ML20236A095

Verification of Acceptability of 1-Pin Burnup Limit of 60 Mwd/Kg for Calvert Cliffs Units 1 & 2

ML20236A095
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Site:	Calvert Cliffs
Issue date:	01/31/1989
From:	ABB COMBUSTION ENGINEERING NUCLEAR FUEL (FORMERLY
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CEN-382(B)-NP, NUDOCS 8903160369
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Text

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i CEN-382(B)-NP 4

' I VERIFICATION OF THE ACCEPTABILITY OF a

A 1-PIN BURNUP LIMIT OF 60 MWD /KG FOR CALVERT CLIFFS UNITS 1 AND 2 l

January 1989 i

l Combustion Engineering, Inc.

I Nuclear Power Systems 1000 Prospect Hill Road Windsor, Connecticut 06095 l

l DO IO$b! $$0b 7

l LEGAL NOTICE THIS REPORT WAS PREPARED AS AN ACCOUNT OF WORK SPONSORED BY COMBUSTION ENGINEERING, INC. NEITHER COMSUSTION ENGINEERING NOR ANY PERSON ACTING ON ITS BEHALF:

A. MAKES ANY WARRANTY OR REPRESENTATION, EXPRESS OR IMPLIED INCLUDING THE WARRANTIES OF FITNESS FOR A PARTICULAR PURPOSE OR MERCHANTABILITY, WITH RESPECT TO THE ACCURACY, COMPLETENESS, OR USEFULNESS OF THE INFORMATION CONTAINED IN THl8 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 LIABILITIES WITH RESPECT TO THE USE OF, OR FOR DAMAGES RESULTING FROM THE USE OF, ANY INFORMATION, APPARATUS, METHOO OR PROCESS DISCLOSED IN THIS REPORT.

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CEN-382(B)-NP VERIFICATION OF THE ACCEPTABILITY OF A 1-PIN BURNUP LIMIT OF 60 eld /Ks FOR CALVERT CLIFFS UNITS 1 AND 2 January 1989 Combustion Engineering, Inc.

Nuclear Power Systems 1000 Prospect Hill Road Windsor, Connecticut 06095 l

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l ABSTRACT Baltimore Gas and Electric has implemented a program to extend the cycle length at its Calvert Cliffs Units from 18 to 24 months.

The maximum 1-pin burnup predicted for.these 24-month cycles exceeds the 52 MWD /kg limit presented in the existing C-E Extended Burnup Operation topical report.

l Calvert Cliffs 2 Cycle 9 (the second 24-month cycle for that unit) will have a number of fuel pins that exceed this current burnup limit.

This report verifies the adequate modelling of those pins to 60 MWD /kg (the new limit required by the implementation of 24-month cycles) by supplementing the existing topical report with additional data and discussions.

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1 TABLE OF CONTENTS l

l Section Eggg ABSTRACT ii TABLE OF CONTENTS iii LIST OF TABLES iv LIST OF FIGURES v

INTRODUCTION 1

DISCUSSION 2

3.3.6.a Cl add i ng Co11 ap s e.................................... 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 add i ng Co11 ap s e...................................

13 4.1.5.a Ductility of Fuel C1 adding..........................

16 4.1.6.a Fission Gas Release.................................

28 4.1.7.a Fuel Thermal Conductivi ty........................... 35 4.1.8.a Fuel Melting Temperature............................

36 4.1.9.a F u el Swe 11 1 n g....................................... 3 7 4.1.10.a F u e l Rod B ow........................................ 3 8 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 F u el Rod G rowt h..................................... 4 3 4.2.1.a Guide Tube Wear.....................................

46 4.2.2.a Fuel Assemoly Length Change and Shoulder Gap Change.

47 4.2.3.a Fuel As sembly Ho1ddown.............................

50 4.2.4.a Grid Irradiation Growth.............................

51 4.2.5.a Spacer Grid Rel axati on..............................

52 4.2.6.a Corrosion of the Fuel Assembly Structure............

53 4.2.7.a Burnabl e Poi son Rod Behavior........................

54 CONCLUSION 63 i

REFERENCES 64 l

iii

l TABLES lid 211 P.Agit 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 4.2.7.a-1 Burnabl e Poi son Rod Detail s........................... 61 W

Y iv

FIGURES i

I Fiaure P_agg l

4.1.2.a-1 0xide vs.

Burnup.......................................

8 4.1.2.a-2 Cladding Peak 0xide Thickness as a Function.of........ 9 Average Burnup j

4.1.2.a-3 Hydrogen Uptake as a Function of 0xide Thickness.....

