Text

Extended-Burnup Lead Test Assembly-Design Rept
	ML19340B637
Person / Time
Site:	Arkansas Nuclear
Issue date:	10/31/1980
From:	; BABCOCK & WILCOX CO.
To:	;
Shared Package
ML19340B633	List: ML19340B632; ML19340B637;
References
	BAW-1626, NUDOCS 8011110508
	Download: ML19340B637 (47)
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i e

BAW-1626 October 1980 i

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lI EXTENDED-BURNUP LEAD TEST ASSDiBLY l

- Design Report -

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!I II II lfl BABCOCK & WILCOX i5 Nuclear Power Group l

Nuclear Power Generation Division P. O.

Box 1260 Lynchburg, Virginia 24505 i

Boilit0546 Babcock & Wilcox

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CONTENTS Page 1.

INTRODUCTION.

1-1 I

2

SUMMARY

2-1 3.

LEAD TEST ASSEMBLY DESIGN 3-1 3.1.

Base Design 3-1 3.2.

Annular Pellet Fuel Rod 3-1 I

3-2 3.3.

Segmented Rods.

4.

FUEL SYSTEM DESIGN.

4-1 4.1.

Fuel Assembly Mechanical Design 4-1 4.1.1.

Hardware Changes in Structural Cage 4-1 4.1.2.

Removable Rod End Fitting Design.

4-2 3

4.2.

Fuel Rod Design 4-3 4.2.1.

Design Analyses - Solid Pellet 4-3 4.2.2.

Design Analyses - Annular Pellet 4-6 4.2.3.

Design Analyses - Segmented Rods.

4-6 4

4.3.

Material Design 4-8 5.

NUCLEAR DESIGN.

5-1 l

5.1.

Physics Characteristics 5-1 5.1.1.

LTA Solid Pellet Fuel Rod 5-1 g

5.1.2.

Annular-Pellet Fuel Rod 5-2 g

5.2.

Segmented Rod Analyses.

5-2 5.3.

Assembly Characteristics.

5-2 6.

THERMAL HYDRAULIC DESIGN.

6-1 61 DNBR Analysis 6-1 6.2 Hydraulic Lift Analysis 6-2 7.

ACCIDENT AND TRANSIENT ANALYSES 7-1 CLOSSARY.

A-1 REFERENCES.

B-1 R

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Babcock & Wilcox

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List of Tables Table Page g

3-1.

Lead Test Assembly Components 3-3 5

3-2.

Lead Test Assembly Fuel Rod Designs 3-4 3-3.

Segmented Rod Design Parameters 3-5 List of Figures j

,Ii' Figure 3-1.

Mark BEB Fuel Assembly, General Arrangement 3-6 3-2.

Fuel Rod Placement Chart, Segmented Fuel Rod h1 Fuel Assemblies.

3-7 3

3-3.

Fuel Rod Placement Chart, Non-Segmented Fuel Rod Fuel Assemblies.

3-8 g

3-4 Mark BEB Pressurized Fuel Rod Assembly 3-9 g

3-5 Mark BEB Annular Fuel Pellet 3-10 3-6.

Segmented Rod Vs Fuel Assembly 3-11 3-7.

Leng Fuel Rod. Segment 3-12 3-8.

Short Fuel Rod Segment 3-13 4-1.

Removable Fuel Rod End Fitting Components.

4-9 4-2.

Removable Fuel Rod UEF Threaded Ring 4-10 j

4-3.

Removable Fuel Rod UEF Components.

4-11 3

4-4 Removable Fuel Rod UEF Components.

4-12 4-5.

Assembled Removable Fuel Rod UEF 4-13 4-6.

Fuel Rod Internal Pressure Vs Burnup for Full-4-14 Length Rods.

4-7.

Intermediate Plug Design 4-15 4-8.

Fuel Rod Segment Internal Pressure Vs Burnup 4-16 l

5-1.

Core Loading Map for ANO-1, Cycle 5.

5-4 5

5-2 PDQ Relative Assembly Powers at 0 EFPD - 177-FA Core, ANO-1, Cycle 5 5-5 I

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Babcock & Wilecx 1

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I 1.

INTRODUCTION I

Increased fuel burnup is widely acknowledged as a straightforward and readily backfittable means for reducing uranium requirements in light water reactors operating in the "once-through" fuel cycle.

Babcock & Wilcox (B&W), in con-junction with the U.S. Department of Energy and Arkansas Power & Light (AP&L) is engaged in a program to develop and demonstrate an extended burnup fuel as-sembly capable of burnups in excess of 50,000 mwd /mtU.1,2 The in-reactor dem-onstration phases of this program call for AP&L to irradiate four first-phase extended burnup 15 by 15 lead test assemblies in the arkansas Nuclear One, l' nit 1 (ANO-1) reactor during cycles 5, 6, and 7.

This report describes and justifies the design of the first-phase extended burnup lead test assemblies (LTAs), which are similar in design to standard I

15 by 15 fuel assemblies except for changes to the fuel rod and fuel assembly structural cage to extend their burnup capability. All four LTAs are to be extensively characterized before irradiation and examined after each cycle of operation.

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l-1 Babcock & Wilcox

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1 lI 2.

SIM!ARY l

l The extended-burnup LTA is termed the Mark BEB.

Four such assemblies will be loaded in the ANO-1, cycle S core. Two of the four will contain segmented fuel rods and will have a specially designed end fitting for removal of these rods in the reactor fuel pool upon assembly discharge.

