cladding by1the pellets at the edges of the gap. The adjustment varies' as a functicn of the length of the gap or unsupported cladding. Ac the-gap considered becomes longer, the results approach the normal CEPAN results. L !' 4.1.4.2.a Effect of Extended Burnuo f Since cladding collapse is a creep-related phenomenon, the longer residence times associated with extended burnup fuel will increr.se the amount of creep of unsupported cladding. The increased creep strain will be accounted for in the analysis of the ability of the fuel red to resist cladding collapse.
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l 4.1.4.3.a Evaluation of Claddino Co11anse Although early experience with densifying 00 2 fuel pellets indicated that cladding' collapse could result in fuel failure, improvements in fuel design, notably the development- of stable fuel pellet types and rod pressurization, have essentially eliminated this ' concern. Current commercial fuel pellets have shown through operating performance that ) significant- axial gaps do not form in the fuel pellet column . during j operation. Without- the occurrence of gaps of sufficient length, cladding t collapse cannot occur and, as . a consequence, the cladding will remain j stable and will not be subject to high local strains from this effect. 1 Furthermore, there is no evidence to indicate that continued operation of - fuel rods having cladding in oval contact with the fuel pellet column is detrimental. C-E has performed cladding collapse calculations with the modified method 3 described in Section 4.1.4.1.a using very conservative input assumptions. The assumed length of the axial gap in the fuel column-bounded the largest hot axial gap in modern C-E fuel (See Section 4.1.4.a). These calculations i
.have shown that the predicted collapse times far exceed the longest ~
residence time ever expected for C-E fuel that is operated to a maximum 1-pin burnup of 60 MWD /kg. It has therefore been concluded that unless significant changes in design or- manufacturing methods are introduced, modern C-E fuel and poison rods for both 16x16 and 14x14 designs are not susceptible to cladding collapse. On this basis, C-E will no longer specifically address cladding collapse for new cores or reload batches unless design or manufacturing changes are introduced which would
. significantly reduce predicted collapse time results. In the event such l, changes do occur, the modified method described in Section 4.1.4.1.a will be used to confirm that cladding collapse will not occur during the design lifetime of the fuel.
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I 4.1.5.a Duetility of Fuel Claddino j 1 l This section replaces Section 4.1.5 of Reference 1. Exposure of the fuel rod Zircaloy cladding to fast neutron irradiation' ) causes the cladding ' material to strengthen and lose some of its ductility. In addition, the fuel rod Zircaloy cladding reacts with water during reactor operation to form a zirconium dioxide (Zr02 ) layer on the outer
- surface of . the fuel rod. Hydrogen is produced by this reaction and a .- fraction.of the liberated hydrogen (approximately 0.18) is absorbed by the l cladding. This hydrogen uptake may also reduce- the ductility of the ;
cladding. The fuel rod design criteria related to strength and ducti,lity , were discussed in Se:tions 3.3.2 and 3.3.3 of Reference 1, respectively. Since the fuel rod _ design calculatior.s are based on the yield strength of unirradiated cladding, the increase in the yield strength of cladding due to neutron irradiation does not pose a strength limitation on the cladding's performance. The loss of ductility due to the neutron irradiation and hydrogen uptake, however, needs to be' evaluated to assure that adequate cladding ductility exists,at extended burnup levels to ensure , that the design strain limits remain valid. The effect of extended burnup operation on the cladding ductility is evaluated in this section.
- The elevated temperature cladding strain design limit used in the C-E FSARs is 1%. A review of the mechanical property data of high fluence cladding (from fuel rods with rod average burnups up to 60 mwd /kg) (
). Since the deformation capability of irradiated cladding during the normal reactor operation and anticipated transients is important, the mechanical properties of irradiated Zircaloy-4 at the defonnation temperatures of about 600*F were considered in the analysis of the extended burnup data.
I The combined effect of the neutron fluence and hydrogen uptake on the mechanical properties of Zircaloy-4 is evaluated below.
L h 4.1.5.1.a 14echanical Procerties of Irradiated Zirealov at Extended Burnues C E has obtained data on the mechanical properties of Zircaloy-4 cladding irradiated in the Fort Calhoun reactor to local burr.ups of up to 62 mwd /kg (Reference 4.1.5.a 1). In addition, mechanical property data have also become available for fuel cladding irradiated in Oconee-1 (Reference 4.1.5.a 2) and Zion (References 4.1.5.a-3, 4.1.5.a 4) to extended burnups. These data were recently analyzed to evaluate the effects of irradiation and hydriding on the mechanical properties of Zircaloy-4 at high fluences (Reference 4.1.5.a 5). These data are described below together with the low burnup data presented in Section 4.1.5 of Reference 1. CE uses ( ) fuel rod cladding (Reference 4.1.5.a6). The increase in elevated-temperature yield strength due to irradiation is illustrated in Figure a.1.5.a-1,(References 4.1.5.a 7 through 4.1.5.a-10). An increase in yield strength has also been observed by CE at extended burnup (Reference 4.1.5.a 5). The increase in the ultimate tensile strength of irradiated Zircaloy due to higher hydrogen levels, on the other hand, does not appear to be significant (see Figure 4.1.5.a2). The data of Evans and Parry (Reference 4.1.5.a 11) shown in , this figure indicate that there is no change in the ultimate strength of irradiated Zircaloy 2 at temperatures above 100'C (210'F) when the hydrogen level is increased from 0 to 200 ppm. The yield strength behaves in a similar manner. [ b 3 The fluence dependence of the [
] is illustrated in
i Fipere 4.1.5.a-3. The data (Reference 4.1.5.a-12) suggest that for ! ( . I
] These tests were conducted at high strain rates.
It has been theoretically predicted by Nichols (Reference 4.1.5.a 13) and Ibrahim and Coleman (Reference 4.1.5.a 14) and experimentally verified by j Ibrahim (Reference 4.1.5.a 15) and Wood (Reference 4.1.5.a 16) that at the ; lower strain rates more appropriate to the creep deformation rates of the fuel cladding, the uniform elongation is greater than estimated from the short term, high strain rate mechanical tests. Irradiation data for a low ; 21 2 fluence (1.8 x 10 n/cm ) nickel free Zircaloy 2 (Reference 4.1.5.a-15) indicate that at a stress of 332 MPa, the creep rupture strain is greater than 5.1%. Two factors need to be considered for in reactor creep of ; cladding with higher fluence. Firstly, at the lower stresses appropriate , to in reactor cladding creep, the creep strains at rupture are expected to l be higher (Reference 4.1.5.a-13). Secondly, with an increasing level of fluence, the creep strain will decrease. Based on the available ductility ' data on high fluence cladding irradiated in power reactors, it is concluded , that the cladding ductility at high burnups will be significantly greater than 1% as a result of the net effect of these two opposing factors. 4.1.5.2.a influence of Hydrocen on Mechanical Proceties ! A fraction of the amount of hydrogen liberated by the Zircaloy corrosion reaction with the primary coolant is absorbed by the cladding. It remains in solution in the Zircaloy until the terminal solid solubility of hydrogen ' is exceeded. At 300*C (572'F), the solubility limit is approximately 100 ppm. Amounts in excess of the solubility limit will precipitate as ,
- b. zirconium hydride platelets.