10 for Zircaloy-4 Cladding in PWRs v

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

[

] Irradiation Temperature 500 to G50*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-Tragsverse 25 Specimens Irradiated to 4.3 x 10 n/cm (E>1Mev) 4.1.5.a-5 Effect of Hydrogen Concentration on the $ educt {on....

26 2

of Area for Zircaloy-2 Irradiated to 10 n/cm 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 i

4.1.14.a-1 Fuel Rod Growth Measurements Compared to C-E.........

45 j

Zircaloy Fuel Rod Growth Model l

4.2.2.a-1 Comparison of Calvert Cliffs Reload Design...........

48 Assembly Length Changes to SIGREEP Predictions 4.2.2.a-2 Comparison of 14x14 Fuel Assembly Shoulder Gap.......

49 I

Changes to SIGREEP Predictions 4.2.7.a-1 Swel l i ng o f A1 0 - 8 C.................................

62 23 4 l

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l l

v

i INTRODUCTION Baltimore Gas and Electric has implemented a program to extend the cycle l

length at its Calvert Cliffs Units from 18 to 24 months.

Cycle 9 is the second 24-month cycle for Unit 2.

The predicted maximum 1-pin burnup of Unit 2 Cycle 9 and subsequent 24-month cycles exceeds the current C-E 1-pin burnup i

~

limit presented in Reference 1, 52 MWD /kg.

This report justifies a 1-pin burnup limit of 60 MWD /kg by supplementing Reference 1 with data and discussions covering the additional burnup range required by the implementation of 24-month cycles, 52 MWD /kg to 60 MWD /kg.

L 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.

i

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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 is numbered identically to its respective (sub)section in Reference 1 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 1 o

which is a general reference that applies to all sections of this report.

i o

4 1

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3.3.6.a Claddina Collaose j

l This section replaces Section 3.3.6 of Reference 1.

i Collapse is the term applied to a condition where a slightly oval cladding i

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

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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 are reduced.

l 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 are not subject to densification.

The cladding collapse model is discussed in Section 4.1.4.a.

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r 4.1.1.a Fatiaue 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 to calculate fatigue damage is applicable to extended burnup operation since the other sections of this report show that the individual components

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of the method (e.g.,

cladding creep and fuel swelling) are adequately l

modeled and the cladding has adequate ductility.

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4 4.1.2.a Claddina Corrosion i

The following subsections append the corresponding subsections of Reference 1.

4.1.2.1.a Corrosion Behavior

[

]

The U enrichment 235 for the future fuel batches of Calvert Cliffs-1 and -2 is expected to increase, but the burnups are not expected to exceed 60 MWD /kg.

Moreover, the total effective full power days for three of the 24-month cycles planned for Calvert Cliffs-1 and -2 is expected to be close to the EFPD experienced by [

]

The lithium-boron coordinated coolant chemistry employed in Calvert Cliffs-1 during the irradiation of SCOUT and PROTOTYPE fuel rods included a maximum lithium level of 2.2 ppm at the beginning of cycle (80C).

Recently, the BOC lithium hydroxide level at Calvert Cliffs was increased I

to 3.5 ppm to reduce the excore radiation levels.

However, for low-coolant-temperature plants, such as Calvert Cliffs-1 and

-2, where 1

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t boiling is minimal, the increase of the B0C lithium level from 2.2 to 3.5 ppm is not expected to have any significant effect on the Zircaloy-4 cladding corrosion rate (References 4.1.2.a-5, 4.1.2.a-6).

Control of the coolant pH and hydrogen level is expected to minimize the formation of crud that could accelerate the corrosion' rate of the cladding.

Recently published high-burnup corrosion data from other PWRs (References 4.1.2.a-7 to 4.1.2.a-10) are presented together with the da'ta of Figure 4-4 of Reference 1 in Figure 4.1.2.a-2.

For a rod average burnup of - 60 MWD /kg, the upper limit of expected oxide thickness from Figure 4.1.2.a-2 o

is about 100 microns.

[

]

Another important aspect of cladding corrosion is the extent of hydrogen 1

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.

Recent hydrogen concentration data from. several PWRs 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.

[

] 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 Claddina Corrosion at Extended Burnuo I

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The cladding wastage due to this level of oxide layer thickness is l.

insignificant with regards to cladding stresses.

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It is, therefore, concluded that cladding corrosion will not impair the integrity of fuel rods irradiated in Calvert Cliffs-1 and -2 to rod average burnups of 60 MWD /kg. !

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I

An oxide layer will, of course, increase the surface temperature of. the cladding.