The segmented rod de-sign, based on the Mark BEB full-length rod, comprises five individual fuel segments, three of which are representative of a full-length rod.

The base Mark BEB design employs fuel rods of a solid pellet design; however, four full-length rods in each assembly and certain segments of the segmented rods contain annular pellets to gain incore high burnup experience with an l

annular fuel design (Mark BEB-A).

The annular pellet was selected because of lE its I wer perating fue temperatures, which result in significantly reduced lW fission gas release from the fuel matrix.

Lower end-of-life fuel rod internal t

pressures result from the annular pellets' combination of lower operating tem-ig peratures and the increased void volume from the pellet central void, i

The heat treatment for the guide tube and instrument tube material was changed from stress reliev'.ng to full annealing to reduce fuel assembly irradiation j

growth, which has been identified as a limiting condition for extended burnup

)

operation of standard Mark B fuel.3 i

Based on mechanical, nuclear, and thermal hydraulic analyses, the loading of feur extended-burnup LTAs in the ANO-1, cycle 5 core will not adversely affect the performance characteristics of the reactor and will be bounded by existing safety analyses.

I I

I 2-1 Babcock & Wilcox il

lI II lI 3.

LEAD TEST ASSEMBLY DEEIGN l

Four extended-burnup LTAs are being fabricated for insertion in the ANO-1, cycle 5 core. The base design for the LTAs, along with some features unique to individual LTAs, acc described below.

Two variations to the base fuel de-sign included in the LTAs are annular peilets and segmented fuel rods. Table 3-1 lists the major components of the LTAc. Figures 3-2 and 3-3 illustrate in detail the locations of the various fuel rod types.

,3.1.

Base Design The LTA is a Mark BEB (15 by 15) fuel assembly that has been designed for ex-tended burnuo (>50,000 mwd /mtU) operation; the assembly is shown in Figure 3-1.

Outside dimensions and external interfaces for both the Mark B and Mark BEB assemblies are the same.

In addition, the envelope dimensions of the base extendei-burnup fuel rod de-i sign (Mark BEB, Figure 3-4) are identical to those of the standard Mark B.

A fuel rod design with extended-burnup capability was obtained by (1) reducing the fuel column stack height to increase plenum volume, (2) decreasing fuel i

rod initial fill gas pressure to reduce end-of-life (EOL) internal pressure, and (3) increasing the cladding thickness to provide a more creep resistant l

rod.

The pertinent fuel rod design parameters and dimensions are given in Table 3-2.

.l l

W All the basic fuel rod internal components, e.g., upper and lower spring spac-ers and upper and lower tubular spacers, are similar in design to those of standard Mark B fuel rods with only slight dimensional changes to naintain interface dimensions.

3.2.

Annular Pellet Fuel Rod (Mark BEB-A)

Some fuel rods in each LTA will be loaded with annular pellets which have a nominal inside diameter of 0.115 inch (Figure 3-5).

The selection of this design is based on the following:

3-1

' Babcock & Wilcox

I 1.

The annulus reduces the maximum fuel temperature. The lower maximum fuel temperature aids in the reduction of fission gas release and in turn re-g duces EOL fuel rod internal pressure.

W 2.

The additional void volume of the annulus contributes to reduced EOL fuel rod internal pressure.

3.

The annular pellet fuel rod has lower predicted EOL creep than the solid g

pellet fuel rod because of a higher beginning-of-life (BOL) internal pres-W sure under operating conditions. The higher BOL pressure is created by the elevated gas temperature within the central annulus of the pellet.

Therefore, the pressure differential across the cladding, which causes creepdown, is reduced.

4.

The reduced smear density of the annular pellet causes an increase in the hydrogen-to-uranium atom ratio, yielding improved uranium utilization.

Annular pellet fuel rod dimensions are presented in Table 3-2, which shows that the annular and solid pellet fuel rod designs are similar. Mixed-oxide annular fuel pellets of similar design (10 vol %) have operated successfully in the past."-'

3.3.

Segmented Rods Two of the four LTAs will each contain eight segmented fuel rods. The seg-mented rod design (comprising five individual fuel segments) is based on both the Mark BEB and Mark BEB-A full length rods; it provides three segments which are essentially abbreviated versions of the full length rods.

Two long sections, identical in design and length, and a shorter middle seg-ment are the three components intended for possible use in a test reactor.

Eight of the 16 middle segments will contain annular fuel pellets and the other eight solid fuel pellets.

Sixteen of the long segments will be loaded with annular fuel pellets and 16 with solid pellets. The upper and lower end segments contain solid fuel pellets and complete the rod, making its length equivalent to the Mark BEB and Mark BEB-A full-length rods. Active fuel lines at the top and bottom of the assembly coincide with those of full length rods.

Figure 3-6 depicts the relationship of a segmented rod to a fuel assembly.

Figures 3-7 and 3-8 show the internal configurations of the long and short sec-tions, respectively. Hafnia-yttria pellets act as flux suppressors in the I

3-2 Babcock & Wilcox

I I

coupling region between the sections. The spring, which holds the internals in place during shipping and handling, also permits fuel growth during opera-tion. The individual segments are joined by inserting their intermediate 4

plugs into a cladding sleeve to which they are welded. Design parameters for the segmented rods are presented in Table 3-3.

Table 3-1.