4
+
- . . . . . . . , - - - . . --.2 .+ ~ . <- ..,-
l l l l It has been established that the _ ductility reduction due to hydrogsn is
~
dependent not only on the quantity of hydrides but also on their l orientation. For example, if the hydrides are precipitated so .that .their. ! major axis is parallel to an applied stress, the reduction in ductility is relatively small. [
'I I
)
1 Evans and Parry (Reference 4.1.5.a.11) determined the temperature above l which the effects of unfavorably oriented hydrides disappear in cold-worked and stress relief annealed Zircaloy 2 cladding irradiated to low fluences. l At temperatures above 200'C (392'F), adversely oriented hydrides up to 200 t ppm did not influence the ductility as measured by the reduction in area j (Figure 4.1.5.a-4). Watkins et al. (Reference 4.1.5.a 17) have conducted j tests on cold worked tubular samples of Zircaloy 2 prehydrided to levels of J l up to 800 ppa which have circumferentially oriented hydrides. Tensile f tests showed that hydrogen concentration had only a minor effect on , ductility at 300'C (572*F) (Figure 4.1.5.a 5). Specimens charged with l hydrogen showed values of the reduction in area at failure in e'xcess of : Thus, it has been concluded that at- elevated temperatures, 60%. circumferential1y oriented hydrides up to 800 ppm do not influence the i ductility of Zircaloy cladding irradiated to fluence levels of 1020 n/cm2 , a i l~ 4.1.5.3.a Combined Effect of Radiation Damace and Hvdridina on the ; Ductility of Claddina at Extended Burnues : 1 The ductility of extended burnup fuel rod cladding with rod average burnups l,. approaching 60 MWD /kg (local burnups to 62.5 mwd /kg corresponding to i 2 cladding fluence levels to 16.2 x 102I, n/cm , E > 0.821 MeV) has been j recently measured by axial and ring tension tests and diametral burst tests (References 4.1.5.a-1 to 4.1.5.a-4). The results of these mechanical tests L. _ __. _ . _ . _ ~ _ . . _ _ _ . . _ _ _ _ _ _ . _ . _ _ . _
i demonstrated the combined effects of neutron damage and hydrogen uptake on ; the mechanical properties of highly irradiated Zircaloy 4. The strain rates resulting from the load application in these tests were also . significantly higher than the fuel rod cladding strain rates expected j during normal steady state operation and also during the anticipated , operational transients of a power reactor. Ring tensile tests at 650*F on ; cladding from 5 cycle rods (rod average burnups 49.5 to 49.9 mwd /kg) j 1rradiated in Oconee-1 (Reference 4.1.5.a 2) show uniform strains in the range of 2 to 3% and total strains in the range of 3.8 to 8.4%. Axial tension tests at 650'F on cladding from the same rods resulted in uniform , strains in the range of 0.93 to 1.43% (average 1.29%) and total strains in the range of 5.68 to 15.31%. Therefore, the Oconee 1 cladding data indicate that at a burnup of about 50 mwd /kg, the cladding can withstand an : additional strain of 1% pricr to plastic instability and about 4% strain prior to failure. Axial tension tests on six cycle Fort Calhoun cladding (local burnups in the range of 57.6 to 63.3 MWO/kg) (Reference 4.1.5.a-1) show that for a deformation temperature range of 392 to 752'F, the uniform strains are in the range of 0.7 to 0.8% and total strains are in the range of 5 to 9%. , Thus, Fort Calhoun cladding tensile data indicate that at an end of life burnup of approximately 60 MWD /kg, the cladding can withstand approximately 1% additional strain prior to the onset of plastic instability and at least . 5% additional strain prior to failure. ' Burst test data on high burnup cladding are available from fuel rods irradiated in Fort Calhoun and Zion. The burst test data on Zion cladding with a rod average burnup of 55.3 MWD /kg show total circumferential strains of 0.79 to 2.69% (Reference 4.1.5.a-3). At lower burnup levels of 38 and 46 MWD /kg, the Zion cladding burst test results show total circumferential I strain values above 3% (Reference 4.1.5.a 4). Burst test data are available on Fort Calhoun cladding with rod average burnups approaching ! 60 MWD /kg (Reference 4.1.5.a-1). At a local burnup of 41.6 MWD /kg, the uniform strain values are 1.12 and 1.21% and total strain values are 6.9 and 5.6%. At a local burnup level of 52.3 53.2 MWD /kg, the uniform strains I are 1.43 to 1.75% and total strains are 4.5 to 4.7%. At a local burnup ! level of 54.7 to 62.5 MWD /kg, the uniform strains are 0.03 to 0.11% and ; total strains are 1.24 to 4.19%. ! The material ductility at 572 to 599'F as a function of fluence is shown in Figure 4.1.5.a 6 (Reference 4.1.5.a 5). For fluence values up to ,
' 21 2
-9x10 n/cm2 (E > 0.821 MeV or 8x1021 n/cm E > 1 MeV) (corresponding to burnups up to -53 mwd /kg), [
l J Moreover, based on a detailed analysis of the f microstructures of the fractured specimens, the fracture mode at burnure greater than 53 mwd /kg was determined to be ductile (Reference 4.1.5.a-5). The observations described above indicate that at a burnup level of 60 mwd /kg, the cladding material has ( ja strain limited cladding failure is not expected at a burnup level of 60 mwd /kg due to an operational transient. Additional confirmation of acceptable cladding performance to rod average burnups up to (
) Acceptable cladding performance to rod average burnups up to approximately 58 MWD /kg was also recently demonstrated for 16x16 fuel assembly designs (ANO 2) (Reference +
!* 4.1.5.a-19). ,
~ . . . . . -. . - . . . . . . . . - . . . . . . . .. .~.. .. _..- _ .. - - _.... . .-,_,......
l I I m b m I <o
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i I i I> ! ; E , I ! 9 i i t i
.. s yggT TEMPERATURE. 'C .
FIGURE 4.1.5.A-2 ULTIMATE TENSILE STRENGTH OF SHORT-TRANSVERSE - SPECIMENS IRRADIATED TO 4.3 X 101I N/CM2 (E > 1 MeV)
. . . _ - . . - . . . . . . . . . . . . - , . . ~ . . - . . ~ . ~ - . . - . . . . . . . . . - - . . . - . . _ . . . . . . ~ - . _ . . _
h l t i W W e . s- - s
- I g
- i
.- a 7d E I
=g l
e , u 15
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- u ,
= =- y ,
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= b L 5
+ l .
Ie I I w .
.. E j i
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% 'NCl1YDNC73 WWOdlNn
\
l e i l l i I i w - 4 . 1 a : L I g
-3 ;
I i l I o : i l.- 1 7:gr 7:wenaruns. *c ,, I FIGURE 4.1,5.A-4 t PERCENT REDUCTION OF AREA FOR SHORT-TRANSVERSE SPECIMENS IRRADIATED TO 4.3 x 1018 N/CH2 (E > 1 MeV) l l'
+,enac.2 a a a K-& a. 4-am esa a s aasuu,.s ama41-*&-+A4*'M Maca.ema m =A-alisb=o.mAnm- u A.= bem a44. b aE-&& dinmaK&Ask hmb M & A- esA 4-- 4 =-4=-n2 44-+mb-nw--6 na- a b ha c, e.r-+-=A 1
i a v 4 4 i s WiAWSN O o e
)
E o v 200*C [ u IRRACIATED , p w G G em I - I : g 8 aa -
' wiaumime
*I A ^
; 1
, we -
, , & ImRADIATED ,
& A A I 9
g l 0 300 400 000 800 I MYDRCGEN. PPM 1 FIGURE 4.1.5.A-5 EFFECT OF HYDROGEN CONCENTRATION ON THE REDUCTION OF ! AREA FOR ZIRCALOY-2 IRRADIATED TO 1020 N/CM 2 '
- 26
L 1 t l
- 1. , , , , ,
DEFORMATION TEMPERATURE 572.$99 4
]
LSWORM TOTAL
- 8 9 PLEL CLADDMG.SLRST TESTS g, . X 9 PLEL CLADDMG, TENSLE TEST . ;
?
- Om TLSE DOGSONE TENSLE TEST j & & QUDE TWE RmG TENSLE TEST g ;
l t l: i a ! 3R I.
, x TOTAL STRAW N !
N: : N . = . T . LSWORM STRAW O === . O S b """- ===== ==- . . A u 9 **'*mmmm % % 9 i e a a e gg @ 8 7 8 9 10 11 12 2 PLUENCE 1981 n/sm ,E>0.821 MeV FIGURE 4.1.5.A-6 FLUENCE DEPENDENCE OF STRAIN FOR IRRADIATED ZIRCALOY-4
- a l
4.1.6.4 Fission Gas Release i The following section supplements Section 4.1.6 of Reference 1. i 4.1.6.1.a Fission Gas Release i The calculation of fission gas release is an integral part of the fuel 1 perfomance calculations involving the temperature distribution and internal pressure of fuel rods. The release of fission product gases plays
. an important role in the calculation of gas conductivity and, therefore, ,
affects the transfer of heat from the U02 pellets to the cladding. C E's l current model for these calculations (FATES 38) was submitted to the NRC in - 1986 (Reference 4.1.6.a-1) and received NRC approval in early 1987 l (Reference 4.1.6.a-2) . The FATE 538 fission gas release model was developed utilizing data from low and high power rods with burnups ranging from 6.5 to 61.5 WD/kg and measured releases of 0.3 to 48.1%. The model includes the results of fission gas release measurements performed on test ; rodlets that were irradiated in a PWR and subsequently ramped to high linear heat rates. Comparisons between measurements and FATES 3B predictions are given in Reference 4.1.6.a-1.
- Additional extended burnup data on fission gas release has been obtained
, since the publication of References 1 and 4.1.6.a-1. These data consist of six-cycle Fort Calhoun (49.7 to 55.7 WD/kg) rods and five-cycle Zion-1 ' rods (54.3 to 59.4 WD/kg) (References 4.1.6.a-3 and -4). All of these i fission gas release measurements were low (less than 2.8% at burnups up to 59.4 WD/kg). These data also show no significant enhancement of fission gas release with burnup at normal operating levels. These data are i presented in Tables 4.1.6.a-1 and 4.1.6.a-2. Microstructural examinations of the Fort Calhoun rods showed the formation ofaporousrim(75to80%TD),150-250 microns thick (References 4.1.6 a-3 : and -5). This porous rim can result in a decrease in local fuel thermal conductivity and thus an increase in pellet temperature. C E believes that this porous layer is a phenomenon associated with local burnup and is well behaved. (
t i
) This increase is not considered significant in low power, high +
burnup fuel. In addition, other high~ burnup effects 'are known' to offset i the temperature increase due to the porous rim. Two such important effects are [' !
) Thus, it is concluded that the effects of a porous rim can be neglected for burnups of up to 60 MWD /kg. ;
Hiah Burnun Data comearisons . The predictive capability of the FATES 38 fuel performance code, was demonstrated with respect to fission gas release by comparing code ; predictions with experimentally measured data in Reference 4.1.6.a-l. The ; high burnup data sets (at and above 50 MWO/kg rod average burnup) analyzed as part of the FATES 3B correlation and verification data bases were characteristic of fission gas release data in the high burnup and high-temperature regime. Additional extended burnup data on fission gas released by test rods (typical of fuel rods operated in C-E designed commercial reactors) have been obtained since the publication of Reference 4.1.6.a-1. Comparisons of these data to FATES 3B ' predictions are presented ' in Table 4.1.6.a-3. These data comparisons provide additional support for FATES 38 fission gas release predictions in the high-burnup, low-temperature ! (low power) regime. These data are described below. 1 Calvert Cliffs Data:
)
High-burnup performance evaluations of Zircaloy 4 clad test fuel rods and i l 'all Zircaloy' fuel assemblies were performed on fuel irradiated in Calvert Cliffs 1. The evaluations were sponsored by combustion Engineering in I conjunction with the Electric Power Research Institute (EPRI) (Reference 4.1.6 a-6). A total of 60 test fuel rods were fabrii:ated for this experiment and were equally distributed among three reconstitutable Batch B assemblies. Fission gas release data comparisons were perfonned for 16 of I l these test rods, with and of life rod average burnups ranging from 18.7 to 44.4 MWD /kgU, in support of the FATES 3B verification effort (Reference 4.1.6>a-1). Five of the modern design test rods, propressurized l l 29-i l
p L
]:
l-rods with modern design non-densifying pellets, were irradiated one i j additional (fifth) cycle to burnups of 49.4 to 54.1 MWD /kgU. The fission ; gas released by the fuel in these rods was measured. A comparison of the l measured gas releases with FATES 3B predicted sas releases-for-thesedive-~ t test rods is presented in- Table 4.1.6.a 3. On the average, FATES 3B [ ] l Fort Calhoun Data: : 4 The Fort Calhoun extended burnup demonstration program was sponsored by the , f Department of Energy (DOE) to demonstrate the performance of C E's standard ! 14x14 fuel design at ' extended burnups (Reference 4.1.6.a-7). Hot cell ; examination work on some of the test rods irradiated through six cycles was performed in a follow on program jointly sponsored by DOE, the C-E Owners , Group, and C-E (Reference 4.1.6.a-3). Fission gas release data comparisons I were performed for four of the 'most highly burned test rods (54.6 to 55.7 s MWD /kg rod average burnup). These four rods resided in positions very , close to each other in the same gnadrant of Assembly D005 through the entire irradiation period. A single FATES 3B case was generated using , design input parameters and an irradiation history that appropriately models all four test rods. The comparisons of measured gas released and - the FATES 3B predicted gas release are also presented in Table 4.1.6.a-3. ; On the average,' FATES 3B [
]
Conclusions:
Additional data comparisons have been performed on fuel rods typical of C E ) current generation fuel that were irradiated under normal low temperature conditions during extended burnup operation to rod average burnups of up to i 55.7 MWD /kg. In general, the low temperature release due to knock-out and
*~
recoil is [ ] at 60 MWD /kg. However, releases j associated with knock-out and recoil are low. Therefore, it can be concluded that FATES 3B adequately models, on a best-estimate basis, the fission gas release of extended-burnup fuel operated under normal conditions in C-E designed commercial reactors. l 1
I l
-]
l 4.1.6.2.a Evaluation of Fission Gas Release The discussion in Section 4.1.6.1.a surveys the situation at C-E with ! respect to the data available and the modeling of the fission gas release of fuel burned to extended burnups. Significant strides have been achieved , in the area of normal operation and in the area of response to ramps. The I conclusions are: (1) Commercial fuel rods operating in PWRs with helium propressurization and nondensifying fuel have been examined and consistently found to . contain very low levels of released fission gases to burnup levels of N
-60 MWD /kg. The relative absence of significant enhancement due to burnup .at normal operating levels is now verified by direct measurement.
I (2) Data evaluated by CE support the FATES 3B model to burnups of 60 MWD /kg. Furthermore, the trends observed in all 002 behaviors are gradual and support the orderly extension of the allowable burnup. 4 9 D t k l l l
. gypr . . . -
1 Table 4.1.6.a-1 FISSION-GAS RELEASE DATA FROM FORT CALHOUN FUEL RODS Rod Time-Avg. Total Gas Rod Burnep. Volume (Xe+Kr) Volume (Xe+Kr) SL Fission Heat Rating, Collected, Released, Generated, Gas'
- i. Number MWD /MIU ikw/ft1 cc SIP i
cc STP cc STPfal Released KJD008 51500 5.38 794.1 23.4 3425.1 0.68 l KJ0015 51400 4.98 769.3 19.2 3417.4 0.56 KJE051 55700 5.36 753.8 46.5 3704.7 1.26 KJE077 55400 5.39 736.5 49.1 3684.6 1.33 i KJE052 54600 5.25 767.6 33.0 3633.1 0.91 KJE006 49700 5.23 737.8 31.6 3309.3 0.95 g, KJD072 53400 5.12 755.2 22.1 3549.6 0.62 O' KJD075 51500 (b) 747.1 , 22.9 3429.3 0.67 i KJE109 52900 5.26 735.8 46.1 3517.2 1.31 KJE068 52600 (b) 751.7 27.4 3500.9 ' 0.78 KJE089 53100 5.50 745.0 40.7 3534.7 1.15 KJE088 52900 5.45 730.5 36.6 3516.6 1.04 l t (a) Assumes production rate of 30 atoms of (Xe+Kr) per 100 fisstens and 200 MeV/ fission. (b) Physics Data used to calculate the time-average heat rating are not available for these rods. t i e s _ _ __ _ __________ _ .__ _ ________ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _- _ _ _ _ _ _ . - __- . _ _ ___ , _ _ _ _ . + .< s _ .- . - ~ . _ , _ _ -- +-w_., e_.,, m._ . . . . ,i.--c.%..
Table 4.1.6.a-2 FISSION-GAS RELEASE DATA FRON ZION I FUEL RODS , Rod Avg. Total Gas Xe + Kr Xe + Kr Fission Burnup Collected Release Generated (a) Gas Release Kr/Xe Rod No. ff8dD/kel fcc at STP) fcc at STP1 fcc at STP1 ft) Ratte A10 55.90 795.25 63.30 3711.91 1.71 0.098 G12 58.25 778.23 91.29 3867.62 2.36 0.107 612 55.34 779.65 43.74 3674.15 1.19 0.115 , 613 57.99 811.62 70.61 3854.77 1.83 0.121 618 55.32 773.01 42.98 3672.56 1.17 0.097 620 55.29 772.67 45.28 3658.88 1.24 0.087 l 638 54.26 783.13 51.76 3606.58 1.44 0.111
- , 640 56.06 777.38 48.20 3737.39 1.29 0.099 f y 648 56.60 773.73 43.25 3774.12 1.15 0.116 650 58.80 851.89 105.72 3903.94 2.71 0.128 I 653 59.43 804.34 90.01 3950.02 2.28 0.108
( 657 54.78 767.56 38.61 3636.84 1.06 0.105 659 56.23 764.05 54.09 3721.29 1.45 ( 0.099 l 665 55.78 785.46 52.47 3707.75 1.42 0.123 < 679 56.62 744.28 49.94 3771.20 1.32 0.008 681 55.01 768.30 41.80 3675.09 1.14 0.103 l 683 56.32 754.43 50.17 3762.88 1.33 0.099 i 685 55.35 766.59 44.08 3667.13 1.20 0.093 l l 693 55.98 782.32 41.38 3704.94 1.12' O.086 6% 58.13 816.72 73.75 3863.88 1.91 0.116 i l ' 697 56.37 738.07 38.08
- 3727.14 1.02 0.112 l (a) Assumptions: Fission Gas Yield - 0.3 atoms (Xe + Kr) per fission Energy Release - 200 MeV/ fission -
l ;
Table 4.1.6.a 3 FATES 3B Predictions of Gas Release from High Burnup, Low Power Test Rods Predicted-Rod Average Measured Predicted Measured Burnup Gas Gas Gas Bad HMQZia Release Release % Release % Calvert Cliffs-1 --- --- SN24 49.4 1.16 SN34 49.4 0.67 SN36 49.5 1.00 SN45 54.1 >2.02 SN59 49.7 1.03 Fort Calhoun Extended Burnuo fuel XJE051 55.7 1.26 , KJE052 54.6 0.91 XJE077 55.4 1.33 XJE109 54.6 1.31 D m
l 4.1.7.4 Fuel Themal Conductivity The following paragraph appends subsection 4.1.7.3 of Reference 1. 4 .1. 7.3. a Evaluation of Fuel Themal Conductivity No new data on the thermal conductivity of irradiated fuel has become l available since the publication of Reference 1. However, the performance j of fuel rods to 60 MWD /kg (References 4.1.7.a 1 and 2) indicates no trend i toward serious degradation of themal conductivity. The ability of the q FATES 38 model to predict the measured gas release data suggests that any l degradation in local fuel thermal conductivity, such as due to the i formation of a porous rim, is implicitly accounted for in the FATES 3B model. This is thought to be accomplished by the density correction in the fuel thermal conductivity equation and through the conservatism that exists in the other parts of the relevant submodels used in the fission gas - release calculation. It ts therefore concluded that the current thermal conductivity equations are adequate to 60 MWD /kg. ' t o k a l. l l b l l
l l l
.4.1.8.4 Fuel Meltina Temeerature l
The following paragraph appends Subsection 4.1.4.1 of Reference 1. j l 4.1.8.1.a Modeline of Fuel Meltina Temeerature and Effect of Increased q New data continue to support the conservatism of the melting point ; expression. The range of the molting point deteminations of unirradiated
. 00 fabricated by C E (5094 5173*F) performed at Pacific Northwest Labs 2 ,
{ (Reference 4.1.8.a-1) exceeds the melting point calculated by the ! expression for unirradiated fuel (5080'F). Work reported by Komatsu, et ; al. (Reference 4.1.8.4-2) showed no effect of burnup on 00 irradiated up 2 to burnups of 30 MWD /kg, and a drop of only -2'F/ MWD /kg for UO 20%Pu0 2 2 irradiated up to burnups of 110 MWD /kg. Thus, it is concluded that the melting point expression is adequate to 60 MWD /kg. i l I
- l l
h i l l 36- _. _ _ _ . . _ . _ . . _. . . _ . _ _ . . _ _ _ _ . _ _ _ . _ _ _ ___ l
4.1.9.4 Fuel Swellina The following paragraph appends Subsection 4.1.9.3 of Reference 1, 4.1.9.3.a Evaluation of Fuel swellina Data for six cycle fuel rods from Fort Calhoun and five-cycle fuel rods
.' from Zion 1 (References 4.1.9.a-1 and -2, respectively) have become available since the publication of Reference 1. Fuel density measurements
... were made on pellet sections with a local burnup of 60.4 MWD /kg from Zion 1 , and 63.3 MWD /kg from Fort Calhoun. These data and lower burnup data from previous cycles of these reactors indicate a swelling raf , of i 0.53%/10 MWD /kg for Zion 1 and 0.70%/10 MWD /kg for Fort Calhoun, which is entirely consistent with the 0.4-0.8%/10 MWD /kg data measured previously , for Fort Calhoun and Calvert Cliffs-1. These results show no enhancement of the fuel swelling rate for local fuel burnups up to 63.3 MWD /kg, ; indicating no change in the fuel swelling mechanism up to this burnup level. Consequently, the current FATES 3B model is valid to these high , burnups.
~
t a
I i
- 4.1.10.a Fuel Red Bow !
1 The discussion provided in Section 4.1.10 of Reference 1 applies to the l proposed increase in the 1 pin burnup limit to 60 MWO/kg.. Rod bow is not a j concern for high burnup fuel rods since their power falloff more than ; compensates for their rod bow penalty. l l t 6 e
)
?
?
[ k 1 4 i i 4.1.!!.a Frettina Wear The discussion provided in Section 4.1.11 of Reference 1 applies to the - proposed increase in the 1-pin burnup limit to 60 WD/kg. No signit'icant fretting wear has been seen during extensive inspections of C E fuel rods { and the degree of stress relaxation of the grid springs and creepdown of the fuel rod changes very little after one operating cycle. e
?
i i e l l-L 1 L
-3g-i
4.1.12.a Pellet /Claddina Internetten The following section replaces Section 4.1.12 of Reference 1. C E has been involved in many ramping experiments and has collected a considerable amount of PCI data. The data plotted in Figure 4.1.12.a-1 came' from rodlets pre-irradiated at Obrigheim and ramped at either the Petten or Studsvik test facilities in Europe (References 4.1.12.a-1, -2, . -3, 4). The data shown are only from rodlets using the standard C E or KWU designs. Other data available in the literature have not been shown because of design differences. It is important tu recognize that comparisons between experimental PCI results are only valid when the important design variables are consistent. All of these rodlets were preconditioned in a PWR at similar power levels and were ramped under PWR conditions at relatively fast and consistent rates (50-110 W/cm/ min). Data are also available at slower ramp rates. The slower ramps are less severe and give improved PCT performance. [
] In addition, as burnup increases, - ~ the capability of the fuel to reach the power levels needed for PCI failure is diminished. This fact (
J
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! 4.1.13.a Claddine Defomation & Ruoture The discussion provided in Section 4.1.13 of Reference 1 applies to the -
proposed increase in the 1-pin burnup limit to 60 MWD /kg. It has been i determined that the LOCA models for cladding deformation and- rupture are ! adequate for use at 60 MWD /kg. [ l
) ,
e k i ( 1 ! ) l.
- l. .
l 42-
P 4.1.14.4 Fuel Rod Growth i The following replaces subsection 4.1.14 of Reference 1. l t It has been well established that Zircaloy-4 clad rods exhibit axial ; elongation or growth when continuously exposed to a neutron flux. A
}
substantial amount of growth data has been obtained on PWR fuel rods of modern design (i.e., pressurized rods with nondensifying fuel) at burnups [ ). This information has been used to modify the fuel rod growth models originally developed with data obtained at lower fluences and , : from rods of older design (densifying fuel with lower initial pressurizationlevels). 4.1.14.1.a Modelino of Fuel Red Growth ' The overall elongation of a Zircaloy clad fuel rod is due to several contributing mechanisms including stress free irradiation growth of the Zircaloy cladding, mechanical interaction between the U0 fuel pellets and 2 , the Zircaloy cladding, and a net positive growth component due to creepdown ' of the cladding under the external coolant pressure. Each of these _ contributing mechanisms are related to the time of operation through accumulated burnup or fluence. Rather than account for individual contributions from each mechanism, overall fuel rod growth is measured and empirically modeled for design purposes. l Growth strain versus fluence (E>0.821 MeV) is linear on a log-log plot. The functional form of such an equation is: ( = A (dt)" where 'e = strain, percent. !
> (t = neutron fluence, n/cm2 (E>0.821 MeV) x 10 21 )
A and n = constants, as shown below. A regression analysis was used to determine the value of the constants A and n and resulted in the following growth eq'Jations:
.- . - - . . - . _ - . . - - - - - - -. . a
i i Upper 955 tolerance limit: c = ; Best estimate equation: ( = . Lower 95% tolerance limit: ( = The growth data used in this analysis covered a fluence range of i t 2
- i 4.1.14.2.a Effect of Ertanded Burnun I l
Rod length measurements performed on rods with fast fluences up to * [ ] have shown - continuous and well-behaved growth with increasing exposure (References l 4.1.14.a 1 through 4.1.14,a8). These data have confirmed that no ! acceleration of the growth rate or other abrupt changes occur up to the ; exposure levels of the examined rods. Furthermore, fuel rod growth at - higher burnups appears to be relatively insensitive to slight design ; differences. [ l t
] co not contribute as much to the overall growth rate at .
higher exposures as would be inferred from measurements taken after only - one or two operating cycles. This observation is supported by measurements taken as part of ' fuel performance evaluation pragrams at Fort Calhoun, l Calvert Cliffs-1, and Arkansas Nuclear One-Unit 2 (References 4.1.14.a 3, ;
-5, -6, 7, -8). !
4.1.14.3.a Evaluation of Fuel Red Growth l l
- Figure 4.1.14.a-1 shows growth measurements obtained on C E fuel rods compared to the CE fuel rod growth model described in Subsection L._
4.1.14.1.a. Data from 14x14 fuel rods at Calvert Cliffs-1 and Fort Calhoun have been obtained for fluences of up to [ ] (References 4.1.14.a-6, -8) while data from 16x16 fuel rods at Arkansas Nuclear One-1 44 1
~ _ . .._ _ _. . _ . _ _ _ _ _ _ . . _ . _ _ _ _ _ - - _ - _ _ _ _
p: i' l Unit- 2 have been obtained to fluences of ( ) (Reference 4.1.14.a-7).- The growth data from the Calvert Cliffs-1 fuel rods have also been used in an analysis of growth published by Franklin which involved more than 700 fuel rod length measurements (Reference 4.1.14.a-9). This analysis confirmed the well-behaved nature of fuel rod growth at high fluence and
.( ).-
The database shown in Figure 4.1.14.a-1 includes measurements from ANO-2 - fuel rods that showed higher growth than other rods in the same batch. The higher growths are believed to be related to the relatively high ca.rbon content of the cladding. A similar association between the carbon content of cladding and fual rod growth was also reported in the 1988 ANS. Topical Meeting'on LWR Fuel Performance by Fragema . describing performance of fuel rods irradiated in TN1. [ e
]
l
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. . - .w .. -
g _
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l . l '. ' i- , Figure " 4.1.14.a - FUEL ROD: GROWTH MEASUREMENTS COMPARED TO C-E ZlRCALOY FUEL ROD GROWTH MODEL - ; - _Z
\.
Z
- t
<3 l
E _Z
< +
e K A t-
, a un <
t I 6-3: o O 1 FLUENCE X E-21 n/cm', (E>0.821 Mev) e 9 1 _ - , . . . c.. _ _ . . . , ~ . . . . _ , , , . . . - ,m, ., , , _ _ . . . . - . . . . . . . _ - .. . . . . ~ , . . . . . _ , . . . . . . . ., , _ _ . , . . . . . ~ , _ , . . ,
[ h p L 4.2.1.a Guide Tube Wear o The discussion provided in Section 4.2.1 of Reference 1 applies to the proposed increase in the 1-pin burnup limit to 60 MWD /kg. An extensive program was initiated in response to the detection of guide tube wear. This program resulted in the development of a guide tube wear sleeve design that essentially eliminated the concern of guide tube wear (Reference 4.2.1.a-1), 4 '*- The current St ;ct;e .. fuel assea.bly design incorporates this guide tube sleeve design w O l l e l 1
v 4.2.2.a Fuel Assembly Lenoth Chance and Shoulder Gao Changa Although most of this section is identical to Section 4.2.2 of Reference 1,
- inclusion of just the changes could be confusing. Therefore, this section replaces Section 4.2.2 of Reference 1 in its entirety.
Fuel assembly length change results from two distinct mechanisms in the .
- Zircaloy guide tubes: irradiation induced growth and compressive creep.
Growth is produced by radiation effects on the Zircaloy crystalline structure,
" and causes the guide tubes to elongate. Compressive creep is the permanent e reduction in length of the guide tubes in response to a net holddown force on the fuel assembly structure.
Change in guide tube length affects the fuel assembly engagement with the reactor internals (thereby affecting the holddown force on the assembly) and , the shoulder gap (the distance between the top of the fuel rods and the bottom of the upper end fitting). The icngth change is important in the evaluation of criteria pertaining to each of these aspects of fuel performance. - Since the holddown force is a function of fuel assembly length, irradiation induced guide tube growth causes.an additional compression of the upper end i fitting springs, increasing the compressive load on the guide tubes. The , higher load in turn causes an increased compressive creep rate of the guide tubes._ Therefore, the net fuel assembly length change at a given time during operation requires a time history analysis to properly account for the combined effects of irradiation growth and creep up to that point in time. 4.2.2.1.a Modelina of Assembly Lenath Chance and Shoulder Gao Chance a) Assembiv Lenath Chance H- Growth and creep characteristics are dependent on the metallurgical state of l E the Zircaloy' guide tubes. All St. Lucie 2 reload fuel assemblies are i fabricated with [ . ] guide tubes. The l analytical models presented in Reference 1 for [ ] are repeated in Table 4.2.2.1.a-1. l l l l
- Dimensional, changes of fuel assembly guide tubes are analytically predicted by the SIGREEP computer code, which.is described in Reference 4.2.2.a-1. . The code utilizes a computerized Monte Carlo technique for establishing resultant
. joint probability density functions by randomly selecting combinations of:
' input values to be used in a time history analysis of dimensional changes.
Inputs assigned statistical uncertainties include component. dimensions, the assembly uplift force, the guide tube growth coefficient, and the guide tube creep coefficient. P
?- The SIGREEP computer code generates a set of randomly selected values for the-input parameters that have been assigned uncertainty distributions, and then uses that set of inputs to perform a time history analysis of the guide tube ,
length change. When the analysis reaches the specified operating time or burnup, the' dimensional change prediction for the fuel assembly is complete. A single value of assembly length change is the result of the time history , calculation. The same steps are' repeated (starting with a different set of randomly selected values for the input parameters) until a sufficient number , of cases (typically 2000) have been generated to define a probability histogram of length change at .end of life (E0L). The resultant histogram represents the statistical variation of E0L length change which can be attributed to the uncertainties of the input parameters. Values can be chosen from the histogram at desired probability levels for comparisons to actual data or appropriate design criteria. Figure 4.2.2.a-1 presents a typical histogram of fuel assembly length change.- b) Shoulder Gao chance Shoulder gaps change with residence time in the reactor due to differential i growth between the fuel rods and the fuel assembly structure (guide tubes). Reference 1 described a technique of evaluating shoulder gap change using the g SIGREEP computer code. With that technique, fuel assembly length change is
- calculated by SIGREEP exactly as described above, but for each time history case for fuel assembly length change, fuel rod length is simultaneously calculated using values for the growth coefficient and beginning of life (B0L) 1 l
l 1
i l dimensions that have been randomly selected from the probability distributions for these parameters. When the time history case reaches the specified time or burnup, shoulder gap. change is calculated as the difference in fuel rod and fuel assembly length changes. A single value of shoulder gap change is the j end product of the time history calculation. The calculation is repeated with different sets of randomly selected values for the input parameters until a sufficient number of cases (again typically 2000) have been generated to define a probability. histogram of shoulder gap at E0L. This method of evaluating shoulder gap change is used on 14x14 fuel designs but, because of the high fuel rod growth rate associated with some ANO-2 Batch C fuel rods, an interim approach of deterministically combining a conservatively high fuel rod growth prediction with a conservatively low fuel assembly growth prediction had been used on 16x16 fuel designs, pending more 16x16 measurement data. - Additional fuel rod growth data- are now available and are presented in Section 4.1.14.a, along with an updated fuel rod growth model based on the data. Also included in Section 4.1.14.a is a discussion of the cause of the high growth rates of the ANO-2 Batch C rods and a justification for no longer applying those high growth rates to current fuel designs (i.e., a change in the material specification of the cladding). The interim approach - is, therefore, no longer necessary and the shoulder gap evaluation technique utilizing the SIGREEP computer code with the updated fuel rod growth model of Section 4.1.14.a can be used for 16x16 fuel designs. A comparison of this technique to shoulder gap measurements taken on 16x16 fuel assemblies with ( ) is included in Section 4.2.2.4.a. 4.2.2.2.a Effect of Extended Burnun As stated in the preceding sections, fuel assembly length change is the net change resulting from irradiation induced growth and compressive creep of the y guide tubes. Since growth is fluence dependent and compressive creep is time . and flux dependent, assembly length change and shoulder gap are affected by i N extended burnup. In general, higher burnups are expected to result in greater Jincreases in. assembly length, greater holddown spring compression, and larger ;
-. changes in shoulder gap. The extent of these changes will be evaluated on the' specific extended burnup operating conditions and the particular fuel Jassembly design. J 4.2.2.3.a Evaluation of Assembly Lenoth Chanae Guide tube length change data for 16x16 fuel assemblies with [- ]-
+ wereipresented in Figure 4-25 of Reference 1, along with SIGREEP predictions using the irradiation induced growth equation and axial creep equation from. Table 4.2.2.a-1. Additional length change data for 16x16 fuel assemblies with [ ]'are now available with bundle average burnups of up to [ ] '. These data are compared to SIGREEP predictions in Figure
-4.2.2.a-2.
Inspection of the figure shows that the SIGREEP predictions are in good agreement with the data, both in the magnitude of the predictions and the trend of_ the predictions, and that the upper and lower 95% predictions - ; represent-conservative estimates of the guide tube length changes. Based on the comparisons of the data and the predictions, it is concluded that both the analytical model and the growth and creep equations are acceptable for use in , _ predicting fuel assembly length change' for designs with [ ') to
-extended burnups.
'4.2.2.4.a Evaluation of Shoulder Gao Chance Shoulder gap change data for 16x16 fuel assemblies with [ ] were presented in Figure 4-29 and 4-30 of Reference 1, along with the limitir.g
# shov1 der gap change predictions using SIGREEP with the models given in y Reference 1. The data from Figure 4-30 of Reference 1 is included on Figure
, 4.2.2.a-3 with the limiting shoulder gap change prediction now made by SIGREEP with the models given in Section 4.1.14.a (fuel rod growth) and Table 4.2.2.a-1 (guide tube growth). The data from Figure 4-29 of Reference 1 is l
l not repeated because it includes the high growth rate fuel rods that have been shown to be non representative of current fuel designs (see Section 4.1.14.a). Additional shoulder gap change data for'16x16 fuel assemblies with [- ) are now available with rod average burnups of up to (
). These data are compared to SIGREEP predictions (again using the models ~ of Secticn 4.1.14.a and Table 4.2.2.a-1) in Figure 4.2.2.a 4 (standard 16x16 design with 150 inch active length) and Figure 4.2.2.a 5 (St. Lucie 2 reload design with 136.7 inch active length).
C Inspection of Figures 4.2.2.a-3, 4.2.2.a-4, and 4.2.2.a-5 shows that the analytical predictions represent conservative bounds of the data. Therefore, it is concluded that the analytical models are acceptable for use in predicting the limiting shoulder gap change for designs with ( ) to extended burnups since:
- 1) the models for the individual components of the shoulder gap change (fuel rod growth and guide tube growth) have been shown to be acceptable to high burnups, and
- 2) the analytical predictions for shoulder gap change represent conservative bounds of the. data to high burnups.
1 i . [ ! Table 4.2.2.a-1 ' Analytical Models for L 1 , 7 1.. Irradiation Growth Model 7 Equation Form: e = A (&t)" + B~ where: e = axial growth strain, in/in t 14 = proportionality factor = [ ] , e: 4t = fluence, nyt x 10-21 (E > 0.821 MeV) n = exponential constant = [ ] B =[ ]
- 2. Comoressive Creen Model ,
Equation Form: e -apa
-where:- c - axial creep strain rate, in/in/hr a -= proportionality factor - 5.37 (Best Estimate)
- 7.39 (Upper 95%)
=
3.35 (Lower 95%)
, ,(,)0.85 g(-6000/RT)(Ake-kt + C) 2
& = fast neutron flux, n/cm -sec (E > 1.0 MeV)
R = 1.987 cal /mol *K T = temperature, 'K A. = constant = [ '] ' t = time, hrs k - constant = [ ], hrs'l C = constant = [ ] o - axial guide tube stress, ksi T I . v.
*- '- i 4
FIGURE 4.2.2.a-1 TYP9 CAL PROSA88LITY SHSTOGRAM FOR FUEL ASSEBABLY LE90GTH CHANGE 1 g . i i SIGREEP-GESIERATED SGISTOGRAte FOft &ledITIBIG . NtMMBER FUEL ASSEGABLY 7 CASES f., 1 i'
+-SSAftGIN TO~
191TERFERE00CE OffE. SIDED UPPER SE% .j PROSA84LITY INTERVAL LeseIT FOR FUEL ASSEGABLY 4 LEf8GTit CIIA00GE \ o l
<E GT.IC..A GE. .
EE G1 C..A E -Gol.. FOR If8TERFERENCE ' . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . . _ . _ _ _ _ . _ _ . _ _ _ _ _ _ . . _ __ _ _ _ _ _ _._, . . _ _ _ ~ _ _ _ . _ _ _ . _ _ _ _ _ . . . . . _ _ _ _ _ _ _ _ _ . _ . _ _ _ _ . . _ _ _ _ _ _ _ . _ =
7 p . , ~ _,..
=
FIGURE. 4.2.2.a--2 -4 COMPARISON OF 16x16 [ GUIDE TUBE
. LENGTH CHANGE TO SIGREEP PREDICTIONS. __
4 12i td O
. k 9 5 m
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. .y . ,
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s FIGURE 4.2.2.a-3 COMPARISON OF ANO-2 BATCH D SHOULDER GAP CHANGES TO SIGREEP PREDICTIONS * :
- Note: AND-2 Batch D used
[ ;J guide. tubes Y N O Z
. 4
- u. I m
e u ' 1 ' 4 O i ll ac w O ' 2 a 3 O i I to l e 6 s ) 6 FUEL ROD FLUENCE, wt a 1921, E > 0.8218WieV i
- y. ~. '.. - :. ~
g - , FIGURE 4.2.2.a-4
~ COMPARISON ~ OF . SHOULDER GAP' CHANGE TO .SIGREEP ' PREDICTIONS
- FOR 16x16 ' FUEL' ASSEMBLIES- WITH [- ] GUIDE TUBES - . ._
~
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FIGURE 4'.2.2.a--5 ' ~~
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COMPARISON- 0F : SHOULDER -GAP CHANGE-TO SIGREEP PREDICTIONS FOR.1ST. LUCIE 21 FUEL ASSEMBLIES WITH. [- ]. GUIDELTUBES
,i, 15
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' 4.2;3.a- Fuel- Assemb1v Holddown .
The discussion' provided in Section - 4.2.3 of Referenca 1 applies to- the - proposed increase in the 1-pin burnup limit to 60 MWD /kg. The holddown
' spring relaxation due'to extended burnup tends to be offset by concurrent growth of the fuel assembly.
e, i i i i i
'l
=; i l
). -
M
i l I 4.2.4.4 Grid feradiation Growth l The discussion provided in Section 4.2.4 of Reference 1 applies to the -
)
proposed increase in the 1-pin burnup limit to 60 MWO/kg. Since the grid
- j. growth data presented in Reference 1 agreed well with all the other growth i measurements (- -] presented in that reference, the good
, agreement between the growth measurements and predictions for [
] presented in Reference 1 supports the adequacy of the grid irradiation growth model to extended burnup.
De I 4
4.2.5.a Soacer Grid Relaxation The discussion provided in Section 4.1.13 of Reference 1 applies to. the , proposed increase in the 1-pin burnup limit to 60 MWD /kg. Thc degree of stress relaxation of the grid springs and creepdown of the fuel rod changes very little after- one operating cycle. Also, the observation of .! superior performance of the grids in the extended t.urnup _ demonstration :
- l. assemblies irradiated. in Calvert Cliffs Unit I and ANO 2 confirm that the ,
relaxation of the fuel assembly spacer grid springs is not a concern for I
, the extended burnup operation of 14x14 or 16x16 fuel assembly designs.
9 O' ( .
a y
- 4.2.6 a [orrosion of the Fuel Assemb1v Structure .] ^ The following paragraphs append Subsection 4.2.6.3 of Reference 1. .
L 4.2.6.3.a Evaluation of Corrosion of the Fuel Assembiv Structure Additional in reactor corrosion data will be obtained from hot cell
- examinations -(metallographic and hydrogen content analyses) to be performed on a five-cycle Calvert Cliffs-1 assembly cage that experienced an assembly average burnup of ( ). Detailed .
poolside visual examinations were performed on this assembly. No indications of anomalous behavior, such as oxide spalling or structural , cracking, were observed. The hot cell data, which will be available in 1989 or 1990 from a joint EPRI, BG&E and C-E program, are expected to support the- current model which predicts the oxide layer thickness to , increase monotonically with time. On review of the available information, it is concluded that, for the coolant conditions typical of ANO-2, the corrosion of the . Zircaloy-4 structure will not preclude the operation of C-E 16x16 fuel assemblies to 1-pin burnups of 60 MWD /kg. For reactors with higher coolant temperatures - and coolant chemistry conditions differing from ANO-2, such as higher lithium concentrations, further evaluations of the assembly structure corrosion behavior would have to be made, b
~
l 65-9'
= - ___ . .
i r 4.2.7.a Burnable Poison Rod Behavior The following subsections replace the corresponding subsections of
- Reference 1.
L L 4.2.7.1.a Modelino of Burnable Poison Rod Behavior 4 Alg0.-B4C Pellet Swellina. The swelling of the burnable poison material, induced by irradiation, results in dimensional changes which can affect cladding strain and poison rod void volume. The neutron absorber material-employed in ' the poison rods is in - a pelletized form and _ consists of a
- dispersion of boron carbide (84 C) par.ticles in an alumina (A10 23 ) matrix.
The 84 C content is established by core neutronic requirements and has ranged to levels on the order of 4 wt%. The dimensional changes of the pellet are predicted by a model which assumes [ t J. Since the A1 230 swelling is the dominant contributor to pellet swelling at high exposure, the A10 23 -B4 C swelling is related to fast fluence in the model. It is recognized; however, that the swelling 'of B C is a- function 4 of thermal flux to the extent that it depends upon the B-10 (n,cr) Li-7 reaction. In relating pellet swelling to irradiation exposure, it is assumed [ J. The B 4C swelling rate used is the same as in C-E's model c for B 4C swelling -in a control element assembly (CEA) as described in Reference 4.2.7.a-1, i.e., a volumetric swelling of 0.3% per percent B-10
- f. burnup. The A10 swelling behavior is - based on the data reported by 23 Keilholtz and Moore for high density (> 99% TD) pellets
-(Reference 4.2.7.a-2). Since Al 23 0 swelling is caused by fast neutron irradiation damage, Keilholtz and Moore correlated their observed Al 0 23 I
i l volume increases with fast fluence-(E > 1 MeV). I A review of the data reported by Keilholtz and Moore (Reference 4.2.7.a-2) l indicates that for gross overall dimensional changes, a two-stage swelling rate model is an appropriate representation for Al 0 swelling. That is, l 23 2 above a - fast fluence of approximately 2.6 x 10 21 n/cm , the swelling of Al23 0 is enhanced by microcracking and grain boundary separation which [ -causes a sharp increase in the apparent overall swelling rate. This l-enhancement of swelling was incorporated'into the previous model'which was
- described in Reference 1. However, since the volume created by ..
microcracking accommodates the gas inventory in the rod, this enhancement of swelling does not reduce- the poison rod internal void volume available to the~ gas inventory. Thus, the more accurate model of void volume reduction due to Al23 0 swelling is represented by the following expression-that- accounts for the matrix swelling of Al230 only: The model assumes that swelling is independent of temperature since poison pellets are not expected to exceed an operating temperature of 500*C in PWR applications. Further, Keilholtz and Moore found no significant temperature dependency for Al 230 swelling in the range of 300 to 600*C. ( I
]a .
two-stage model is used for the composite Al23 0 -84 C pellet cwelling model. The volumetric swelling rate for B C 4 (i.e., 30% at 100% B-10 depletion) was used in conjunction with Equation (1) for A1 230 to arrive at the following i expressions for the volumetric swelling of the compo:ite A10 23 -8 4 C pellet. l l s - _ b [ o , t t The above relationship for swelling as a function of ' fluence,is plotted in Figure 4.2.7.a-1 for A10 -8 C with a B C content of 3 wtL Also plotted 23 4 4 are volumetric swelling values calculated -from- diametral swelling data-which were obtained in:C-E sponsored-post-irradiation examination programs to verify, the. performance of the Al23 0 x and A10 23 -B 4 C pellets. These -. data consist of direct diameter measurements ~on 42 whole A123 0 -84C and 16 whole A1 02 3 pellets which 'were from poison rods discharged after = 1 cycle of exposure. The results of the post-irradiationLexamination of:these 1-cycle 'i A1230 -84 C pellets substantiated the assumption of isotropic swelling i behavior (i.e., equal axial and diametral swelling rates). It was also-found that , swelling was independent of initial pellet density in.the'- ; density range of 85 to 98% TD. In addition, indirect diametral swelling-
. data- were obtained, at higher exposures, by profilometry measurements on-
[ p unpressurized burnable poison rods of the early 14x14 design (described in Table 4.2.7 a-1) discharged after 2, 3 and-4 cycles of reactor irradiation. The pellet dia'netral swelling in these rods was inferred by conserv'atively assuming that the Zircaloy-4 cladding had crept down to contact the pellets. This approach had the advantage of directly determining the mechanical performance characteristics of interest at high fluence: (1) the . _ .
cladding strain. as affected by pellet swelling and (2) by inference, the restrained swelling behavior of the Al23 0 -B4 C pellets. It was found that ' even after 4 cycles of reactor operation, the average cladding strain was - still- negative, exhibiting only a slight tendency to be less negative than , the 1-cycle value. Moreover, after 4 cycles, the cladding had completely crept down to contact the pellets and conformed to the pellet shapes as 3 shown by the profile traces. The inferred Al 0 -8 C pellet swelling in .; 23 4 these rods, shown- in Figure 4.2.7.a-1, was calculated from the irradiated diameter profiles, the . as-fabricated cladding wall thickness, and the , as-fabricated pellet diameter. It should be noted that, because of the different measurement - techniques, the 1-cycle pellet data represent an , unrestrained condition, while the higher exposure data derived from rod profiles represent a restrained condition. L A comparison of the performance data with the model in Figure 4.2.7.a-1 indicated the following: o The swelling of Al23 0 -B4 C pellets, as well as that of Al 02 3 pellets, that occurred during the first-cycle of irradiation up to a fluence of about 3.5 x 10 21 n/cm2 (E' > 1 MeV) are reasonably predicted by Equations (2) and (3). The data scatter indicated that several 1-cycle Al23 0 -84C pellets apparently swelled more than predicted by the model, most likely due to pellet microcracking. o There was no measurable diametral swelling of the pellets contained in the early 14x14 design burnable poison rods exposed to additional irradiation up to 4 cycles, equivalent to 8.2 x 10 21 n/cm 2 (E>1MeV). The reason for the lack of apparent diametral swelling is related to the following overall swelling behavior mechanisms:
- h. (a) 8 4C particle swelling caused by the B-10 (n,a) Li-7 reaction .,
induces microcracking and grain boundary separation in the pellet structure. 1 l l
._-___--_______--_N
1
. k e ,
(b). The resulting early apparent swelling (while the B-10 is -) depleting) is enhanced by this void contribution when the pellet )
- is .not restrained (This may account for any underprediction of -
)
1-cycleswelling).
~
(c) At higher fluence (i.e., after 100% B-10 depletion) some of these new voids are acconnodating the A102 3 matrix swelling due ' to l cladding - restraint. Once the accommodation is completed, ' diametral swelling, 'and therefore, volumetric swelling, would. E proceed at the swelling rate indicated by Equation (3). .. i The subsections of Gas Releasg> Poisen Rod Growth, and Poison Rod Claddina , Sr.gga of Reference 1 apply to the proposed increase in the 1-pin burnup limit to 60 MWD /kg. Poison Rod Internal Pressure. The internal pressure at operating conditions is predicted by an analysis involving the calculation of the poison rod void volume, gas temperature, and pellet temperature at , operating conditions. Each of the conditions discussed above represents either a time-dependent, fluence-dependent, or power history dependent b mechanism which will produce changes in the poison rod internal pressure through changes in the void volume and the amount of helium released. l 1' J w Calculation of the EOL internal pressure is predicted -for appropriate E0L [ conditions which-include the number of moles of helium-(prepressure plus gas released from the pellets), gas temperature (the 100% depleted poison pellets produce only a small' amount of heat flux due to ganna heating), and L the void volume (reflecting changes due to different temperatures, pellet swelling, poison rod growth, and cladding creepdown). Also, for the extended-burnup reference designs, pellet open porosity at , [' BOL is nonexistent (Table 4.2.7.a-1). i
)
1 I I
I 4.2.7.2.a. Effect of Extended Burnuo en Burnable Poisen Rod Behavior I A1;0;-B;c Pellet Swellina. The swelling of A10 23 '04 C pellets is strongly fluence - dependent; therefore, the mechanical behavior of the burnable poison- rod is affected by extended burnup. While the cladding may not be ) strained because of the large diametral gap in the new designs, the rod void volume will be decreased by the diametral and axial swelling of the pellets. Gas Release. As discussed in Reference 1, helium is generated and released- - primarily in the first cycle of irradiation when the poison rod is ' operating at its highest temperature. Extended burnup, therefore, will,not -? result in significant additional helium release. This behavior has already been verified by gas release measurements on burnable poison rods exposed i for up to a cycles. Axial Growth and Diametral creen. Extended-burnup operation will result in additional elongation of the burnable poison rods. As discussed in Reference 1, the ' growth of the poison rods is no more limiting than- the growth of the fuel rods. 4 The increment of diametral cladding creep associated with extended-burnup operation should be extremely small due to low cladding temperatures and low differential pressure across the cladding dur'ing this period of time. ; Full diametral cor. tact between the pellets and cladding is not predicted so outward creep of the cladding due to swelling of the pellets is not' expected.
, Rod Internal Pressure. Internal pressure will increase during extended burnup operation due to a reduced void volume within the rod caused k principally by pellet swelling. Rod growth and creepdown are second order - -
effects on the void volume when compared to pollet swelling, but are accounted for. No additional gas is predicted to be released from the pellets due to extended burnup . 4.2.7.3.a Evaluation of Burnable Poison Rod Behavior
~
Well defined models exist for all- fluence dependent and time-dependent aspects of burnable poison rod behavior. When used in combination with the design improvements in the extended-burnup poison rod designs, they will demonstrate that there is margin to the strain, clearance, and internal pressure criteria for the poison rods. me 9 f l
.. - ~
l
- 1
1 i v Table 4.2.7.a 1 Burnable Poison Red Details - Extended Extended Early Burnup Early Burnup Parameter 14r14 Damien 14x14 Damien 16x16 Desian 16x16 Desian
. Pellet 0.0., 0.376 0.379 0.362 0.310 0.307 in. -
Cladding 0.0., 0.440 0.440 0.382 0.382 in. . Cladding 1.D., 0.388 0.384 0.332 0.332 In.
- Expressed as a percent of the total pellet volume.
e
yw - F90MRE 4.2.7.e.1 SINELLINGOF As 230 -SqC ,
' ' ' 8 3
. 5 5 g e . 3A.
*h
- R o p es.. awo i
i- 9 Jat oss 3.2 7.5 - REACTOR A'
- O Jat ele 33 L
A405103 2.s REACTORS
- AHS let 2.0 00AASETER 8" -
ESTIh4ATED FeteeA IIEACTOft ASEASURESSENTS REPftESENTSY3 3gentR b IaRoOrnOFR.Es 2.e I e-O A8 23 0 lAVG.OF 4 PELLETSI !F s E
-4
. s 1 a me 8 -
- Ey 4
- 4.5 -
g en
- g ,
- D 00, g
- M 3 ve v 9 I 3.e -
1.0 z i 20 DATA rOINTS - t 1.5 .- o ! Og a j g lesCYCLE _l O 'I l 2 CYCLES 3 CYCLES 4 CYCLES O' 8 8 8 A a a a a a 1 1 2 3 4 5 s a 7 e , ,e 1 FLUENCE,le 32 Wes=2 (E > 1 assep f 1 , t 1 I
CONCLUSION The objective of this report is to justify the validity of C E methods and models concerning the St. Lucie 2 reload fuel assembly design and safety analysis for 1 pin burnups up to 60 MWD /kg. The present C-E licensing document on fuel burnup limits (Reference 1) justifies a 1 pin limit of 52 l MWD /kg. The data presented in this report justify the extension of this 1 pin = limit to the new 1 pin limit required by the implementation of longer fuel l cycles, 60 MWD /kg. As such, the overall and individual conclusions presented in Reference 1 are shown to be valid for the extension of t'he 1 pin burnup , limit to 60 MWD /kg for the St. Lucie 2 reload fuel assembly design. The conclusions of this report regarding fuel assembly length change and . shoulder gap change are applicable to the St. Lucie 2 fuel assembly design employing ( ). , Also, since the various fuel performance topics discussed in Reference 1 have no explicit dependence on batch average burnup, the batch average discharge limit of Reference 1 is no longer required and can be deleted, a REFERENCES 1 CEMPD-269-P, Rev. 1 P, ' Extended Burnup Operation of Combustion - Engineering PWR Fuel," July 1984. 4.1.2.4-1 G. P. Smith, 'The Evaluation and Demonstration of Methods for Improved Fuel Utilization, End of Cycles 6 and 7 Fuel Examinations,' DOE /ET/34010 10, CEND 414, combustion Engineering, Inc., October 1983. 4.1.2.a A. M. Garde, ' Hot Cell Examination of Extended Burnup Fuel From Fort Calhoun," 00E/ET/34030 11, CEND 427, Combustion Engineering, Inc., September 1986. 4.1.2.a-3 G. P. Smith, 'The Evaluation and Demonstration of Methods for Improved Nuclear Fuel Utilization; lith Progress Report," 00E/ET/34013-14, CEND 431, to be issued. 4.1.2.a-4 M. A. Shubert, " Examination of the PROTOTYPE and 1H038 Assemblies After Reactor Cycle 9 in Calvert Cliffs Unit 1," CENPSD 493 P, January 1989. 4.1.2.a-5 E. Hillner and J. N. Chirigos, 'The Effect of Lithium Hydroxide and Related Solutions on the Corrosion Rate of Zircaloy in 680*F Water," WAPD-TM 307, Bettis Atomic Power Lab, August 1962. 4.1.2.a 6 M. Darrouzet, P. Beslu and Ph. Billot, 'Zircaloy Corrosion PropertiesunderLWRCoolantConditions(PartII),"RPX101-01, Final Report, NFIR Report, NFIR RP 01-702, Nuclear Fuel Industry Research Group, October 1987. 4 4.1.2.a 7 L. W. Newman, 'The Hot Cell Examination of Oconee 1 Fuel Rods After Five Cycles of Irradiation," 00E/ET/34212-50, BAW-1874, Babcock & Wilcox, October 1986. 4.1.2.a 8 M. G. Balfour, W. R. Smalley, J. A. Kuszyk and P. A. Pritchett,
' Hot Cell Examination of Zion Fuel Cycles 1 through 4,* Research
~
Report EP8016 Final Report, Empire State Electric Energy Research Corporation, April 1985. 4.1.2.a 9 U. P. Nayak, H. Kunishi and W. R. Smalley, ' Hot Call Examination of Zion Fuel Cycle 5,' Research Report EP80-16, Final Report, Empire State Electric Energy Research Corporation, June 1985. L 4.1.2.a-10 R. 5. Kaiser, R. S. Miller, J. E. Moon and N. A. Pisano, -
- Westinghouse High Burnup Experience at Farley 1 and Point Beach 2," Proc. International Topical Meeting in LWR Fuel Performance Williamsburg, VA, April 17 20, 1988 American Nuclear Society.
4.1.2.a 11 A. M. Garde, " Effects of Irradiation and Hydriding on the Mechanical Properties of Zircaloy-4 at High Fluence,' Paper Presented at the Eighth International ASTM /IAEA Symposium on Zirconium in tte Nuclear Industry, San Diego, CA, June 1988, and to be Published in Special Technical Publication 1023 which will . cover the proceedings of the Symposium. 4.1.3.a-1 M. A. Shubert, " Examination of the PROTOTYPE and 1H038 Assemblies After Reactor Cycle 9 in Calvert Cliffs Unit 1,* CENPSD 493 P, January 1989. 4.1.3.a-2 G. P. Smith, "The Evaluation and Demonstration of Methods for Improved Fuel Utilization,' 00E/ET/34010-10, CEND 414, October
, 1983.
4.1.4.a-1 'CEPAN Method of Analyzing Creep Collapse of Oval Cladding,' - EPRI NP 3966-CCM Volume 5, April 1985.
-77
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4.1.5.a 1 A. M. Garde, ' Hot Cell Examination of Extended Burnup Fuel From Fort Calhoun,' 00E/ET/3403011, CEND.427, Combustion Engineering, . September 1986. i 4.1.5.a 2 L. W. Newman, 'The Hot Cell Examination of Oconee 1 Fuel Rods i After Five Cycles of Irradiation,' 00E/ET/34212-50, BAW-1874, Babcock and Wilcox, October 1986. I . 4.1.5.a-3 U. P. Nayak, H. Kunishi and W. R. Smalley, " Hot Call Examination of Zion Fuel, Cycle 5,' WCAP-10543, Final Report EP8016. Empire l State Electric Energy Research Corporation, June 1985. -F i 4.1.5.a-4 M. G. Balfour, W. R. Smalley, J. A. Kuszyk and P. A. Pritchett,
" Hot Cell Examination of Zion Fuel Cycles I through 4,"
WCAP-10473, Final Report EP80-16, Empire State Electric Energy Research Corporation, April 1985. l 4.1.5.a-5 A. M. Garde, ' Effects of Irradiation and Hydriding on the Mechanical Properties of Zircaloy 4 at High Fluence," Paper Presented at the Eighth International ASTM /IAEA Symposium on - Zirconium in the Nuclear Industry, San Diego, CA, June 1988, and to be Published in Special Technical Publication 1023 which will cover the proceedings of the Symposium. 4.1.5.a 6 System 80 D Standard Safety Analysis Report Final Safety Analysis Report (CESSAR FSAR), STM-50 470 F, Combustion Engineering, Inc., October 1978. ' g 4.1.5.a-7 J. F. McLehan. ' Yankee Core Evaluation Program Final Report," WCAP-3017-6094, Westinghouse Atomic Power Division, January 1971. 4.1.5.a 8 R. L. Knecht and P. J. Pankaskie, 'Zircaloy 2 Pressure Tubing," BNWL 746, Battelle Pacific Northwest Laboratory, December 1968.
4.1.5.a 9 L. M. Howe and W. R. Thomas, 'The Effects of Neutron Irradiation on the Tensile Properties of Zircaloy-2,' AECL-809, Atomic Energy of Canada Ltd., March 1959. 4.1.5.a 20 J. E. Irvin, ' Effects of Irradiation and Environment on the Mechanical Properties and Hydrogen Pickup of Zircaloy,' Zirconium and Its A11ovn. Electrochemical Society, New York, NY,1966.
. 4.1.5.a-11 W. Evnns and G. W. Parry, 'The Deformation Behavior of Zircaloy-2 Containing Directionally Oriented Zirconium Hydride Precipitates,' Elaetrochem. Tech.. 4,225(1966).
l 4.1.5.a-12 W. A. Pavinich and T. P. Papazoglou, ' Hot Call Examination of Creep Collapse and Irradiation Growth specimens - End of Cycle 3," LRC 4733-8, Babcock and Wilcox Co., March 1980. 4.1.5.a-13 F. A. Nichols, " Evidences for Enhanced Ductility During Irradiation Creep,' Mater. Sci. Ena. 6, 167 (1970). 4.1.5.a-14 E. F. Ibrahim and C. E. Coleman, 'The Effect of Stress
- Sensitivity on Stress Rupture Ductility of Zircaloy 2 and Ir 2.5 wt% Nb," can. Met. Quart.. 12,285(1973).
4.1.5.a 15 E. F. Ibrahim, " Creep Ductility of Cold Worked Ir 2.5 wt% Nb and Zircaloy-2 Tubes In-Reactor,' J. Nucl . Mat. . 96,297(1981). 4.1.5.a-16 0. S. Wood, "High Deformation Creep Behavior of 0.6 in. Diameter Zirconium Alloy Tubes Under Irradiation," ASTM STP 551, 274 (1974). 4.1.5.a-17 B. Watkins et al., " Embrittlement of Zircaloy-2 Pressure Tubes," Aeolientions Related Phenomena for Zirconium and Its 411ovs, ASTM-STP-458, 1968.
4.1.5.a 18 M. A. Shubert, ' Examination of the PROTOTYPE and 1H038 Assemblies After Reactor Cycle 9 in Calvert Cliffs Unit 1 "
~
CENP50-493-P, January 1989. l 4.1.5.a 19 G. P. Smith, 'The Evaluation and Demonstration of Methods for Improved Nuclear Fuel Utilization; lith Progress Report,"
.. 00E/ET/34013-14, CEND 431, to be issued.
4.1.6.a 1 " Improvements to Fuel Evaluation Model,' CEN-161(B)-P Supplement 1-P, Combustion Engineering, Inc., April 1986. - 4.1.6.a-2 Letter from S. A. McNeil (NRC) to J. A. Tiernen (BG&E), ' Safety Evaluation of Topical Report CEN-161(B) P Supplement 1-P, Iinprovements to Fuel Evaloation Model," February 4,1987. 4.1.6.a 3 A. M. Garde, "Het Cell Exuiination of Extended Burnup Fuel from Fort Calhoun," DOE /ET/34030-11 CEND 427, Comhustion Engineering, September 1986. 4.1.6.a 4 U. P. Nayak at al, " Hot Cell Examination of Zion Fuel Cycle 5." WCAP 10543, Westinghouse, June 1985. 4.1.6.a-5 S. R. Pati, A. N. Garde and L. J. Clink, " Contribution of Pellet Rim Porosity to Low Temperature Fission Gas Release at Extended Burr.ups," Proc. ANS Tonical Meetino on LWR Fuel Performance, Williamsburg, VA, April 17-20, 1988, p. 204. 4.1.6.a-6 " Test Fuel Rod Irradiation in 14x14 Assemblies at Calvert
- Cliffs 1: Task A Research Project 586-1," CE NPSD-280, Combustion Engineering Topical Report.
4.1.6.a-7 'The Evaluation and Demonstration of Methods for Improved Fuel Utilization," DOE /ET/34010-11, CEN-415, November 1983. i 4.1.7.a 1 U. P. Nayak, et al, ' Hot Cell Examihttion of Zion Fuel Cycle 5,* WCAP-10543, Westinghouse, June 1985. 4.1.7.a-2 A. M. Garde, ' Hot Cell Examination of Extended Burnup Fuel from Fort Calhoun,' 00E/ET/34030 11, CEND 427, Combustion Engineering, September 1986. 4.1.8.a 1 B. J. Wrona, et al, ' Thermal Properties of Urania-Erbia,' Battelle Northwest Laboratories, dated June 1988. 4.1.8.a-2 J. Komatsu, et al, 'The Melting Temperature of Irradiated Fuel,' J. Nuclear Materials, No. 154 (1988), pp. 38-44. 4.1.9.a-1 A. M. Garde, ' Hot Cell Examination of Extended Burnup Fuel from Fort Calhoun,' 00E/ET/34030 11, CEND 427, combustion Engineering, September 1986. 4.149.a-2 U. P. Nayak, et al, ' Hot Cell Examination of Zion Fuel Cycle 5," WCAP-10543, Westinghouse, June 1985. 4.1.12.a 1 T. Hollowell, et al., 'The International Over Ramp Project at St'adsvik,' ANS Topical Meeting, LWR Extended Burnup-Fuel Performance and Utilization April 4 8, 1982, Williamsburg, Virginia. 4.1.12.a-2 5. Djurle et al., 'The Studsvik Super Ramp Project Final Report,' 1983, STSR-32. 4.1.12.a-3 J. C. LaVake and M. Gaertner, "High Burnup PWR Ramp Test Program-Final Report,' 00E/ET/34030-10, December 1984. b 4.1.12.a-4 R. Holzer and H. Stehle, 'Results and Analysis of KWU Power Ramp Investigations," KTG/ ENS /JRS Meetino on Ramping and load Following Behavior of Reactor Fuel, Petten, Netherlands, November 30 - December 1, 1978. l m . . . . . . . . .. . . . . .. ... , ,. . . . . .. . . . -.. _. .. ... . .. ., , . . . . . . . . . . . . . . . . . . ...
I i 4.1.14.a-1 D. E. Bassette et al., 'C E/EPRI Fuel Performance Evaluation I Program RP586-1 Task A: Examination of Calvert Cliffs ! Test Fuel . Assemblies at End of Cycles 1 and 2," CENPSD-72 Combustion { Engineering, Inc., September 1978. 4.1.14.a-2 E. J. Ruzauskas et al., 'C E/EPRI Fuel Performance Evaluation
, Program: RP586-1 Task A, Examination of Calvert Cliffs ! Test Fuel Assembly After Cycle 3 " CENPSD 87, Combustion Engineering,
,- Inc., September 1979.
.. j 4.1.14.a-3 E. J. Ruzauskas et al., 'C E/EPRI Fuel Performance Evaluation Program, RP586-1 Task A: Examination of Calvert Cliffs ! Test )
Fuel Assembly After Cycle 4," CE!4PSD-146, Cembustion Ergineering, ) Inc., October 1981. j 4.1.14.a-4 R. G. Weber et al., 'EPRI/C E Fuel Ferformance Evaluation ; Program, RP586-1 Task B: Examination of Arkansas Nucleat-One-Unit 2 Characterized Fuel Assemblies After Cy:le 1," CENPSD 174, Combustion Engineering, Inc., July 1982. ; 4.1.14.a-5 E. J. Ruzauskas et al., 'CE/EPRI Fuel Performance Evaluation Program, RP526-1 Task A: Examination of Calvert Cliffs-I Test Fuel Assembly After Cycle 5,* CENPSD-241, Combustion Engineering, Inc., July 1984. . 1 4.1.14.a-6 G. P. Smith, 'The Evaluation and Demonstration of Methods for Improved Fuel Utilization, End-of-Cycles 6 and 7 Fuel L Examinations," DOE /ET/34010-10, CEND 414, Combustion Engineering, October 1983.
.I 4.1.14.a-7 G. P. Smith, "The Hondestructive Examination of Fuel Assemblies ~~
with Standard and Advanced Design Rods After Three Cycles of Irradiation,' DOE /ET/34013-12, CEND-426. Combustion Engineering, I Inc., November 1986. L _
4.1.14.a-8 M. A. Shubert, " Examination of the PROTOTYPE and IH038 Assemblies After Reactor Cycle 9 in Calvert Cliffs Unit 1," CENPSD 493-P, January 1989. - 4.1.14.a 9 D. G. Franklin, "Zircaloy-4 Cladding Deformation During Power Reactor Irradiation," Fifth Symoosium on Zirconium in Nuclear Anolications, ASTM, August 4-7, 1980, Boston, MA. 4.2.1.a-1 Letter, A. E. Scherer (C-E) to C. O. Thomas (NRC), "CEA Guide Tube Wear Sleeve Modification," LD 84 043, August 3, 1984. ,, 4.2.1.a 2 "Palo Verde Nuclear Generating Station Unit 1 End of Cycle 1 + Fuel Examination Report," CE NPSD-428 P, Combustion Engineering, Inc., December. 1987. 4.2.1.a-3 "Palo Verde Nuclear Generating Station - Unit 2 End-of Cycle 1 : SLrveillanca Fuel Examination Report," CE NPSD-455-P, Combustion .i Engineering, Inc., May, 1983. : 4.2.2.a-1 " Application of CINPD-198 to Zircaloy Component Dimensional Changes", CEN-1G3(B), Combustion Engineering, Inc., September, 1981. - 4.2.7.a-1 System 80 D Standard Safety Analysis Report Final Safety Analysis Report (CESSAR FSAR), STN 50-470F, Combustion Engineering, Inc., j October 1978. 4.2.7.a-2 G. W. Keilholtz and R. E. Moore, " Irradiation Damage to Aluminum Oxide-Exposed to Sx10 21 Fast Neutrons /Cm 2
," Nuclear Aeolications, ;
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.e 83-
. - - .}}