[

]

On a rod-average basis, the temperature increase 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 rate of the peak rod in the core.

Thus, it is 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 at-end-of-life, i

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Figure 4.1.2.a-2 Cladding Peak 0xide Thickness as a Function of Average Burnup 80 6

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5 10 20 50 100 OXIDE THICKNESS, pm Figure 4.1.2.a-3 Hydrogen Uptake as a Function of 0xide Thickness for Zircaloy-4 Cladding in PWRs f

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I 4.1.3.a Claddina Creen l

Append Subsection 4.1.3.3 of Reference I with the following material:

1 4.1.3.3.a Evaluation of Creen

(

] 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 [

r

] 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 Referer.ce 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 show 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 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 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 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 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 operation the axial 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 reactor startup reduces this cold gap to 10.3 inches.

Thus, the largest hot gap inferred from all post irradiation examinations of modern C-E fuel was 0.3 inches.

This conclusion is q

supported by the corrosion patterns observed during visual examinations.

I i - - - - - _ - - _ _

l 4.1.4.1.a Modelina of Claddina Collaose 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 i

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 roc 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 l-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 account 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 results approach the normal CEPAN results.

4.1.4.2.a Effect of Extended Burnuo Since cladding collapse is a creep-related phenomenon, the longer residence times associated 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 cladding collapse. - _ _ _ _ _ _

4.1.4.3.a Evaluation of Claddina Collaose l

Although early experience with densifying U0 fuel pellets indicated that l

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 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.

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 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 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 addres; cladding collapse for new cores or reload batches unless design or manufacturing changes are introduced which would significantly reduce predic'ed 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 lifetime of the fuel. _....

i l

4.1.5.a Ductility of Fuel Cladding This section replaces Section 4.1.5 of Reference 1.

l 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 2

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 cladding.

This hydrogen uptake may also reduce the ductility of the cladding.

The fuel rod design criteria related to strength and ductility 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 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.

[

].

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 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.

1 4.1.5.1.a Mechanical Properties of Irradiated Zircalov at Extended Burnuol C-E has obtained data on the mechanical properties of Zircaloy-4 cladding I

1 l

i irradiated in the Fort Calhoun reactor to local burnups of up to 62 mwd /kg (Reference 4.1.5.a-1).

In addition, mechanical property data have also 4

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.

I C-E uses

[

]

fuel rod claading f

a (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 I

through4.1.5.a-10). The increase in yield strength is also observed after extended burnup (Reference 4.1.5.a-5).

The increase in the yield 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) (Figure 4.1.5.a-2) when the hydrogen level 1

is increased from 0 to 200 ppm.

The yield strength behaves in a similar manner.

[

]

The fluence dependence of the [

]

is illustrated in Figure 4.1.5.a-3.

The data (Reference 4.1.5.a-12) suggest that for

[

l

] These tests were conducted at high strain rates. i

l J

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 (Reference 1

4.1.5.a-15) indicate that at high stresses, the creep rupture strains span the range of 6-18% strain. Therefore, at the lower stresses appropriate to cladding creep, the creep strains at rupture are expected to. be greater than 6% (Reference 4.1.5.a-13).

4.1.5.2.a Influence of Hydroaen on Mechanical Properties 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 l

zirconium hydride platelets.

It has been established that the ductility reduction due ' to. hydrogen is I

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 anplied stress, the reduction in ductility is relatively small.

[

I

]

Evans and Parry (Reference 4.1.5.a.11) determined the temperature above j

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 -

l l

l 1

l l

l (Figure 4.1.5.a-4).

Watkins et al. (Reference 4.1.5.a-17) have conducted tests on cold-worked tubular samples of Zircaloy-2 prehydrided to levels of I

l up to 800 ppm which have circumferential1y oriented hydrides.

Tensile tests showed that hydrogen concentration had only a minor effect on l

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 i

60%.

Thus, it has been concluded 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 Rydridina on the Ductility of Claddina at Extended Burnung j

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 21 2

cladding fluence levels to 16.2 x 10, n/cm, E > 0.821 MeV) has been 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 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 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 on 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% prior 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 MWD /kg) (Reference 4.1.5.a-1) show that for a i

1 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 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 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 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

-9x10 n/cm2 (E > 0.821 MeV or 8x1021.n/cm2 E > 1 MeV) (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 burnups greater than 53 mwd /kg was 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 [

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TEST TEMPER ATUR E. 'C FIGURE 4.1.5.A-2 ULTIMATE TENSILE STRENGTH OF SHORT-TRANSVERSE SPECIMENS IRRADIATED TO 4.3 X 1019 N/CM2 (E > 1 MEV),

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FIGURE 4.1.5.A-4 PERCENT REDUCTION OF AREA FOR SHORT-TRANSVERSE 19

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SPECIMENS IRRADIATED TO 4.3 x 10 N/CM2 (E > 1 MEV),

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FIGURE 4.1.5.A-5 EFFECT OF HYDROGEN CONCENTRATION ON THE REDUCTION OF AREA FOR ZIRCALOY-2 IRRADIATED TO 1020 2

N/CM _ _ _ _ _ _ - _ _ - - _ _ _ _.

12 DEFORMATION TEMPERATURE 572 599 *F UDGFORM TOTAL 0

8 FUEL CLADDMG, BURST TESTS X

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FIGURE 4.1.5.A-6 FLUENCE DEPENDENCE OF STRAIN FOR IRRADIATED ZIRCALOY-4.

1 l

l I

4.1.6.a Fission Gas Release j

The following section supplements Section 4.1.6 of Reference 1.

l 4.1.6.1.a Fission Gas Release The calculation of fission gas release 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, 1

affects the transfer of heat from the U02 pellets to the cladding.

C-E's I

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-2)

The FATES 3B 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 48.1%.

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 I

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 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 rim can result in a decrease in local fuel thermal conductivity and thus an increase in pellet temperature.

C-E believes that l

this porous layer is a phenomenon associated with local burnup and is well behaved.

[ /

1 l

l l

l l

1 l

]

This increase is not considered significar.t in low power, high burnup fuel.

In addition, other high burnup effects are known to offset 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 Burnuo Data Comparisons The predictive capability of the FATES 3B 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-1.

The high-burnup data sets (at and above 50 MWD /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) has been obtained since the publication of Reference 4.1.6.a-1.

Comparisons of these data to FATES 3B predictions are presented 4

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 fabricated for this

~

experiment and were equally distributed among three reconstitutable Batch 8 l

l assemblies.

Fission gas release data comparisons were performed for 16 of

]

these test reds, with end of life rod average burnups ranging from 18.7 to I

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, prepressurized rods with modern design non-densi fying 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:

I The Fort Calhoun extended burnup demonstration program was sponsored by the Department of Energy (D0E) 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 l

I Group, and C-E (Reference 4.1.6.a-3).

Fission gas release data comparisons were performed for four of the most 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 I

design input parameters and an irradiation history that appropriately nodels 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 55.7 MWD /kg.

In general, the low temperature 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 fiss'on gas release of extended-burnup fuel operated under normal conditions in C-E designed commercial reactors.

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 extend burnups.

Significant strides have been achieved in the area of normal operation and in the area of response to ramps.

The conclusions are:

4 (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 due to burnup at normal operating levels is now verified by direct measurement.

(2) Fuel rods which were irradiated in a PWR and subsequently ramped to high linear heat rates (LHRs up to 20 kw/ft) show higher releases of fission gas. The amount of fission gas released is strongly dependent on lir. ear heat rating (temperature) and the grain size of the U0 2 pellets.

These data, when plotted as a function of LHR, display an apparent enhancement of fission gas release due to burnup for burnups up to approximately 25 MWD /kg.

As the burnup of these test rodlets increases, the data show little additional burnup enhancement which is probably due to an improved gap conductance resulting from better fuel-clad contact at higher burnups.

(3) Data evaluated by C-E support the FATES 3B model to burnups of 60 MWD /kg.

Furthermore, the trends observed in all UO behaviors are 2

gradual and support the orderly extension of the allowable burnup.

(4) Design improvements including helium prepressurization, nondensifying UO, reduced pellet-cladding gaps and the use of pellets with larger 2

grain sizes have all shown improved behavior relative to fission gas release. \\

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L i

i 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 End MWD /ka Release Release %

Release %

C.alvert 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 KJE051 55.7 1.26 KJE052 54.6 0.91 KJE077 55.4 1.33 KJE109 54.6 1.31 e

4

_m_

1 1

1 l

4.1.7.a Fuel Thermal Conductivity The following paragraph appends Subsection 4.1.7.3 of Reference 1.

4.1.7.3.a Evaluation 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 l

of fuel rods to 60 MWD /kg (References 4.1.7.a-1 and -2) indicates no trend toward serious degradation of thermal conductivity.

The ability of the l

FATES 3B model to predict the measured gas release data suggests that any degradation in local fuel thermal conductivity, such as due to the 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 is therefore concluded that the current thermal conductivity equations are adequate to 60 MWD /kg.

l l,

______m________

l 4.1.8.a Fuel Meltina Temperature The following paragraph appends Subsection 4.1.8.1 of Reference 1.

j 4.1.8.1..a Modelina of Fuel Meltina Temperature and Effect of Increased l

Burnuo 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 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.a-2) showed no effect of burnup on U0 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.

O !

.q l

4.1.9.a fuel Swellina i

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 rate of 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.

6.

I i

4.1.10.a Fuel Rod Bow I

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 fall off more than compensates for their rod bow penalty.

e l

l.

l

. t j

4.1.ll.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 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.

I

'h, i

4 0 j

-)

l' c

4.l~.12:a ' Pellet /Claddina Interaction S

The following section replaces Section 4.1.12 of Reference 1.

I 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 4 odlets 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 to 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 PCI performance.

[

]

In addition, as burnup increases, the capability of the fuel to reach the power levels needed for PCI failure is diminished. This fact [

]

O +

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4.1.13.a Ciaddina Deformation & 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 determined that the LOCA models for cladding deformation and rupture are adequate for use at 60 MWD /kg.

[-

]

1 l I

c.

J

i 4.1.14.a Fuel Rod Growth The following paragraph replaces the first paragraph of Subsection 4.1.14 of Reference 1.

It has been well established that Zircaloy-4 clad rods exhibit axial elongation or growth when continuously exposured 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 verify existing fuel rod growth models originally developed with data obtained at lower fluences and from rods of older design (densifying fuel with lower initial pressurizationlevels).

The following subsections repl ace the corresponding subsections of Reference 1.

4.1.14.2.a E?fect of Extended Burnuo Rod length measurements performed on rods with fast fluences up to [

] have shown continuous and well-behaved growth with increasing exposure (References 4.1.14.a-1 through 4.1.14.a-8).

These data have confirmed that no acceleration of the growth rate or other abrupt chu,es 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.

[

] do 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 (References 4.1.14.a-3, -5, -6, -7, -8). I s

4.1.14.3.a Evaluation of Fuel Rod Growth Figure 4.1.14.a-1 shows growth measurements obtained on C-E fuel reds compared to the C-E fuel rod growth model described in Reference 4.1.14.a-9.

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-Unit 2 have been obtained for fluences of up to [

]

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-10).

This analysis confirmed the well-behaved nature of fuel rod growth at high fluence and

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l

i 4.2.1.a Guide Tube Wear l

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 eliminates the concern of guide tube wear (Reference 4.2.1.a-1).

The current 14 x 14 fuel assembly design incorporates this guide tube sleeve design.

e

4.2.2.a fuel Assembly Lenath Chanae and Shoulder Gao Chance l

l l

The following subsections append the corresponding subsections of Reference 1.

4.2.2.3.a Evaluation of Assemb1v Lenath Chanae Additional assembly length change data are now available for several Calvert Cliffs Unit 1 fuel assemblies with bundle average burnups of up to

(

).

These data are compared to SIGREEP predictions in Figure 4.2.2.a-1.

Based on that comparison, it is concluded that the material i

behavior equations and the analytical model (the SIGREEP computer code) are acceptable for use in predicting the irradiation induced length change of fuel assemblies with [

] guide tubes to guide tube fast fluences of [

].

4.2.2.4.a Evaluation of Shoulder Gao Chanae a) 14x14 Fuel Assembly Design Additional shoulder gap change data are now available for Calvert Cliffs Unit 1

fuel rods with axial average burnups of up to

[

].

Representative data are plotted on Figure 4.2.2.a-2, along with predictions using SIGREEP.

As discussed in Reference 1, the original [

]

growth equation (the upper equation in Table 4-7 of Reference 1) was used in the SIGREEP computer runs.

Based on the comparison of the data and predictions in Figure 4.2.2.a-2, it is concluded that the SIGREEP methodology is acceptable for use in evaluating shoulder gap change in 14x14 fuel assemblies to fuel rod axial average fast fluences of [

].

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4.2.3.a Fuel Assembly Holddown The discussion provided in Section 4.2.3. of Reference 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 6 -

i

4.2.4.a Grid Irradiation Growth The 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 1 agreed well with all the other growth mear:urements [

] presented in that reference, the good agreement between the growth measurements and predictions for [

] presented in other sections of this document supports the adequacy of the grid irradiation growth model to extended burnup.

I J

J l i

4.2.5.a S.pAger Grid Relaxation I

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 PROTOTYPE extended burnup demonstration assemblies irradiated in Calvert Cliffs Unit I confirm that the relaxation of the fuel assembly spacer grid springs is not a concern for extended burnup operation.

i I

i e

1 ;

i 4.2.6.a Corrosion of the Fuel Assembiv Structure The following paragraphs append Subsection 4.2.6.3 of Reference 1.

4.2.6.3.a Evaluation of Corrosion of the Fuel Assembly Structure Additional in-reactor corrosion data will be obtained from hot cell 1

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 l

cracking, were observed.

The hot cell data, which will be available in l

1989 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 Calvert Cliffs, the corrosion of the Zircaloy-4 structure will not preclude the operation of C-E fuel assemblies to 1-pin burnups of 60 MWD /kg.

e

l 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 0 -8 C Pellet Swellina.

The swelling of the burnable poison material, g3 4 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 l

dispersion of boron carbide (8 C) particles in an alumina (A10 ) matrix.

4 23 The B 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 [

].

Since the A1 0 swelling is the dominant contributor to pellet swelling at 23 high exposure, the A10 -8 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 B 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 A1023 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) indicates that for gross overall dimensional changes, a two-stage swelling rate model is an appropriate representation for A10 swelling.

That is,

)

23 21 2

above a fast fluence of approximately 2.6 x 10 n/cm, the swelling of

]

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 1

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 A1 0 swelling is represented by the following expression j

23 that accounts for the matrix swelling of A1 0 U"IY 23 I

1 l

l l

i 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 A1 0 swelling in the range of 300 to 600*C.

23 1

[

f

]

a two-stage model is used for the composite A10 -8 C pellet swelling model.

]

23 4 The volumetric swelling rate for 8 C (i.e., 30% at 100% B-10 depletion) was j

4 used in conjunction with Equation (1) for Al 0 to arrive at the following 23 expressions for the volumetric swelling of the composite Al 0 -0 C pellet.

23 4 j

j I 4 1

I

l f

The above relationship for swelling-as a function of fluence is plotted in Figure 4.2.7.a-1 for Al 0 -8 C with a B C content of 3 wt%.

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 A10 and A10 -8 C pellets.

These data 23 23 4 consist of direct diameter measurements on 42 whole A10 -8 C and 16 whole 23 4 Al 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 behavior (i.e., equal axial and diametral swelling rates).

It was also found that swelling was independent of fnitial 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 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 conservatively 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 4 I

cladding strain as affected by pc11et swelling and (2) by inference, the restrained swelling behavior of the Al 0 -B C pellets.

It was found that l

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 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 -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.

A comparison of the performance data with the model in Figure 4.2.7.a-1 indicated the following:

o The swelling of A10 '0 C pellets, as well as that of Al 02 3 pellets, 23 4 that occurred during the first-cycle of irradiation up to a fluence of 21 2

about 3.5 x 10 n/cm (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 bahavior mechanisms:

(a)

B C particle swelling caused by the B-10 (n,o) Li-7 reaction 4

induces microcracking and grain boundary separation in the pellet structure. __ _ _ _ _ -

t 4

(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-cycle swelling).

(c) At higher fluence (i.e., after 100% B-10 depletion) some of these new voids are accommodating the Al 0 matrix swelling due to l

23 N

cladding restraint.

Once the accommodation is completed, diametral swelling, and therefore, volumetric swelling, would proceed at the swelling rate indicated by Equation (3).

The subsections of Gas Release, Poison Rod Growth, and Poison Rod Claddina Creep 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 mechanism which will produce changes in the poison rod internal pressure through changes in the void volume and the amount of helium released.

Calculation of the E0L 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 peilets produce only a small amount of heat flux due to gamma heatirg), and the void volume (reflecting changes due to different temperatures, pellet swelling, poison rod growth, and cladding creepdown).

s Also, for the extended-burnup reference designs, pellet open porosity at BOL is nonexistent (Table 4.2.7.a-1).

r..

E

l 4.2.7.2.a Effect of Extended Burnuo on Burnable Poison Rod Behavior Al;0 -8 C Pellet Swellina.

The swelling of A10 -B C pellets is strongly 3 4 23 4 fluence dependent; therefore, the mechanical behavior of the burnable s

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 l

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 Creeo.

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 less than the limiting fuel

~

rod growth prediction.

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.

d Rod Internal Pressure.

Internal pressure will increase during extended burnup operation due to a 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 accounted for.

No additional gas is predicted to be released from the pellets due to extended burnup.

59-

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 j

design improvements in the extended-burnup poison rod designs, they will i

demonstrate that there is margin to the strain, clearance, and internal pressure criteria for the poison rods.

e T,

Table 4.2.7.a-1 Burnable Poison Rod Details Extended Extended Early Burnup Early Burnup i

Parameter 14x14 Desian 14x14 Desian 16x16 Desian 16x16 Desian Pellet 0.D.,

0.376-0.379 0.362 0.310 0.307 in.

l 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.

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CONCLUSION The objective of this report is to justify the validity of C-E methods and l

models concerning Calvert Cliffs fuel 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 l

presented in this report justifies the extension of this 1-pin limit to the new 1-pin limit raquired by the implementation of 24-month cycles at -the Calvert Cliffs Units, 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.

Also, since the various fuel performance topics discussed in Reference 1 have no explicit deper.dence on batch average burnup, the batch average discharge limit of Reference 1 is no longer required and can be deleted.

i I

1 l

l

~

1 i

1

• i

REFERENCES 1

CENPD-269-P, Rev.1-P, " Extended Burnup Operation of Combustion Engineering PWR Fuel," July 1984.

4.1.2.a-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-2 A. M. Garde, " Hot Cell Examination of Extended Burnup Fuel From Fort Calhoun," D0E/ET/34030-11, CEND-427, Combustion Engineering, Inc., September 1986.

4.1.2.a-3 G. P. Smith, "C-E/00E Corrosion Data on ANO-1 Unit 2 Fuel Rods,"

to be issued, October 1988.

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 Properties under LWR Coolant Conditions (Part II)," RPX101-01, Final Report, NFIR Report, NFIR PP-01-702, Nuclear Fuel Industry Research Group, October 1987.

4.1.2.a-7 L. W. Nevman, "The Hot Cell Examination of Oconce 1 Fuel Rods After Five Cycles of Irradiation," 00E/ET/34212-50, BAW-1574, Babcoc'k & Wilcox, October 1986.

I,

____m__. - _ _ _ _ _. _ _ _ _ _. _ _ _ _ _.. _ _

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 j

Report EP80-16, Final Report, Empire State Electric Energy Research Corporation, April 1985.

1 4.1.2.a-9 U. P. Nayak, H. Kunishi and W. R. Smalley, " Hot Cell Examination of Zion Fuel Cycle 5," Research Report EP80-16, Final Report, h

Empire State Electric Energy Research Corporation, June 1985.

4.1.2.a-10 R. S. 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 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.

4.1.3.a-1 H. 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 I "CEPAN Method of Analyzing Creep Collapse of Oval Cladding,"

EPRI NP-3966-CCM Volume 5, April 1985. -

4.1.5.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.

4.1.5.a-2 L. W. Newman, "The Hot-Cell Examination of Oconee 1 Fuel Rods After Five Cycles of Irradiation," DOE /ET/34212-50, BAW-1874, Babcock and Wilcox, October 1986.

4.1.5.a-3 U. P. Nayak, H. Kunishi and W. R. Smalley, " Hot Cell Examination of Zion Fuel, Cycle 5," WCAP-10543, Final Report EP80-16, Empire.

State Electric Energy Research Corporation, June 1985.

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 1 through 4,"

WCAP-10473, Final Report EP80-16, Empire State Electric Energy Research Corporation, April 1985.

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.

TM 4.1.5.a-6 System 80 Standard Safety Analysis Report Final Safety Analysis Report (CESSAR FSAR), STM-50-470 F, Combustion Engineering, Inc.,

October 1978.

'u 4.1.5.a-7 J. F. McLehan. " Yankee Core Evaluation Program, Final Report,"

WCAP-3017-6094, Westinghouse Atomic Power Division, January 1971.

i l

4.1.5.a-8 R. L. Knecht and P..l. Pankaskie, "Zircaloy-2 Pressure Tubing,"

]

l BNWL-746, Battelle Pacific Northwest v.aboratory, 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-10.J. E. Irvin, " Effects of Irradiation and Environment on the Mechanical Properties and Hydrogen Pickup of Zircaloy," Zirconium and Its A11gy.i. Electrochemical Society, New York, NY,1966.

4.1.5.a-11 W. Evans and G. W. Parry, "The Deformation Behavior of Zircaloy-2 Containing Directionally Oriented Zirconium Hydride Precipitates," Electrochem. Tech.. 4, 225 (1966).

4.1.5.a-12 W. A. Pavinich and T. P. Papazoglou, " Hot Cell 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," Mator. Sci. Eno. 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 Zr-2.5 wt% Nb," Can. Met. Quart.. 12, 285 (1973).

4.1.5.a-15 E. F. Ibrahim, " Creep Ductility of Cold-Worked Zr-2.5 wt% Nb and Zircaloy-2 Tubes In-Reactor," J. Nucl. Mat.. 96, 297 (1981).

4.1.5.a-16 D. 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

8. Watkins et al., " Embrittlement of Zircaloy-2 Pressure Tubes,"

Acolications Related Phenomena for Zirconian and Its Alloys, ASTM-STP-458, 1968.

4.1.5.a-18 M. A. Shubert, " Examination of the PROTOTYPE and 1H038

)

i Assemblies After Reactor Cy:le 9 in Calvert Cliffs Unit 1,"

l CENPSD-493-P, January 1989. '

(

l l-ll-l 4.1.6.a-1

" Improvements to Fuel Evaluation Model," CEN-161(B)-P Supplement l

'l-P, Combustion Engineering, Inc., April 1986.

4.1.6.a-2 Letter from S. A. McNeil (NRC) to J. A. Tiernen (BG&E), " Safety l

Evaluation of Topical Report CEN-161(B)-P Supplement 1-P, Improvements to Fuel Evaluation Model," February 4,1987.

~

l 4.1.6.a-3 A. M. Garde, " Hot Cell Examination of Extended Burnup Fuel from Fort Calhoun," DOE /ET/34030-11, CEND-427, Combustion Engineering, September 1986.

4.1.6.a-4 U. P. Nayak et al, " Hot Cell Examination of Zion Fuel Cycle 5,"

WCAP-10543, Westinghouse, June 1985.

4.1.6.a-5 S. R. Pati, A. M. Garde and L. J. Clink, " Contribution of Pellet Rim Porosity to Low Temperature Fission Gas Release at Extended Burnups," Proc. ANS Tooical 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," 00E/ET/34010-11, CEN-415, November 1983.

4.1.7.a-1 U. P. Nayak, et al, " Hot Cell Examination of Zion Fuel Cycle 5,"

WCAP-10543, Wer.tinghouse, June 1985.

l 4.1.7.a-2 A. M. 'larde, " Hot Cell Examination of Extended Burnup Fuel from i

Fort Calhoun," 00E/ET/34030-11, CEND-427, Combustion Engineering, September 1986.

4.1.8.a-1 B. J. Wrena, et al, " Thermal Properties of Urania-Erbia,"

Battelle Northwest Laboratories, dated June 1988.

! '__________________________?___-_____-

?

l l

l 4.1.8.a-2 J. Komatsu, et al, "The Melting Temperature of Irradiated Fuel,"

J. Nuclear Materials, No. 154 (1988), pp. 38-44.

l 4.1.9.a-1 A. M. Garde, " Hot Cell Examination of Extended Burnup Fuel from Fort Calhoun," DOE /ET/34030-11, CEND-427, Combustion Engineering, l

September 1986.

4.1.9.a-2 U. P. Nayak, et al, " Hot Cell Examination of Zion Fuel Cycle 5,"

WCAP-10543, Westinghouse, June 1985.

I.

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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.

4.1.14.a-1 D. E. Bassette et al., "C-E/EPRI Fuel Performance Evaluation Program RP586-1 Task A: Examination of Calvert Cliffs I 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 T,ssk A, Examination of Calvert Cliffs 1 Test Fuel Assembly After Cycle 3," CENPSD-87, Combustion Engineering, Inc., September 1979. '

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i s

g cy 4

4.1.14.a-3 E. J. Ruzauskas et al., "C-E/EPRI Fuel Performance Evaluatica l '. _

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l 4.1.14.a-4 ' R. G. Weber et al., "EPRI/C-E Fuel Performance Evaluation Program, RP586-1 Task B: Examination of Arkansas Nuclear One-Unit 2 Characterized' Fuel Assemblies After' Cycle 1,"

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'e

' 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.

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4.1.14.a-7 G. P. Smith, "The Nondestructive Examination of Fuel Assemblies with Standard and Advanced Design Rods After Three Cycles of Irradiation," DOE /ET/34013-12, CEND-426, Combustion Engineering, Inc., November 1986.

/

4.1.14.a-8 M. A. Shubert, "Exarrination of the PROTOTYPE and 1H038 Assemblies After Reactor Cycle 9 in Calvert Cliffs Unit 1,"

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4.1.14.a-9 "In-Reactor Dimensional Changes in Zircaloy-4 Fuel Assemblies,"

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.l 4.1.14.s-10 D. G. Franklin, "Zircaloy-4 Cladding Deformation During Power.

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TM

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0xide Exposed to 5x10 Fast Neutrons /Cm," Nuclear Acolications, 3, 686, November 1967.

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