Lead Test Assembly Components in No. of components

)

~g per assembly j

NJ023P("

NJ023R E

I Assembly component NJ023Q

?UO23s Removable rod upper end fitting 1

4 j

Mark B upper end fitting 1

Mark B lower end fitti<..g 1

1 Mark BEB fuel rod 196 204 Full-length annular pellet fuel rod 4

4 Segmented fuel rod 8

.g Annealed guide tubes 16 16

.5 (a) Assembly identification.

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.I 3-3 Babcock & Wilcox

I Tab 2 3-2.

Lead Test Assembly Fuel Rod Designs Mark B Mark BEB Mark BEB-A (nominal) solid pellet annular pellet Parameter design design design Cladding W

l OD, in.

0.4300 0.4300 0.4300 j

ID, in.

0.3770 0.3710 0.3710 Wall, in.

0.0265 0.0295 0.0295 Pellet OD, in.

0.3686 0.3635 0.3635 ID, in.

0.115 1

l Length, in.

0.600 0.418 0.418 Density, % TD 95 95 95 WI i

ruel Rod Stack height, in.

141.80 138.251 138.25 Fuel rod length, in.

is 6875 153,625 153.625 UO loading, g 2528.6 2400.0 2200.0 i

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3-4 Babcock & Wilcox I

1-

]W Table 3-3.

Segmented Rod Design Parameters I

Lower Upper end Long Short end Fuel rod j

segment segment segment segment total No. per assembly 8

16 8

8 8

l No, per full-length rod 1

2 1

1 5

Segment length, in.

25.932 43.307 20.005 21.077 153.625 Cladding length, in.

25.057 42.120 18.818 20.2025 148.32 Fuel stack length, in.

14.75 34.75 14.00 10.813 109.63 i

Fuel stack UO, loading, 256.06 603.25 243.03 187.71 1893.29 i

g 553.05 222.81 1772.67 Type of pellet Solid Solid Solid Solid annular annular i;g Lower spring free 3.5 None None None J

1ength, in.

i Upper spring free 2.4687 3.4687 2.4687 6.5928 length, in.

No. of hafnia pellets 2

5 3

1 16 t

l No. of tubular spacers 2

2 1

2 9

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3-5 Babcock & Wilcox l

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l Figure 3-1.

Mark BEB Fuel Assembly, General Arrangement

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Fuel Rod Placement Chart, Segmented Fuel Rod Fuel Assemblies 1

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LEGEND

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@ PRECHARACTERIZED SOLIO FUEL PELLET R00 ASSEMBLIES - 20 EACH I

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@NON-REMOVABLESEGMENTEDFUELR00 ASSEMBLIES-4EACH

@ REMOVABLE SEGMENTED FUEL R00 ASSEMBLIES - 4 EACH X Gul0E TUBE LOCATION - 16 EACH I

Note: Remaining locations are filled with non-precharacterized solid fuel pellet rod assemblies, 176 each.

3-7 Babcock & Wilcox

I Figure 3-3.

Fuel Rod Placement Chart, Non-Segmented Fuel Rod Fuel Assemblies A

I 2

3 4

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@ PRECHARACTERIZED SOLID FUEL R0D ASSEM81.lES - 20 EACH R INSTRUMENT TUBE LOCATION - 1 EACH x GUIDE TUBE LOCATIONS - 16 EACH Note: Rema ining locat ions are filled with non-precharacterized solid fuel pellet rod assemblies, 184 cach.

3-3 Babcock & Wilcox I

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3-10 Babcock & Wilcox I

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Segmented Rod Vs Fuel Assembly i

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3-11 Babcock & Wilcox i

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FUEL SYSTD4 DESIGN ll W

4.1.

Fuel Assembly Mechanical Design

g The LTA analysis includes those areas in which the design or service condi-tions of the assemblies differ from those considered in the evaluation of the sta

.rd Mark P fuel r sembly. Where no differences exist, the analyses per-I formed on the standard Mark B assembly apply.

Normal operation, transient events, emergency and faulted conditions, and handling are all addressed M l

the LTA analyses, but it is primarily in the area of normal operation that substantial differences exist. The changes in assembly hardware and the anal-yses of these

~.anges are discussed below.

4.1.1.

Hardware Changes in Structural Cage The structural cage is defined as all components of the fuel assembly except the fuel rods.

Changes were made in three areas:

guide tubes and instrument tube, assembly holddown springs, and upper end fitting.

Fuel assembly growth has been identified as a limiting condition for extended burnup operation of standard Mark B fuel.3 To reduce fuel assembly irradia-tion growth, the guide tube and instrument tube material heat treatment for the LTAs has been changed from stress-relief to full recrystallization anneal-ing since the growth rate of the fully annealed material is about one-quarter of that for cold worked stress-relieved material. Weld strength and clastic

w buckling, the limiting structural criteria for the assembly, are unaffected by the material change; thus, this change does not reduce the load-carrying ca-pacity of the guide tube assembly. With the alteration in the guide tube heat treating process, the LTA has an assembly burnup limit of 62,000 mwd /mtU from fuel assembly growth, which exceeds the target burnup of 50,000 mwd /mtU for the LTAs.

Because of the decrease in fuel assembly irradiation growth, fuel rod growth becomes the limiting constraint in the LTA design, yiciding an as-sembly burnup limit of 58,000 mwd /mtU.

I 4-1 Babcock & Wilcox

I The LTA is about 24 lb lighter than the standard Mark B assembly because of a reduction in fuel loading. To compensate for this weight reduction, the g

Inconel X-750 (No. 1 temper) holddown spring of the standard Mark B assembly W

has been replaced with a stronger Inconel 718 spring. As in the standard Mark B spring, the dimensions are (1) wire diameter 0.472 in., (2) coil diameter 4.665 in., and (3) free height 5.9 in.

The Inconel 718 spring increases the minimum holddown force by 70 lb producing an increase in net holddown force of 46 lb.

In addition, the increased strength of Inconel 718 results in im-proved stress margins for springs fabricated from this material when compared to the Inconel X-750 r.pring.

4.1.2.

Removable Rod End Fitting Design The upp r end fittings (UEF) of two lead test assemblies allow for removal of four fuil rods from each assembly at the reactor site. The requisites for the removable rod UEF design were as follows:

1.

Allow removal of designated fuel rods on-site, preferably in the spent fuel pool.

2.

Permit reinsertion of a dummy rod in the vacated location.

g 3.

Maintain compatibility with handling equipment and interfaces of the other W

fuel componen;s.

Each removable rod has a special end cap (see Figure 4-1).

The corresponding UEF has a threaded ring welded in the grillage (Figure 4-2).

A hollow bolt and locking ring are installed subsequent to the positioning of the end fit-ting.

The bolt is hollow to accommodate the fuel rod end cap, which passes through it.

After assembly of the components, both sides of the locking ring g

are crimped onto the slotted area of the bolt (Figure 4-3).

=

Since the end fitting grillage is a highly redundant structure, weakening of a local area merely shifts the stress to other webs, resulting in a negligible change in maximum stress. The analysis of the grillage conservatively assumes g

no structural contribution from the plug and shows an increase of less than 1%

W in maximum stress. The plug itself has also been analyzed and has been shown S

adequate for impact by a fuel rod under accident conditions and for normal p

operating loads.

After irradiation, the rod can be removed in the spent pool by unlocking and completely unscrewing the bolt with a special mating tool and inserting an El Babcock & Wilcox 4-2

15

[g instrument which lifts the rod, bolt, and ring in a single operation.

A re-placement rod can be installed by reversing this procedure.

Figure 4-2 shows the location of the threaded rings in the UEF.

An exploded view of the parts as they mate is presented in Figure 4-3.

Figure 4-1 shows

!W the completed assembly. Prototype parts and the a1sembled prototype are shown in Figures 4-4 and 4-5.

The assembling procedure for the fuel assemblies with the removable rods is the same as for the standard Mark B assembly.

4.2.

Fuel Rod Design

=

The LTAs will contain solid pellet fuel rods, rods loaded with annular pellets, and segmented fuel rods.

In order to evaluate the performance of these fuel i

rods, various design analyses were performed, a

4.2.1.

Design Analyses - Solid Pellet

.g All the LTAs will have fuel rods containing solid pellets. The analyses de-scribed below were performed to determine the effects of extended burnup on ju this design.

as 4.2.1.1.

Cladding Collapse

!5 Using the CROV computer code *, the fuel rod was designed to preclude creep collapse within the design life. The creep collapse analyses were performed l

using power histories that track the most limiting assembly so that the most l

1imiting collapse time was obtained.

The collapse time was conservatively determined to exceed a design life of 35,500 EFPH (corresponding assembly burn-up 50,000 mwd /mtU) which is greater than the anticipated LTA incore residence time of m31,000 EFPH.

ll 4.2.1.2.

Cladding Stress and Strain

(

lW Stress and strain limits are imposed to ensure that the cladding stresses are l

less than the allowable material strength and that the strain capability of the cladding is not exceeded.

The following design critaria were used for the stress and strain analyses:

I I

See clossary, page A-1.

4-3 Babcock & Wilcox

I 1.

Primary membrane stresses (which are not relieved by small material de-formation) are not to exceed two-thirds of the minimum unirradiated yield strength.

d W

2.

Primary membrene plus ben ing stresses are not to exceed the minimum un-irradiated yield strength.

3.

The average circumferential strain is not to exceed 1% inelastic strain

(+0.4% elastic strain).

The stress analysis was performed using thich shell equations with stresses evaluated at both the inside and outside diameters. This analysis follows the format and procedures outlined in Section III of the ASME Boiler and Pres-sure Vessel Code (1971) generally used to organize stresses into various cate-a gories; the stresses are combined to determine stress intensity. The cladding 7

strain criterion above, based on work by O'Donnell, addresses failure due to plastic instability of the cladding. The criterion described in reference 7 shows that the allowable hoop strain in the temperature range of interest

(>600F) is 2%.

Hence, the use of 1% as a criterion is conservative.

Using the aforementioned techniques, the Mark BEB solid pellet fuel rod has been designed to operate to a maximum fuel rod average burnup of more than 60,500 mwd /mtU.

4.2.1.3.

Cladding Fatigue Combinations of system operating transients were evaluated to ensure that the cumulative usage factor as defined by the ASME Code,Section III, Paragraph NB-5222.4, would be less than 0.9 of the allowable material fatigue life (this is an additional conservatism over the ASME Code recommendation of 1.0).

The system transients considered were events causing cyclic stress, such as heat-up and cooldown.

A cumulative usage f actor was calculated from f atigue curves based on the O'Donnell and Langer curves.a The cumulative usage factor for the Mark BEB solid-pellet fuel rod design was lower than the design criterion of 0.9.

4.2.1.4.

Thermal Design The thermal design objective for the extended-burnup LTA fuel is to provide a conservative fuel rod design based on the following criteria:

Il I

Babcock 8.Wilcox 4-4 I

I 1.

The LTA shall not restrict core allowable local power limits (kW/ft);

that is, the LTA fuel shall not be limiting in terms of linear heat rate E

to fuel melt (HRTM) relative to the Mark B fuel. The Mark B design com-prises the remainder of t'ae ANO-1 core and forms the basis for the maxi-mum allowable heat rates used in reload fuel cycle design and in reactor protection system trip limits.

2.

The LTA fuel rod (or segment) internal pressure shall not exceed nominal reactor coolant system pressure (2200 psia) during normal operations up to a rod average burnup of 60,500 mwd /mtU.

The HRTMs and rod internal pressures were calculated using the TACO 2 computer code. " This fuel performance program includes models for fuel densification, swelling, cladding creep, and fission gas production and release. The TACO 2 calculations were based on bounding fuel densification kinetics, in which the maximum densification was assumed for temperature and heat rate calculations lg to maximize the fuel-cladding gap and minimize the active fuel stack height, 5

ind the minimum densification was assumed for pressure calculations to minimize the rod free volume. This technique yields conservative predictions for both i

fuel temperature and rod internal pressure.

Both fuel rod temperature and internal pressure are affected by the release of fission products (xenon and krypton).

Because the fission gas release model in TACO 2 is temperature-dependent, fuel rod power is important since it di-i rectly affects temperature and thus flesion gas release.

Conservatism is built into the TACO 2 predi tions by using bounding axial flux and burnup shapes and a bounding fuel rod power history.

The very conservative assumed fuel rod I

power history envelops both the ANO-1 cycle 5 fuel cycle design fuel rod peaks and burnups and the standard Mark B fuel rod power history.

The minimum HRTM for the full-length solid-pellet fuel rod was calculated to be 21.1 kW/ft based on design peaking limits.

Since this maximum allowable I

heat rate is higher than the cycle 5 minimum HRTM (20.15 kW/it based on stan-dard Mark B fuel) the full-length rod with solid pellets will not make the !.TA 3

the limiting assembly in the core.

I 4

4-5 Babcock & Wilcox

I The fuel rod maximum internal pressure of 2196 psia is shown as a function of burnup in Figure 4-6.

The predicted rod internal pressure does not exceed nominal reactor coolant (RC) system pressure, 2200 psia. The actual rod in-ternal pressure will be less than that shown in the figure by an amount de-pendent on the dif ference between the design power peaks and the actual power peaks and burnups experienced.

4.2.2.

Design Analyses - Annular Pellet Design analyses (cladding collapse, stress, strain, fatigue, and thermal eval-uation) as previously discussed in section 4.2.1, were also performed for the annular fuel pellet design. These evaluations have shown that the annular-pellet fuel rod design will operate to extended burnup with lower rod internal pressure and lower EOL creep relative to the solid-pellet rod design.

The minimum HRTM for the rod containing annular fuel pellets is greater than that calculated for both the full-length rod with solid pellets and the cur-rant Mark B fuel based on the same bounding densification kinetics and design peaking. Hence, the full-length annulcr-pellet rods are not the limiting rods in the core.

The maximum predicted internal pressure, 1635 psia, is shown as a function of burnup for the full-length annular-pellet rod in Figure 4-6.

The predicted internal pressure does not exceed nominal RC system pressure. The lower in-W ternal pressures of the annular pellets (as compared to solid pellets) is evident and is the result of both greater void volume and lower fuel tempera-tures in the annular-pellet rod.

As with the full-length solid-pellet rod, the predicted internal pressure of the annular-pellet rod will be reviewed before each subsequent cycle of LTA residence along with the comparison of the design fuel rod power history envelope to the current cycle actua:. fuel rod peaks and burnups.

4.2.3.

Design Analyses - Segmented Q,

The specification for the fuel pelletr l cladding for the segmented rod are the same as those of the Mark ESS full-length rod.

Additional mechanical analyses performed to specifically delineate the characteristics of the seg-mented rod included intermediate plug stress, creep collapse, and fuel rod l

cladding strain.

I 4-6 Babcock & Wilcox I

E 4.2.3.1.

Intermediate Plug Stress As previously mentioned in section 3.3, the individual sections of the rod are g

jof ned by inserting the intermediate plugs into a cladding sleeve to which they are welded. The design of these " plugs" must meet two requirements:

1.

The entire segmented rod, when welded together, must satisfy the full length cladding straightness criterion of 10.010 in./ft.

2.

Segment integrity must be maintained.

An added design feature of the intermediate plug is the threaded area, indi-cated by dotted lines (see Figure 4-7), which facilitates handling of the seg-ments after separation.

A mechanical analysis was performed for the intermediate plug; the following results were obtained:

the cladding, weld, and plug showed very little ther-mal stress; strese.es due to thermal gradients in the sleeve weld joining the plugs were insiginficant; the plug is structurally adequate for system pres-sure loads and withdrawal drag loads.

4.2.3.2.

Cladding Collapse Cladding creep collapse analyses were performed using the CROV* code in which the collapse time is a function of power history, temperature, changes in fuel rod pressure throughout life, fast flux, and cladding dimensions.

Except for

!=

the upper end segments, each segment was analyzed for the same initial minimum fill gas pressure as the Mark BEB full-lei.ph rod.

Since the upper end seg-ment had a small amount of fuel (due to the design of the segmented rod), its large plenum volume / fuel volume ratio resulted in low internal pressures near l

EOL.

Thus, a slightly higher initial pressure was required to provide the desired creep collapse margin. Collapse times for all segments were deter-mined conservatively to be >35,500 EFPil, which is greater than their expected incore residence time of S31,000 EFPil.

4.2.3.3.

Cladding Strain A cladding strain analysis was conducted for the segmented fuel rod using the TACO 2 computer code' to simulate pellet / cladding strain - a maximum segment lI

[

See Glossary, page A-1.

W 4-7 babcock & WilCOX

I average burnup of 62,000 mwd /mtU was modeled.

Cladding transient strain was calculated for a pellet burnup of 73,000 mwd /mtU.

Using conservative fuel rod dimensions and transient conditions, uniform transient strain was con-firmed to be less than the design limit of 1.0*.'.

4.2.3.4.

Thermal Design The thermal design criteria and methods for rod segments are the same as de-g scribed in section 4.2.1.4 for full-length rods. The minimum HRTM calculated W

for any of the rod segments is 21.1 kW/f t, which is greater than the cycle 5 minimum HRD! and equal to the full-length solid-pellet LTA rod IIRTM.

Hence, none of the rod segments will make the LTA the limiting assembly in the core.

The maximum internal pressure as a functior of burnup for each rod segment is shown in Figure 4-8.

The segment maximum internal pressures are less than those of the full-length solid-pellet LTA tod because of differences in vol-umes and fuel temperatures.

Rod inte nal premure does not exceed the nominal RC system pressure. As with the other two types of LTA fuel rods, the rod seg-ment predicted internal pressures will be reviewed before each subsequent cycle of LTA irradiation along with the comparison of the design fuel rod power his-tory envelope to the current cycle actual fuel rod peaks and burnups.

W 4.3.

Material Design The chemical compatibility of the fuel cladding / coolant / assembly interactions for the LTAs is identical to that of the standard Mark B fuel.

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Fuel Rod Internal Pressure Vs Burnup for Full-Length Rods 2500 l

NOMINAL SYSTEM PRESSURE

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Fuel Rod Segment Internal Pressure Vs Burnup i

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4-16 Babcock 8.Wilcox I

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I 5.

NUCLEAR DES 1CN The core loading map for ANO-1 cycle 5 is shown in Figure 5-1.

The Mark BEB LTAs will be located in symmetric core locations M12, M4, E4, and E12. The enrir'sent will be 2.95 wt % 23sU, the same as the standard design Mark B batch 7 fuel assemblies.

Batch / is the feed batch for cycle 5 of ANO-1.

Figure 4-2 shows the quarter core power distribution at the beginning of cycle 5.

5.1.

Physics Characteristics The LTA was modeled using the PDQ, NUL1F and DOT computer codes.* Physics characteristics for the fuel and the assembly were examined to determine the maximum power peaks that would be experienced.

5.1.1.

LTA Solid Pellet Fuel Rod The overall nuclear characteristics of the LTA solid-pellet fuel rod are simi-lar to those of the Mark B fuel rod.

However, the smaller pellet diameter (0.3635 Vs 0.3686 in., Table 3-2) causes the Mark BEB fuel rod to have a higher water-to-uranium ratio, which has two effects: (1) more neutron moderation, resulting in a slightly larger thermal-to-fast flux ratio (a positive reactiv-ity ef fect) and (2) a larger soluble boron-to-fuel ratio, which - combined with the larger thermal flux - results in a greater soluble brron reactivity wo rt h.

At the beginning of cycle 5 (BOC-5) when the soluble boron concentra-tion is large, the positive reactivity effect of the larger thermal flux is I

overshadowed by the larger absorption rate of the soluble boron. The net re-sult at BOC-5 is that the Mark BEB fuel rod has a 1.8% lower relative power density than the comparable Mark B fuel rod.

As the soluble boron concentra-tion decreases over the course of cycle 5, the positive reactivity effect of the larger thermal flux predominates, and by the EOC-5, relative power density of the Mark BEB rod exceeds that of a comparable Mark B fuel rod by 2.0%.

I

See Clossary, page A-1.

5-1 Babcock & Wilcox

t 5.1.2.

Annular-Pellet Fuel Rod Four fuel rods with annular pellets are placed in the corner fuel rod locations of each LTA.

The annulus represents a 10% decrease in fuel volume compared to that of he LTA solid pellet.

Consequently, the lower uianium loading results in a relative increase in the thermal-to-fast flux ratio and a larger soluble boron worth. The annular pellet also has a reactivity gain due to its lower average fuel temperature. Cumulatively, these changes cause a reduction in the fuel rod power density to 1.9% below that of the LTA solid-pellet fuel rod at BOL.

During the cycle, as the soluble boron concentration decreases, the power of the annular pellet fuel rod gradually increases to within 0.3% less than that of an equivalent LTA solid-pellet fuel rod.

5.2.

Segmented Rod Analyses In addition to its own plenum volume, each section of the segmented fuel rod has a coupling to connect the segmento. This coupling-plenum region creates a gap in the f uel stack, s hich can cause power peaking increases at the end of the f tel stacks and in suric'inding f uel rods if steps are not taken to control them. The power peaking effects caused by the coupling-plenum region were analyzed using the DOT two-dimensional transport code in cylindrical geometry.

The analyses were conducted for the condition of the worst tolerance buildup on g

the location and size of the coupling-plenum region.

5 The conbination of inserting 0.400 inch long hafnia-yttria pellets in the g

coupling-plenum region and placing the coupling region under th2 Inconel spacer 5

grid will reduce power peaking, thereby eliminating the need for an additional peaking penalty. The fuel pellets ac each end of the coupling-plenum region will have a power level approximately 6% below that of the non-segmented rod.

Surrounding fuel rodn in the vicinity of the coupling region will experience a power level increase of approximately 1%.

The 1% increase is well below the 2.6% local power peaking penalty already taken for peaking between Inconel spacer grids.l' 5.3.

Assembly Characteristics At BOC-5 the relative power density of the LTA in core location M12 is 1.8%

below the power of the comparable Mark B assembly in symmetric core location Nil. The LTA power gradually increases during cylce 5 to the same power as the Mark B assembly at 100 EFPD and to 2.0% above the Mark B assembly power at Babcock & Wilcox 5-2

l 5

I EOC-5.

At no time during cycle 5 does the LTA contain the highest assembly powe r o r hi ghe st radial-local peaking factor in tne core.

The nuclear characteristics of the Mark BEB design LTA represent a small de-parture f rom those of the standard Mark B fuel assembly.

Consequently, the 1,iA will have a negligible effect on the nuclear performance characterist ics of the ANO-1 core.

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5-3

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Figure 5-1.

Core Loading Map for ANO-1, Cycle 5 1

2 3

4 5

6 7

8 9

10 11 12 13 14 15 6

6 6

6 A

K4 K2 K14 K12 6

6 6

SB 6

6 6

7 7

M4 L3 N3 P6 N13 L13 M12 C

7 7

7 H13 K6 L5 L11 K10 08 D

7 7

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6 5B SB 5B SB 6

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6 6

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6 g

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6 5B SB SB SB SB 6

6 B9 E10 Fl P12 P4 N2 FIS E6 B7 SB SB 5B SB SB SB SB 5B SB L14 H14 H9 N14 011 D2 H7 H2 F2 6

6 SB SB SB 5B dB 6

6 P9 M10 L1 D14 B12 B4 L15 M6 P7 6

6 SB SB SB 5B SB 6

6 m

N9 012 K1 F14 G8 B6 K15 04 N7 6

6 5B 5B SB SB 6

6 7

7 010 L9 LTA N4 R6 RIO N12 LTA L7 06 6

6 5B SB SB 6

6 7

7 7

N11 Hll R7 B8 R9 E8 NS 6

6 6

6 7

C8 G6 F5 Fil C10 H3 6

6 6

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6 6

p E4 F3 D3 B10 D13 F13 E12 6

6 6

6 R

7 G4 G2 G14 G12 x

Batch Note: LTA: Lead test assembly.

xxx.

Cycle 4 Location I

5-4 Babcock & Wilc0X I

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l Figure 5-2.

PDQ Relative Assembly Powers at 0 EFPD -

W 177-FA Core, ANO-1, Cycle 5 8

9 10 11 12 13 14 H

10 1.26 1.17 1.32

.l.17 1.18 0.54 0.55 K

1.26 1.26 1.25 1.19 1.29 1 23 1.00 0.49 L

1.17 1.25 0.71 1.18 1.12 1.28 0.96 0.39 il M

1.32 1.20 1.19 1.08 l 1.27 '!N 1.13 0.69 i

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N 1.18 1.30 1.12

.29 1.24 1.01 0.45 I

O 1.18 1.24 1.29 1.1 '.

1.01 0.57

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'Q'fqNy x.xx g-Assembly (s.7.. @,

I 5-5 Babcock & Wilcox

5 I

I 6.

TilERMAL llYDRAULIC DESIGN The thermal-hydraulic design objective for the LTA was to provide a conserva-tive fuel assembly design based on the following criteria:

1.

The L3 shall cause no reduction sa core thermal margin-that is, the LTA shall not be the limiting assembly in the core in terms of minimum DNBR (departure from nucleate boiling ratio), and its insert ion in the core shall not reduce the thennal margin of the limiting assembly.

2.

The margin to fuel assembly lift-off shall be equal to or greater than that for the standard Mark B fuel assemblies.

These design object Ives will ensure that the LTA is bounded by the cycle 5 reload safety analyses and operational limits and that it will in no way l

restrict normal operation of the core.

l 6.1.

DNBR Analysis Thermal bydraulic analyses assume design radial power distributions and axial power shapes (1.714 radial x local x 1.5 cosine) for DNBR calculat ions per-formed in the determination of initial conditions for accident analyses, re-i actor protection system trip limits, and Technical Specification operating limits. Maximum allowable peaking (MAP) limits are generated to ensure that safety evaluations and DNBR margins based on the design radial and axial power i

distributions are maintained during actual plant operations. The MAP limits represented by a f amily of power shape curves that are equivalent to the are I

design power shape in terms of DNBR. These curves depict the limiting total fuel rod peak as a function of axial power shape and axial peak location. The MAP curves are then used in plant maneuvering margin analyses to determine operating limits for the core.

The LTA has been analyzed for both steady-state minimum DNBR and maneuvering margin. The LTA is not limiting in terms of DNBR.

Based on the cycle 5 fuel I

cock & Mcox 6-1

I cycle design peaking, the LTA minimum DNBR will be greater than the minimum DNBR in the core. Hydraulically, the LTA is virtually identical to other fuel assemblies in the core; however, the shortened LTA fuel stack height will tend to increase its surface heat flux and decrease its minimum DNBR. This is con-servatively accounted for by reducing the MAP curves used for the LTA maneuver-ing margin analysis by 1%.

Thus, it was shown that the LTA will at no time re-duce the cycle 5 thermal margins.

Similar analyses will be performed before each subsequent cycle to demonstrate that the LTA will not be the DNBR-limiting assembly in the core.

The DNBR penalty asociated with rod bow is a function of assembly burnup.

References 11,12, and 13 established a procedure for defining the penalty, I

and on the basis of this procedure, there is no DNBR penalty for fuel assem-E bly burnup values below 16,500 mwd /mtU.

Beyond this burnup value the penalty increases with burnup to a value of 6% at 40,000 mwd /mtU.

Since the predicted burnup on the LTA is less than 16,500 mwd /mtU during cycle 5, no rod bow pen-alty was applied to the LTA for cycle 5.

As discussed above, core thermal-hydraulic analyses employ design peaking dis-tributions. For typical reload cycles the assemblies most closely approaching the design peaking factor are the fresh fuel assemblies. Thus, it is antici-pated that, for subsequent cycles when the LTA burnup is high enough to result in a rod bow penalty on DNBR, the power output of the LTA (assembly radial peaking factor) will be reduced enough to fully offset the rod bow penalty.

This will be verified before each cycle of LTA irradiations as a part of the

=

thermal-hydraulic evaluation.

I 6.2.

Hydraulic Lift Anal,vsis As previously stated, hydraulically the LTA is virtually identical to the standard Mark B fuel assembly. The modification to the LTA upper end fitting to allow removal of the segmented rods has been assessed and was found to have a negligible effect on LTA hydraulic resistance. Therefore, the hydraulic lift force sn the LTA is the same as that for the standard reload fuel. The fuel assembly weights and holjddown spring forces, however, are not the same.

Be-cause of a shortened fuel stack height and annular fuel pellets, the LTA is lighter than a standard Mark B assembly. To compensate for this lighter weight.

(as discussed in section 4.1.1) a stronger holddown spring is used and the re-sulting margin to lif t for the LTA is larger than for the standard Mark B 6-2 Babcock 8.Wilcox

I

' I assembly.

llence, the LTA is not predicted to lift and is not the limiting assembly in the core.

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,I Babcock & Wilcox 6-3

'I t

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'I 7.

ACCIDE4T AND TRANS1ENT ANALYSES I

As described in section 5, the core power distribution remains almost com-pletely unaf fected by the presence of the four LTAs. Minor local reactivity perturbations do occur, but their effect is negligible because there are only four LTAs out of a total of 177 fuel assemblies in the core. However, the presence of the LTAs in the ANO-1 cycle 5 core was modeled during the genera-tion of physics parameters for cycle 5 and was evaluated as part of the cycle 5 reload analyses. The same procedure will be followed in subsequent cycles.

In terms of maximum fuel temperatures and fuel rod internal pressures, the LTAs are bounded by the standard Mark B fuel assemblies. Therefore, the current loss-of-coolant accident (LOCA) limits developed for Mark B friel are applicable for the LTAs.

The loading of four extended-burnup lead test assemblies in the ANO-1, cycle 5 core will not adversely affect the nuclear, mechanical, or thermal-hydraulic I

character of the reactor, nor will it af fect the existing safety analysis.

I I

7-1 Babcock & Wilcox

I

!I

.I GLOSSARY l3 CROV A computer code that calculates creep-induced fuel rod ovalizatton during reactor operation for creep collapse analyses. "

PDQ A computer code (PDQ07) that derives one, two-or three-d imensional l

solutions to the neutron diff usion depletion problem in one to five j

lethargy groups.15 i

NULIF A computer code that computes neutron energy spectra over the energy range 0.0 to 15 MeV and permits generation of data for PDQ tablesets.16 l

DOT A computer code (DOT 2, 3, 5) that solves two-dimensional, energy-dependent, linear Boltzmann transport equations with general aniso-tropic scattering for (X,Y), (R,Z), and (R, THETA) geometries.

TACO A computer code that computes the fuel and cladding temperature distri-but ion, fission gas production and release, cladding creep, fuel densi-fication and swelling, and fuel-to-cladding gap closure within a cylin-j drical fuel rod.'

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I REFERENCES 1

T. A. Coleman, et al., Development of an Extended-Burnup Mark B Design -

First Semi-Annual Progress Report, July-December 1978, BAW-1532-1, Babcock

& Wilcox, Lynchburg, Virginia, Novermber 1979.

2 T. A. Coleman, et al., Development of an Extended-Bumup Mark B Design -

Second Semi-Annual Progress Report, January-June 1979, BAW-1532-2, Babcock

& Wilcox, Lynchburg, Virginia, December 1979.

T. A. Coleman, et al., Qualification of the B&W Mark B Fuel Assembly for High Burnup - First Semi-Annual Progress Report, July-December 1978, BAW-1546-1, Babcock & Wilcox, Lynchbrug, Virginia, August 1979.

M. D. Freshley, " Annular Mixed Oxide Pellet Fuel Irradiations For Thermal Reactors," Trans. American Nuclear Society, 15 (1972), p.

756.

5 Plutonium Utilization in Boiling Water Reactors - