ML20247E323
| ML20247E323 | |
| Person / Time | |
|---|---|
| Site: | Arkansas Nuclear  |
| Issue date: | 06/30/1989 |
| From: | ABB COMBUSTION ENGINEERING NUCLEAR FUEL (FORMERLY |
| To: | |
| Shared Package | |
| ML19292J329 | List: |
| References | |
| CEN-386-NP, NUDOCS 8907260133 | |
| Download: ML20247E323 (90) | |
Text
_ _ - _ _
CEN-386-NP VERIFICATION OF THE ACCEPTABILITY OF A 1-PIN BURNUP LIMIT OF 60 MWD /Ks FOR COMBUSTION ENGINEERING 16X16 PWR FUEL June 1989 Combustion Engineering, Inc.
Nuclear Power Systems 1000 Prospect Hill Road Windsor, Connecticut 06095 c 726Cgg[~fc 4
LEGAL, NOTICE THIS REPORT WAS PREPARED AS AN ACOOUNT OF WORK SPONSORED BY COMBUST:ON ENGINEERING, INC. NEITHER COM80STION ENGINEERING NOR ANY PERSON ACTING ON ITS BEHAI,F:
A.
MAKES ANY WARRANTY OR REPRESENTATION. EXPRESS OR llWLIED INCLUDING THE WARRANTIES OF FITNESS FOR A PARTICULAR PURPOSE OR MERCHANTABILITY, WITH RESPfCT TO THE ACCURACY, COMPLETENESS, OR USEFULNESS OF THE INFORMATION CONTAINED IN THIS REPORT, OR THAT THE USE OF ANY INFORMATION, APPARATUS, METHOD, OR PROCESS DISCLOSED IN THIS REPORT MAY NOT INFRINGE PRIVATELY OWNED RIGHTS:OR B. ASSUMES ANY LIA8813 TIES WITH RESPCCT TO THE USE OF, OR FOR DAMAGES RESULTING FROM THE USE OF, ANY INFORMATION. APPARATUS, METHOD OR PROCESS DISCLOSED IN THIS REPORT.
y ABSTRACT f
Several Utilities' using Combustion Engineering 16x16 fuel assembly designs have implemented programs to extend their fuel cycle lengths from 12 to 13 months and beyond.
The maximum 1-pin burnup oredicted for these extended burnup cycles exceeds the 52 MWD /kg limit presented in the existing C-E f'.
Extended Burnup Operation topical report.
For example, Arkansas Nuclear One Unit 2 Cycle 8 (the third 18-month cycle for that unit) will have a number of fuel pins that exceed this current burnup limit.
This report verifies the g '
adequate modelling of those 16x16 fuel design pins to 50 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 Combustion Engineering 16x16 fuel assembly designs employing
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TA8LE OF CONTENTS 19S119A EAER A8STRACT 11 l
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TABLE OF CONTENTS
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UE LIST OF TABLES iv 3l LIST OF FIGURES v
INTRODUCTION 1
DISCUSSION 2
3.3.6.a-Cl add i ng Co11 ap se.................................... 3 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 addi ng Co11 a p s e...................................
13 4.1.5.a1 Ductil i ty of Fuel t1 adding........................... 16 4.1.6.a Fission Gas Release.................................
28 4.1.7.a Fuel ' Thermal Conducti vi ty........................... 35 l
4.1.8.a Fuel Mel ting Temperature............................ 3 6 4.1.9.a
-Fuel Swe111ng.............,,.........................
37 i
4.1.10.a Fuel Rod Bow.........................................
38 l '
4.1.11.a ' Fretting. Wear.......................................
39 j
4.1.12.a Pellet / Cladding Interaction.........................
40 4.1.13.a Cl adding Deformation and Rupture.................... 42
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'4.1.14.a Fuel Rod Growth.................................... 43 4.2.1.a Guide Tube Wear.....................................
47 4.2.2.4 FueF Assembly Length Change and Shoulder Gap Change. 48 4.2.3.4 Fuel As sembly Ho1 ddown.............................. 62 4.2.4.a Grid Irradiation Growth.............................
63 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 Burnabl e Poi son Rod Behavior........................ 66 1
l CONCLUSION 75 1
i REFERENCES 76 1
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L' TABLES o;.. -
D f1SA 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 Low Power Test Rods 4.2.2.a-1 Analytical Models For [
J...............
53 4.2.7.a-1 Burnable Poison Rod Details...........................
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i, FIGURES fiEEN EAER 4.1.2.a-1 0xide vs.
Burnup.......................................
8 4.1.2.a-2 Cladding Peak 0xioe Thickness as a Function of........ 9 Average Burnup 4.1.2.a-3 Hydrogen Uptake as a Function of Oxide 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 Strencth at a Function of Fluence for..........
2Z
[
] Irradiation Temperature 500 to 650*F, Elevated Temperature Test Ultimate Tensile Strength of Shog n/cmTrangverse
..... 23 4.1.5.a-2 Specimens Irradiated to 4.3 x 10 (E.$1Mev) 4.1.5.a-3 Uniform Elongation as a Function of Flu c ce for......
24
[
] Zircaloy, Irradiation Temps. 560 - 610*F 4.1.5.a-4 Percent Reduction of Area for Shgt-Tragsverse 25 Specimens Irradiated to 4.3 x 10 n/cm (E>1Mev) 4.1.5.a-5 Effect of Hydrogen Concentration on the geductjon.... 26 2
of Area for Zirealoy-2 Irradiated to 10 n/cm 4.1.5.a-6 Fluence Dependence of Strain for Irradiated..........
27 Zirealoy-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 Guide Tube Length Change Data to.......
55 SIGREEP Predictions for ANO-2 Fuel Assemblies 4.2.2.a-3 Comparison of Guide Tube Length Change Data to.......
56 SIGREEP Predictions for SONGS Fuel Assemblies v
m.__m._____.___.--
FIGURES Fiaure' g
4.2.2.a-4 Comparison of Guide Tube Length Change Data to.......
57 SIGREEP Predictions for PVNGS Fuel Assemblies 4.2.2.a-5 Comparison of Guide Tube Length Change Data to.......
58 SIGREEP Predictions for St. Lucie 2 Fuel Assemblies 4.2.2.a-6 Comparison of Shoulder Gap Change Data to............
59 SIGREEP Predictions for ANO-R Fuel Assemblies 4.2.2.a-7 Comparison of Shoulder Gap Change Data to...........
60 SIGREEP Predictions for SONGS Fuel Assemblies 4.2.2.a-8 Comparison of Shoulder Gap Change Data to............ 61 SIGREEP Predictions for PVNGS Fuel Assemblies 4.2.7.a-I Swel l i ng o f Al 0 - B C................................. 74 23 4 e
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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 presentec' in Reference 1, 52 MWD /kg. For example, Arkansas Nuclear One Unit 2 Cycle 8 (the third 18-month cycle for that Unit) will have a number cf fuel pins that exceed this current burnup limit. This report justifies a 1-pin burnup limit of 60 MWD /kg for 16x16 fuel assembly designs by supplementing Reference I with data and discussions covering the additional burnup range 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 Je1 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 I 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.
The conclusions of this report regarding fuel assembly length change and shoulder gap change are applicable to combestion Engineering 16x16 fuel e.ssembly designs employing [
f 3-l _
DISCUS $10N The contents of the following update sections generally follow the format of their respective section in Chapter 3 or 4 of Reference 1.
Eacii (sub)section is numbered identically to its respective (sub)section in Reference I with the addition of ".a".
Each section has an introduction which specifies how the succeeding subsections should be treated, i.e., whether they append or replace 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.
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3.3.6.a Claddino co11anse This section replaces Section 3.3.6 of Reference 1.
Collapse is the term spplied 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 beginnir.g of life. The driving force for this creep is supplied by the differential pressure across the fuel or poison rod cladding.
C-E dcsign characteristics which mitigate cladding collapse are:
o Fuel and poison rods are prepressurized with helium which offsets the effects of external pressure tc 'the extent that cladding long term creep and cladding ovalization are reduced.
o "Non-densifying" or stable fuel pellets are used to prevent the formation of significant axial gaps within the fuel column.
This allows the fuel pellets to support the cladding later in life when the fuel-cladding gap closes.
o Poison rods behave in a similar fashion to fuel rods except the pellets
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are not sub, ject to densification.
.j The cladding collapse model is discussed in Section 4.1.4.a.
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4.1.1.a Faticue The discussion provided in Section 4.1.1 of Reference 1 applies to the proposed increase in the 1-pin burnup limit to 60 MWD /kg.
The method used i
to calculate fatigue damage -is applicable to extended burnup operation since the other sections of this report show that the individual components l
of the method (e.g.,
cladding creep and fuel swelling) are adequately modeled and the cladding has adequate ductility.
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'4.1.2.a Claddina Corrosion The. following subsections append the corresponding ' subsections of' Reference 1.
4.1.2.1.a r trosion Behavior a
M 0xide thickness data from three C-E PWRs, Calvert Cliffs-1, Ft. Calhoun, and - ANO-2, for red 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 maximum 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
.16x16 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 a-3 of Reference 1 in Figure 4.1.2.a-1.
The U enrichment level 235 for these high-burnep ro;s was between 3.03 and 4.00%. The U235 enrichment for future-fuel batches is expected to increase, but the burnups are not expected to exceed 60. MWO/kg.
T'he available oxide thickness data (,n irradiated fuel cladding approximately covers the naximum 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 16x*6 (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 + 30) oxide thickness of [
] microns at 60 L J/kg.
A similar fit to the Calvert Cliffs-1.14x14 Jata yields a best-estimate thickness of [ ] microns and an upper bound of [
]
microns.
Recently published high-burnup corrosion data from other PWRs (References 4.1.2.a-7 tr.- 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 enrichments lower than 4% U and several irradiation cycles of the order 235 of 12 months duration.
The heat rates of these fuel rods are generally lower than those expected for futu e high-burnop fuel rods.
l 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 micror.s.
This is in reasonable agreement with the upper bound estimate presented above for the ANO-2 data [
] and the Calvert Cliffs-1 data [
3 l
Another important aspect of cladding corrosion is the extent of hydrogen uptake by tha cladding. A fraction of the amount of hydrogen liberated by the Zircaloy corrosion reaction is absorbed by the cladding. As discussed
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in Section 4.1.5.a, the absorbed hydrogen may reduce the ductility of the cladding.
Hydfogen concentrations measured on cladding speciment 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 3
18% represents a reasonable upper limit on hydrogen absorption by cladding at high burnups.
This pickup fraction translates to a ciadding hydrogen level of about [
).
The relationship between hydrogen level and cladding ductility is further discussed in Section 4.1.5.2.a.
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4.1.2.2.a Evaluation of Claddinn Corrosion at Extended Burnuo
[
Based on the limitations discussed in Section 4.1.2.1.a, the 3a upper bound
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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 oxida layer thickness is insignificant with regard to
)
cladding stresses.
Although some tendency towards [ oxide spalling] was 1
reported at a level of oxide layer thickness of about [
] microns (References 4.1.2.a-4 and 4.1.2.a-7), fuel rod integrity was not impaired.
l It is, therefore, concluded that cladding corrorion is not likely to impair j
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 exampley the maximum local temperature incret.se at the metal-oxide interface due to a 3a upper bound oxide layer (on the 16x16 or I
14x14 design fuel rods), assuming a local fuel rod linear heat rate of
[
] kw/ft, is calculated to be about [
].
On a rod-average l L t
i basis, the temperature increase at the metal-oxide interface will be considerably less. The largest impact of the insulating oxide layer occurs at end-of-life when the linear heat rate of the fuel rod is significantly lower than the linear heat rata of the peak rod in the core.
Thus, it is j
concluded that the effect of oxide build-up on fuel temperature and stored energy is essentially counteracted by the lower linear heat rates that occur towards end-of-li fe.
- Thus, corrosion, based on observed 4
oxide-thicknesses at 60 MWD /kg in operating reactors, will not be limiting.
- However, additional factors in the future must be considered.
Specifically, the factors that need to be considered are EFPD (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.
Tnese factors will be monitored and corrosion evaluations performed as necessary, particularly if 1) the EFPD to achieve maximum burnup are considerably r.horter, 2) the reactor inlet temperatures t
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 impact 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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Figure 4.1.2.a-E Cladding Peak Oxide Thickness as a Function of Average Burnup 90
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Figure 4.1.2.a-3 Hydrogen Uptake as a Function of 0xide Thickness for Zircaloy-4 Cladding in PWRs
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'4.1.3.a C1addina Creen Append Subsection 4.1.3.3 of Reference I with the following material:
4.1.3.3.a fyaluation of Creen 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 it a contir.uous function to rod average burnups of 60 MWD /kg and that the model discussed in Reference 1 is adequate for 1-pin burnups of up to 60 MWD /kg.
The diametral strain data presented in Figure 4.1.3.a-1 shod that the fuel rod diameter does not change significantly during extended burnup operation.
Early-in-life, prior to the establishment of fuel-cladding contact, cladding creepdown occurred due to coolant pressure.
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4.1.4.a-Claddina Collaose This section replaces Section 4.1.4 of Reference 1.
l 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 geometiry. Under the high temperature and neutron flux conditions in the reactor, the Zircaloy cladding creeps in
~
response to the bending stresses. The resulting creep strain increases the deviation from the circuler 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 an unstable condition is reached and the cladding " collapses
- into a distorted shape.
Observations indicate that no significant axial gaps form in the fuel pellet column during the operation of Combustion Engineering's modern design fuel, which has prepressurized fuel rods and stable "nondensifying" fuel pellets.
Such gaps would be evidenced by unusual local ovalities of the fuel rod cladding, a distinct region of atypical crud depositian 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 operation the ax'al gaps in a fuel column are usually only a fraction of the length of a pellet.
The 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 rea;: tor startup reduces this cold gap to 5, 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. _
4.1.4.1.a-Modelina of Claddina Co11anse l
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 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 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 acceptable analytical procedure for determining the minimum time to collapse for C-E Zircaloy 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 acount for the support provided to the cladding by the pellets at the edges of the gap.
The adjustment varies as a function of the length of the gap or unsupported cladding.
As the gap considered becomes longer, the res~..a spproach the normal CEPAN results.
4.1.4.2.a Effect of Extended Burnun i
Since cladding collapse is a creep-related phenomenon, the longer residence f
times tssociated with extended-burnup fuel will increase the amount of creep of unsupported cladding.
The increased creep strain will be accounted for in the analysis of the ability of the fuel rod to resist l
cladding collapse.
A i - - _
4.1.4.3.a Evaluation of Claddino Co11anse b
Although early experience with densifying UO fuel pellets indicated that 2
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 i
significant axial gaps do not form in the fuel pellet column during operation.
Without the occurrence of gaps of sufficient length, cladding collapse cannot occur and, as a consequence, the cladding will remain stable and will not be subject to high local strains from this effect.
Furthennore, there is no evidence to indicate that continued operation of fuel rods having cladding in ovat contact with the fuel pellet column it detrimental.
C-E has performed cladding collapse calculations with the modified method 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 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 conc 1Med 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 6
unless design or manufacturing changes are introduced which would significantly reduce predicted collapse time results.
In the event such 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 l
lifetime of the fuel.
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4.1.5.a Ductility of Fuel Claddina m
This section replaces Section 4.1.5 of Reference 1.
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 (Zr0 ) layer. on the outer 3
2 surface of. the fuel rod.
Hydrogan is produced by this reaction and a l
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.f uel rod design criteria related to strength and auctility were discussed in Sections 3.3.2 and 3.3.3 of Reference 1, respectively.
Since the fuel rod design calculations are based on the yield strength of l
unirradiatr cladding, the increase in the yield streogth of cladding due to neutron irradiation does net pose a strength limitation on the cladding's performance.
The loss of ductility due to the ~ neutron irradiation and hydrogen uptake, however, ::seds to be. evaluated to assure that adequate cladding ductility exists at extended burnup levelt to ensure thbt i.he design strain limits remain valid. The effect of extended burnup operation on the cladding ductility is evaluated in this section.
i The elevatnd 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 md 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 deformation temperatures of about 600*F were considered in the analysis of the extended burnup data.
- The combined effect of the neutron fluence and hydrogen uptake on the
]
mechanical properties of Zircaloy-4 is evaluated below.
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I 4.1.5.1.a Mechanical Pronerties of Irradiated 71rcalov at Extended Butgyg C-E has obtained data on the mechanical properties. of Zircaloy-4 cladding irradiated in the Fort Calhm nactor to local burnups 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 burnug data presented in hElica 4.1.5 of Reference 1.
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C-E uses
[
]
fuel rod -cladding (Reference.4.1.5.a-6). The increase in elevated-temperature yield strength due to irradiation is illustrated in Figure 4.1.5.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.a-2).- 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.
.-l l
[
?
3 The fluence dependence of the [
] is illustrated in i
_ _ _ _ _ _ _ _ _ _. if ~ _
]
Figure 4.1.5.a-3 The data (Reference 4.1.5.a-12) suggest that for
[
l l
] 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 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 fluence (1.8 x.1021' n/cm ) nickel free Zircaloy-2. (Reference 4.1.5.a-15) 2 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 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 Procerties A fraction of the amount o 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 zirconium hydride platelets.
0 W.
It has been estabitshed that the ductility reduction due to hydrogen is dependent not only on the quantity of hydrides but also on their 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.
[
f'
)
Evans and Parry (Reference 4.1.5.a.11) determined the temperature above which the effects of unfavorably oriented hydrides disappear in cold-worked and stress-relief-annealed Zircaloy-2 cladding irradiated to low fluences.
At temperatures above 200*C (392*F), adversely oriented hydrides up to 200 ppm did not influence the ductility as measured by the reduction in area (Figure 4.1.5.a-4).
Hatkins et al. (Reference 4.1.5.a-17) have conducted I
tests on cold-worked tubular samples of Zircaloy-2 prehydrided to levels of up to 800 ppm which have circumferentially oriented hydrides.
Tensile
}
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 hydrogen showed values of the reduction in area at failure in excess of 60%.
- Thus, it has been concluced that at elevated temperatures, circumferential1y oriented hydrides up to 800 ppm do not influence the 20 2
ductility of Zircaloy cladding irradiated to fluence levels of 10 n/cm,
4.1.5.3.a Combined Effect of Radiation Damaae and Hydridino on the Luctility of Claddino at Extended Burnuos The ductility of extended burnup fuel rod cladding with rod average burnups approaching 60 MWD /kg (local burnups to 62.5 mwd /kg corresponding to 2
cladding fluence levels to 16.2 x 1021, n/cm, E > 0.021 MeV) has been recently measured by axial and ring tension tests and diametral burst tests l
(References 4.1.5.a-1 to 4.1.5.a-4).
The results of these mechanical tests l
l _-
y 1
i s
- 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 a
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) irradiated 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 claddine; data indicate that at a burnup of about 50 mwd /kg, the cladding can withstand an additional strain of 1% prior to plastic instability and about 4% strain prior to failure.
Axia'l tension tests on six-cycle fort Calhoun cladding (local burnups in the range of 57.6 to 63.3 MWD /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 HWD/kg show total circunferential 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 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
s
. and 5.6%.
At a local burnup level of 52.3-53.2 MWD /kg, the uniform strains 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 showr in Figure 4.1.5.a-6 (Reference 4.1.5.a-5).
For ' fluence values up to
-9x10 n/cm2 (E > 0.821 MeV or 8x1021 2
2I n/cm E > 1 FeV) (corresponding to burnups up to -53 mwd /kg), [
]
Moreover,. based on a detailed analysis of the microstructure of the fractured specimens, the fracture mode at btrnups 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 [
]a 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_______.
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TEST 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) - -
F 0
1 6
0 6
5 S
i P M
R E OT F
N E O C I NT E A U I 3 L D
- F A A
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Z F A
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55c4E3 Ia2E3c
-~.*
i
~
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1
-I se i
8 I
~
D E
E TEST TEMPER ATUR E. 'C 4
FIGURE 4.1.5.A-4 PERCENT REDUCTION OF AREA FOR SHORT-TRANSVERSE SPECIMENS IRRADIATED TO 4.3 x 1019 N/CH2 (E > 1 ME ) l
UNIARADIATED o
e
=
.0 300*C IRRADIATED a_,_
-D-w I 80 i
a h
^
E N
o E#
i N'h.NIRRADIATED U
g A
1
.3s IRmAotATEC A
w%
4 A
i
+
o 200 800 900 800 HYDMOGEN. PPM v
l FIGURF. 4.1.5.A-5 EFFECT OF HYOROGEN CONCENTRATION ON THE REDUCTION OF AREA FOR ZJ.i!CALOY-2 IRRADIATED TO 1020 2
ht/CH L.
I s
-s DEFORMATION TEMPERATURE 572 599 *F l
UNFORM TOTAL 0
W FUEL CLA00mG. BURST TESTS
- FUEL CLADOMG. TENSLE TEST go V
V GUIDE TUBE DOGBONE TENSLE TEST A
& GUCE TUBE RWG TENSEE TEST g
5 at
.g, TOTAL STRAM
=g a
\\
N:
i x
,N N
UWORM STRAW 8
2
~
0 b"""="========---
3 e
8 f
i I
I EC $
8 7
8 9
to 11 12 FLLENCE,1021,f,,2,E>0.821 MeV l
FIGURE 4.1.5.A-6 FLUENCE DEPENDENCE OF STRAIfi FOR IRRADIATED ZIRCALOY-4 l
1 _ _ _ _ _ _ _ - _ _ _ - _ _ _
p 1;
1 4.1.6.a Fission Gas Release The following section supplements section 4.1.6 of Reference 1.
4.1.6.1.a Fission Gas Relgggg The calculation of fission gas ralease is an integral part of the fuel performance 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 UO2 pellets to the cladding.
C-E's 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
~
(Reference 4.1.6.a-?)
Th6 FATES 38 fission gas release model. was developed utilizing data from low' and high power rods with burnups ranging from 6.5 to 61.5 MWD /kg and measured releases of 0.3 to 48.15.
The model includes the results of fission gas. release measurements performed on test rodlats that were irradiatsd in a PWR and subsequently ramped to high linear heat rates.
Comparisons between measurements and FATES 3B predictions are given in Referer.ca 4.1.6 a-l.
Additional extended burnup data on fission gas release has been obtained since the publication of References I and 4.1.6.a-1.
These data consist of six-cycle Fort Calhoun (49.7 to 55.7 MWD /kg) rods and five-cycle Zion-1 rods (54.3 to 59.4 MWD /kg) (References 4.1.6.a-3 and -4).
All of these fission gas release measurements were low (less than 2.8% at burnups up to 59.4 MWD /kg).
These data also show no significant enhancement of fission gas release with burnup at normal operating levels.
These data are-presented in Tables 4.1.6.a-1 and 4.1.6.a-2.
Microstructural examinations of the Fort Calhoun rods showed the formation of a porous rim (75 to 80% TD), 150-250 microns thick (References 4.1.6.a-3 and -5).
This porous cia 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.
[ l
.)
E-'
].
This increase is. not considered significant in ' low power, high j
burnup ' fuel.. In addition, other high burnup effects ' are known to offset-the temperature increase due to the porous rim. Two such important effects i
are ['
L.-
] Thus, it is concluded that the effects of a-porous rim can'be neglected for burnups of up to 60 WD/kg.
Rich Burnuo Data Comparisons i
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 WD/kg rod. average burnup) analyzed as part of the FATES 3B correTatton. 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 y
commercial reactors) have:been obtained since the publication of Reference 4.1.6.a-l. ' 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 3B fission gas release predictions in the high-burnup, low-temperature (low power) regime. These data are described below.
Calvert Cliffs Data:
High-burnup performance evaluations of Zircaloy-4 clad test fuel rods and "all Zircaloy" fuel-assemblies were performed on fuel irradiated in Calvert Cliffs 1.
The evaluations were sponsored by Combustion Engineering in conjunction with the Electric Power Research Institute (EPRI) (Reference 4.1. 6. a-6). -
A total of 60 test fuel rods were fabrii:ated fer this
' experiment and were equally distributed among three reconstitutable Batch B
~
assemblies.
Fission gas release data comparisons were perfomed'for 16 of these test rods, with and of life rod average burnups ranging from 18.7 to 44.4 WD/kgu, in support of the FATES 3B verification.
effort j.
(Reference 4.1.6.a-1).
Five of the modern design test rods, prepressurized.
s-m_.______m.__. - - - - - - - - - - -
- - = = - - - - -
p
]
L l-1' l
rods with radern design ncn-densifying pellets, were irradiated one 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 measured gas releases with FATES 3B predicted gas releases for these five test rods is presented in Table 4.1.6.a-3.
On the average, FATES 3B
[
]
~
Fort Calhoun Data:
Tne Fort Calhoun extended burnup demonstration program was sponsored by the 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 rome 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 were performed for four of the mo;t highly burned test rods (54.6 to 55.7 MWD /kg rod average burnup).
These four rods resided in positions very close to each other in the same quadrant 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-L 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 55.7 MWD /kg.
In general, the low temperi.ture release due to knock-out and recoil is [
] at 60 MWD /kg.
However, releases 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. 1 m______
4.1.6.2.a Evaluation of Fission Gas Releagg I'
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 conclusions are:
(1) Commercial fuel rods operating in PWRs with helium prepressurization and nondensifying fuel have been examined and consistently found to contain very low levels of released fission gases to burnup levels of
-60 MWD / kg.
The relative absence of significant enhancement dueNto burnup at normal operating levels is now verified by direct.
measurement.
~
(2) Data evaluated by C-E support the FATES 3B model to burnups of behaviors are 60 MWQ/kg.
Furthermore, the trends observed in all U02 gradital and support the orderly extension of the alloweble burnup.
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Table 4.1.6.a-3 L
FATES 3B Predictions of Gas Release from High Burnup, Low Power Test Rods Predicted-7.
Rod Average Measured Predicted Measured Burnup Gas Gas Gas and MWQZka Release Release %
Felease %
Calvert Cliffs-1 SN24 49.4 1.16 SN34 49.4 0.67 SN36 49.5 1.00 SN45 54.1
>2. d'2 SN59' 49.7 1.03 Fort Calhoun Extender Burnuo Fuel KJE051 55.7 1.26 KJE052 54.6 0.91-KJE077 55.4 1.33 KJE109 54.6 1.31 6
M a
'd 4 _ - _ _ _ _ _ _ _ - _ _ - - - -
h 1
l 4.1.7.a Fuel Themal Conductivity The following paragraph appends Subsection 4.1.7.3 of Reference 1.
s 4.1.7.3.a fyv,21uation of Fuel Thermal Conductivity No new data on the thermal conductivity of irradiated fuel has become available since the publication of Reference 1.
However, the performance 3
i of fuel rods to 60 MWD /kg (References 4.1.7.a-1 aad -2) indicates no trend tcward serious degradation of thermal conductivity.
The ability of the FATES 3B model to predict the measured gas release data suggests that any
]
degradation in local fuel thermal conductivity, such as due to, the j
formation of a porous rim, is implicitly accounted for in the FATES 3B I
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 reltvant submodels used in the fission gas release calculation.
It is therefore concluded that the current thermal conductivity equations are adequate to 60 MWD /kg.
).
4 _
__----_-____-_____________-______-_a
e 1
h 4.1.8.a Fuel Meltino Temperature o
The following paragraph appends Subsection 4.1.8.1 of Reference 1, 4.1.8..l.a Modelino of Fuel Meltino Temperature and Effect of Increased i
New data continue to support the conservatism of the melting point expression. The range of the melting point determinations of unirradiated U0 fabricated by C-E (5094-5173*F) performed at Facific Northwest Labs 1
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 U0 1rradiated 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.
4 - _ _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _
9 i
4.1.9.a Fuel Swellina j
l The' following paragraph appends Subsection 4.1.9.3 of Reference 1.
4 4.1.9.3.a Evaluation of Fuel Swelling i
Data for six-cycle fuel rods from Fort Calhoun and five-cycle fuel rods I
from Zion 1 (References 4.1.9.a-1 and
-2, respectively) have become tvallable since the publication of Reference 1.
Fuel density measurements were made on pellet sections with a local burnup of 60.4 WD/kg from Zion 1 and 53.3 WD/kg from Fort Calhoun.
These data and lower burnup data from previous cycles of these reactors indicate a
swelling rate, of 0.53%/10 WD/kg for Zion I and 0.70%/10 WD/kg for Fort Calhoun, which is i
entirely consistent with the 0.4-0.8%/10 WD/kg data measured previously for Fort Calhoun and Calvert Cliffs-1.
These results show no enhancement of the fuel swelling rate for-local fuel burneps up te 63.3 WD/kg, indicating no change in the fuel swelling mechanism up to this burnup level.
Consequently, the current FATaS?B model is valid to these high burnups.
l 4
)
'1 j
j O
r 4.1.10.a Fuel Rod Bow The discussion provided in Section 4.1.10 of Reference 1 applies to the proposed increase in the 1-pin burnup limit to 60 MWD /kg. Rod bow is not a concern for high burnup fuel rods since their power fallt.,ff more. than cortpensates for their rod bow penalty.
il 5
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3
p L
4.1.11.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 MWD /kg.
No significant fretting wear has been seen during extensive inspections of C-E fusi rods and the degree of stress relaxation of the grid springs and creepdown of the fuel rod changes very l'ittle after one' operating cycle.
e i
G I
a 4
i -
t 4.1.12.a Pellet /Cladditm Interaction 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..0ther data available in the literature have not been shown because of. design differences.
It is important to recognize that comparisons between experimental PCI results are only valid when the
. important design variables-are consistent, All of these redlets 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 w
is diminished. This fact [
.I
I h
b b
ix 9<y
'o O
ce WE E
5 3
s l
as M
W E
1#MX 'ONi1VW AY3H WV3 Nil n
n. - - - -
I 4.1.13.a Claddina Deformation & Runture l
l-The discussion provided in Section 4.1.13 of Reference 1 epplies to the proposed increase in the 1-pin burnup limit to 60 WD/kg.
It has been deterwined that the LOCA models for cladding deformation and rupture are adequate for use at 60 WD/kg. [
3 O
O 4
4 4
0 _ _ _ _ _ - _ _ _ _ - _ _ _ _ _ _ _ _ _.
4.1.14.a Feel Rod Growth The following replaces Subsection 4.1.14 of Reference 1.
p L
lt 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).
I 4.1.14.1.a Model ha of Fuel Rod Growth The overall elongation of a Zircaloy clad fuel rod is due to several contributing mechanisms including stress-free irradiattan growth of the Zircaloy cladding, mechanical interaction between the UO fuel pellets and I
2 the Zircaloy cladding, and a net positive growth component due to creepdown of the cladding under the external coolant pressure.
Eiich 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.
Growth.: train versus fluence (E>0.821 MeV) is linear on a log-log plot.
The functional form of such an equation is:
A (pt)"
e =
where
'e st min, percent.
=
neutron fluence, n/cm2 (E>0.821 MeV) x 10-21
- t
=
~'
A and n constants, as shown below.
=
l A regression analysis was used to determine the value of the consunts A and n and resulted in the following growth equations:,
i
__-]
p
]
Upper 95% tolerance limit: e =
)
Best estimate equation:
e =
Lower 95% tolerance limit: e =
The growth data used in this analysis covered a fittence range of
[
]
n i
4.1.14.2.a Effect of Extended Burnuo Rod length measurements performed on rods with fast fluences ~ up to
[
] htve shown continuous and well-behaved growth with increasing exposure (liefednces i
4.1.14.a-1 through 4.1.14.a-8).
These data have confirmed that no i
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.
[
] 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 programs at Fort Calhoun, Calvert Cliffs-1, and Arkansas Nuclear One-Unit 2 (Referent.es 4.1.14.a-3,
-5,-6,-7,-8).
4.1.14.3.a Enluation of Fuel Rod Greyth Figura 4.1.14.a-1 shows growth measurements obtained on C-E fuel rods compared to the C-E fuel rod granth model described in Subsection 4.1.14.1.a.
Data from 14x14 fuel rods at Calvert Cliffs-1 and Fort Calhoun have been ebtained for fluences of up to [
] (References 4.1.14.a-6, -8) while data from 16x16 fuel rods at Arkansas Nuclear One-
3 a,
m m
t.g.
m s
,. M
'UI, i qwo
. l';..
Ol bV p 4-l Unit 2 have beta obtained to fluences of [
1 (Reference M
4.1.14.a-7).
[
N Tha growth date from the t'alvert Cliffs-1 fuel rods have also been used in-en malysis of growth. published by Franklin which involved more than 700 feel rod length mersuremets (Reference 4.1.14.a 9).-
This analysis cotifinud tha well-behtved rdture of fuel rod growth at high fluence and
[
- ).
Die database. shon in Figure 4.1.14.a-1 includes measurements from ANO-2 fuel rods that showd higher growth than other rods in the same batch. The hfgher growths are belfrved to be related to the relatively high carbon content of the' cladding. A similar association betwatn the carbon content of cladding and fuel rod growth was also reported in the 1988 ANS Topical Meeting on LWR Fuel Performance by Fragema, describing performance of fuel rods irradicted in TNI
[
3
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A 5i 4 R $EEtm It3ouO gaai
a l
<4.2.1.a1 Guide Tebe W1E The discussion provided in section 4. 2.1 of Reference' I applies to the proposed ir. crease in the 1 pin burnup limit.to 60 MWD /kg.
An extensive program was initiat2d in response to the. detection of. guide tube wear.
This progrcm resulted in (a) the development of-a-guide tube wear sleeve.
design that essentially eliminates the concern of guide. tube wear (Reference 4.2.1.a-1),
and (b) the developinent of an unsleeved fuel assembly design that reduces the guide tube wear to acceptable levels.
The only unsleeved-designs being fabricated are the fuel assemblies. for the System 80 reactors.
Post-irradiation examinations of these assemblies (Reference-4.2.1.a-2 and Reference 4.2.1.a-3) have verified the c
conservatism of the analytical predictions used to justify the unsleeved assemblies.
As discussed in Reference 1, the defense of the un:leeved assembly design for extended bur.nup operation [
]
~
9 - - - _ _ _ - _ _ _ _ _ _ - - - _ _ _ _ _ _ _ _ _ _ _ _ _ _
l l
1 4.2.2.a Fuel Assemb'y Lenath Chance and Shoulder Gao Chance
.j This section replaces Section 4.7.2 of Reference 1 in its entirety, l
Fuel assembly length change results from two distinct mechat isms in the j
Zircaloy guide tubes:
irradiation induced growth and compressive creep.
j Growth is produced by radiation effects on the Zircaloy crystalline structure, and causes the guide tubes to elongate. Compressive creep is the permanent reduction in length of the guide tubes in response to net holddown force. on the fual assembly structure.
1 I
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 fuc.1 rods and the bottom of the upper end ff tting).
The length change is important in the evaluation of criteria pertaining to each of these aspects of fuel j
performance.
1 Since the holddown force is a function of fuel assembly length, irradiation induced guide tube growth causes en aaditional compression of the upper end fitting ' springs, *ncreasing 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
.~
o operation requires a time history analysis to properly account for the combined effects of irradiation growth and creep up to that esiat in time.
4.2.2.1.a Modelina of Assembiv Lenath Chance and Shouldet Gho Chanae l
a)
Assembiv Lenath Chanae Growth and creep characteristics are dependent on the metallurgical state of the Zircaloy guide tubes. The analytical models presented in Reference 1 for
[
] have
- 1.,
r P
been updated, based on all the available guide tube length change data'on C-E fuel assemblies with [
).
The general foms of the equations presented in. Reference 1 for the irradiation induced growth model and the axial creep model were maintained while the proportionality factors and exponential constants were adjusted to obtain a best fit of the data.
Uncertainties on the guide tube length change predictions were based on an evaluation of the errors between measured data and best estimate predictions.
The result was a fluence dependency of the length change uncertainty.
The updated irradiation induced growth model and axial creep model for [
] are summarized in Table 4.2.2.a-1, along with the uncertainty function on the guide tube length change.
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 and the probability / confidence factor of the guide tube length change model (see Item 4 of Table 4.2.2.a-1).
The SIGREEP computer code generates a set of randomly selected values for the w
input purameters 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 steos are repeated (starting with a different set of randomly selected values for the input parameters) until a sufficient number of castes (typically 2000) have been generated to define a probability histogram of length change at end of life (EOL).
The resultant histogram represents the statistical variation of EOL 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 a
actual data or appropriate design criteria.
Figure 4.2.2.a-1 presents a typical histogram of fuel assembly length change.
b)
Shoulder Gan Chance Shoulder gaps change with residence time in the reactor due to differential growth between the fuel rods and the fuel assembly st'ructure (guide tubes).
Reference 1 described a technique of evaluating shoulder gap change using the SIGREEP computer code.
With that technique, fuel assembly length change is calculated by SIGRETP 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 (BOL) 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 tte difference in fuel rod and fuel assembly length changes. A single value of shoulder gap change is the end product of the time history calculation.
The calculation is repeated with different sets of randomly selected values for the input parameters uritil a sufficient number of cases (again typically 2000) have been generated to define a probability histogram of shoulder gap at EOL.
This method of evaluating shoulder gap change is used on 14x14 fuel desions 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 grcwth rates of the ANO-2 Bat :h C rods and a Dstification for no ionger applying those high growth rates, to current fuel
)
designs (i.e., a change in the material specification of the cladding).
The i
l interim approach is, therefore, no longer necessary and the shoulder gap evaluation technique utilizing the S:'.EEP computer code with the updated 3
j fuel rod growth model of Section 4.1.14.a can be used for 15x16 fuel c'esigns, l
50
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 Extendod.lgragn As stated in the preceding sections, fuel assembly length change is the net
[
change resulting from irradiation induced growth and compressive craep of the guide tubes. Since growth is fluence dependent and compressive creep'is time and flux dependent, assembly length change and shoulder gap are affected by extended burnup.
In general, higher burnups are expected to result in l
~
greater increases in assembly length, greater holddown spring compression, and larger changes in shoulder gap.
The extent of these changes will be evaluated based on the specific extended burnup operating conditions and the particular fuel assembly design.
4.2.2.3.a Evaluation of Assemb1v Lenoth Ch3.n.q,e l
Guide tube length change data for C-E fuel assemblies with [
]
{
are shown in Figures 4.2.2.a-2 thru 4.2.2.a-5, along with SIGREEP predictions using the irradiation induced growth equation and axial creep equation from Table 4.2.2.a-1.
Figures 4.2.2.a-2 thru 4.2.2.a-5 present det a and predictions for fuel designs that have different axial loads on the fuel assembly.
The different axial loads result from differences in holddown spring forces and/or uplift forces on the fuel assembly spacer gritis.
These i
differences affect the axial creep component of the guide tube length change so the various sets of data must be presented on separate figures.
Inspection of the four figures shows that the best estirnate SIGREEP predictions are in good agreement with the data, both in the magnitude of the predictions and the trend of tiie predictions.
In addition, the upper and lower 95% predictions represent conservative estimates of the guide tube j
length changes. The good agreement between the data arm the SIGREEP I
predictions in all four figures
- confirmr, the creep model's correct 4 %m,--_ _ _ _ _ _ _ _ -. _. _. _ _ _ _ _
I sensitivity to the axial stress on the guide tubes. 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 predictir.g fuel assembly ~1ength change for designs with [
] to extended j
burnups.
4.2.2.4.a Ey.glqation of Shoulder Gao Chance
)
1 Shoulder gap change data for C-E fuel assemblies with l
] are 1
shown in Figures 4.2.2.a-6 thru 4.2.2.a-8 along with the limiting shoulder gap change prediction using the technique described in Section 4.2.2.1.a.
The SIGREEP predictions shown on the figures were generated using a typical ratio between the. fuel rod fluence and the guide tube fluence.
Three separate figures are provided because the fuel assembly designs associated with each figure have different guide tube length change characteristics (see above section)
- which, in
- turn, affect the shoulder gap change characteristics.
Inspection of the three figures shows that the analytical predictions represent conservative bounds of the data. Shoulder gap change data for fuel assemblies with [
] are available to fluences of approximately
[
] uvt. Fuel rod growth data exist to considerably higher fluences
~
(over [
] nyt, per Figure 4.1.4.a-1) and were included in the development of the revised fuel rod growth model.
Since the fuel rod growth is the predominant component in the shoulder gap change and the technique of predicting the limiting shoulder gap change employs the fuel rod growth equation that properly models the high fluence rod data, the fuel rod growth portion of the shoulder gap change analysis is acceptable for use to extended burnups. The guide tube growth portion of the shoulder gap change analysis uscs the model verified to extended burnups in the above section. Therefore, it is concluded that the analytical technique for predicting shoulder gap changes (SIGREEP) can be used to conservatively predict shoulder gap changes c
for designs with [
] to extended burnups.
1,
e Table 4.2.2.a-1 i-Analvtieal Mqdels_for I 1
1.
Overall Lenoth Chance Model l
Length Change - Irradiation Growth - Compressive Creep i Uncertainty l
2.
Irradiation __ Growth Model
- A (pt)"
Equation Form: (
where: c - axial growth strain, in/in-A = proportionality factor - [
]
1 pt - fluence, nyt x 10-21 (E > 0.821 KTV) n '- exponential constant = [
]
3.
Compressive Creen Model
- a p (a)N Equation Form: (
where:
c - axial creep strain rate, in/in/hr a - proportionality factor - [
]
g, (4)0.85,(-6000/RT) (Ake kt, g) 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 = [
]
C - constant - [
]
a - axial guide tube stress, ksi N = exponential constant = [
]
4.
Uncertainty Model Equation Form: Uncertainty - KS where: K - probability / confidence factor e.g.1.96 for 95/95
$ = standard deviation - [
]
pt - fluence, nyt x 10-21 (E > 0.821 MeV) _ - _ - _ _ - _ - _ -
n---
. -. - - -., _ _. = - -
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. 4.2.3.a Fuel Assembly Holddown
- The-discussion provided in Section 4.2.3 of Reference I applies to' the proposed increase in the 1-pin burnup limit to 60 NWD/kg.
The holddown spring relaxation due to extended burnup tends to be offset by concurrent growth of the fuel astsably.
4 m
49 9
e i
62-
-4.2.4.a Grid Irradiation Growth Th2 discussion provided in Section 4.2.4 of Reference 1 applies to. the proposed increase in the 1-pin burnup limit to 60 MWD /kg. Since the grid growth data presented in Reference I agreed well with all the other growth measurements [
] presented in that reference, the. good agreement between the growth measurements and predictions for [
] presented in Refarence I supports the adequacy of the grid irradiation growth model tc,i4 tended burnsp.
i, 9
i
)
i l
1 l
a le Y
' 4 I
L_-
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.
The 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 burnup demonstration 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 the extended burnup operation of 14x14 or 16x16 fuel assembly designs.-
S M
m 6
9 l
4.7.6.a - Corrosion of the Fuel _ Asnmb1v Structure The following paragraphs append Subsection 4.2.6.3 of Reference 1.
4.2.6.3.a Evaluation of Corrosion of the Fuel Assembiv Structure kiditional in-reactor corrosion data will be obtained from hot cell u aminations (metallographic and hydrogen content analyses) to be performed on a five-cycle Calvert Cliffs-1 assembly cage that experienced an assembly average burnto of (
).
Detailed poolside visual examinations were performed on this assembly.
7 indications of anomalous behavior, such as oxide spalling or structui, I
cracking, were observed.
ii.e het 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 i
increase monotonically with time. -
On reytew 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 y
lithium concentrations, further evaluations of the assembly structure corrosion behavior would have to be made.
1 e
1
Q i;
4.2.7.a Burnable Poison Rod Behavior The following subsections replace the. corresponding subsections. of Reference 1.
4.2.7.1.a Modelina of Burnable Poison Rod Behavior Al 0j-B C Pellet Swellina.
The swelling of the burnable poison material, g
4 induced - by irradiation, results in dimensional changes which can affect cladding strain and poison rod void volume. The neutron absorbor material employed in the poison rods is in a pelletized form and censists of a dispersion of boron carbide (B C) particles in an alumina (A10 ) matrix.
4 23 The 8 C content is. established. by core neutronic requirements and has 4
ranged to levels on the order of 4 wt%.
The dimensional changes of the pellet are predicted by a model which assumes [
I 1
Since the A1 0 swelling is the dominant contributor to pellet swelling at 23 high exposure, the A10 -B C swelling is related to fast fluence in the 23 4 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,a) Li-7 reaction.
In relating pellet swelling to irradiation exposure, it is assumed [
].
The B C swelling rate used is the same as in C-E's model 4
for 8 C swelling in a control element assembly (CEA) as described in 4
Reference 4.2.7.a-1, i.e., a volumetric swelling of 0.3% per percent B 10 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 A10 swelling is caused by fast neutron 23 irradiation damage, Keilholtz and Moore correlated their observed Al 0 23 j _ _ __ ___ _
r i
volume increases with fast fluence (E > 1 IleV).
J A review of the data reported by Keilholtz and Moore (Reference 4.2.7.s-2) indicates that for gross overall dimensional changes, a two-stage swelling rate model is an appropriate representation for A10 swelling.
That is,
{
23 21 nje,2, the swelling of above a fast fluence of approximately 2.6 x 10 A1 0 is enhanced by microcracking and grain boundary separation which 23 causes a sharp increase in the apparent overall swelling rate.
This enhancement of swelling was incorporated into the previous model which was described in iteference 1.
- However, since the volume created by microcracking accommodates the gas inventory in the rM, 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 A1 0 swelling is represented by the following expression 23 that accounts for the matrix swelling of A10 only:
23
.. ~
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 dep adency for A1 0 swelling in the range of 300 to 600*C.
23
[
]a two-stage model is used for the composite A10 -8 C pellet swelling model.
23 4 The volumetric swelling rate for B C (i.e., 30% at 100% B-10 depletion) wts 4
used in conjunction with Equation (1) for A10 to arrive at the following 23 expressions for the volumetric swelling of the composite A10 -B C pellet.
23 4 i
t i
l
)
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 8 C content of 3 wt%.
Also plotted
{
23 4 4
are volumetric swelling values calculated from diametral swelling data i
which were obtained in C-E sponsored post-irradiation examination programs i
to verify the performance of the A10 and Al 0 -B C pellets.
These data
-j 23 23 4 consist of direct diameter measurements on 42 whole A1 0 -8 C and 16 whole 23 4 A1 02 3 pellets which were from poison rods discharged after 1 cycle of exposure. The results of the post-irradiation examination of these 1-cycle A1 0 -8 C pellets substantiated the assumption of isotropic swelling 23 4 behavicr (i.e., equal axial and diametral swelling rates).
It was also found that swelling was independent of initial pallet density in the j
L density range of 85 to 98% TD.
In addition, indirect diametral swelling
)
data wera obtained, at higher exposures, by profilometry measurements on I
I 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 diametral swelling in these rods was inferred by conserv'atively i
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
1 cladding strain as affected by pellet swelling and. (2).by inference, the restrained -swelling behavior of the. A10 -8 C pellets.
It was found that-23 4 even after: 4 cycles of reactor. operation, the average cladding strain was still negative',-exhibiting only a slight tendency to be less negative'than j
' 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 shown by 'the profile traces.
The inferred A10 -B C pellet swelling in-23 4 these rods, shown in Figure 4.2.7.a-i, was calculated from the. Irradiated diameter profiles, the as-fabricated cladding. wall thickness, and the 7
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 dataL derived from rod profiles represent a restrained coadition.
A comparison of the performance data with the model in figure 4.2.7.a-'1 Indicated the following:
o The swelling of A10 -8 C pellets, as 'well as that of A102 3 pellets, 33 4 that occurred during the firtt-cycle of irradiation up to a. fluence of about 3.5 x 1021,fe,2 (E > 1 MeV) are reasonably predicted 'by Equations (2) and (3).
The data scatter indicated that several 1-cycle A10 -B C pellets apparently swelled more than predicted by 23 4 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 21 2
irradiation up to
- 4. cycles, equivalent to 8.2 x 10 n/cm (E > 1 MeV).
The reason for the lack of apparent diametral swelling is related to the following overall swelling behavior mechanisms:
c (a) 8 C particle swelling caused by the B-10 (n,a) Li-7 reaction 4
induces microcracking and grain boundary separation in the pellet structure. -
[~1 ;k 1
(b) The. resulting early apparent swelling (while the B-10. is depleting) is enhanced ty this void contribution when.the pellet-is not restrcined (This may account for any underprediction of L j
1-cycleswelling).
.(c) At, higher fluence (i.e., after 2005 B-10 depletion) some of these new voids are - tecommodating the A10 satrix swelling due to 23 cladding restraint.
Once~ the accommodation is completed, diametral swelling, and therefore, volumatric swelling, would proceed at the swelling rate indicated by Equation (3).,
The subsections of Gas Release, Poison Rod Growth, and Poison Rod Claddino Creep of Reference 1 apply to the proposed increase in the 1-pin burnup limit to 60 MWD /kg.
Poison Red 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 represants either a time-dependent, fluence-dependent, or power-history dependent mechanism which will produce changes in the poison rod internal pressure through changes in the void volume and th' amount of helium released.
o e
Calculation of the EOL internal pressure is predicted for appropriate E0L conditions which include the number of moles of helium (precressure plus gas released from the pellets), gas temperature (the 100% depleted poison pellets produce only a small amount of heat flux due to gamma heating), and 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 B01. is nonexistent (Table 4.2.7.a-1),
1
4.2.7.2.a Effect cf Extended Burnun on Burnable Poison Rod Behavior j
A1 0;-B C ' Pellet Sve111na.
The swelling of. A10 -8 C pellets' is strongly 4
23 4 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 J..
void volume will be decreased by the diametral and #Nial 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 for up to 4 cycles.
. Axial Growth and Diametral Cr.g.tg.
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.
~ 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 during this period of time.
Full diametral contact between the pellets and cladding is not predicted so outward creep of the cladding due to swelling of the pellets is not expected.
Rod Intern,a1 Pressure.
Internal pressure will increase during extended burnup operation due to e reduced void volume within the rod caused principally by pellet swelling. Rod growth and creepdown are second order effects on the void volume when compared to pellet swelling, but are accourited for.
No additioral 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.
w 9
l l
l 1 '
]
Table 4.2.7.a-1 Burnable Poison Rod Details Extended Extended Early Burnup Early Burnup Parameter 14x14 Desi.ED 14x14 Desian 16x16 Desian 16x16 Desian Pellet 0.0.,
0.376-0.379 0.362 0.310 0.307 in.
Cladding 0.D.,
0.440 0.440 0.382 0.382 in.
Cladding I.D.,
0.388 0.384 0.332 0.332 In.
~
- Expressed as a percent of the total pellet volume.
1 i
1
)
j 1 -
-]
0 8
9 3
2 1
C4 8
M 3
e 3
Oy lA S
T N
i ES ER P
ER C4 B
S S
T 1
s a.
3 T
- 7. O N
N l
2 E
O
- 2. A l
R M P
4 E
E F E
A T R T
RO EU N
A UG MS E
I A
D A
GN A
0 D I E
M 2
I F L OM N
L E
,f W
S A
S R
R O
O R
T T
O
)S C
C T
T l
l'
%0 A
A C
E E
E A
L E
R R
E L
R E
L a
P C
M Y
B W n20. 9 O S 4 9
C0 C
4/
E F
2 R
FI O
9 L
ts 1
9 F
D O G
ET R V
s A P A
.o 3 l 0
nMt MD 3
e i
N I
D e
O O I
S S T R 2
mHH O
S 0
l R
AA Ea A o0
, O 9
5 0
s.
0 5
9 7
G 4
3 1
i s eA.* 3 9 a o>
c c
i2'
CONCLUSION The objective of this report is to justify the validity of C-E methods and models concerning the 16x16 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 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 cycles, 60 MWD /kg.
As such, the overall and individual conclusions presented in Reference 1 are shown to be valid for the extension of the 1-pin burnup limit to 60 MWD /kg for 16x16 fuel assembly designs.
The conclusions of this report regarding fuel assembly length change and
~
shoulder gap change are applicable to Combustion Engineering 16x16 fuel assembly designs employing [
).
Also, since the various fuel performance topics discussed in Reference 1 have no explicit dependence on batch average burnep, the batch average discharge limit of Reference 1 is no longer required and can be deleted.
O L.
9 __._
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! ~
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g
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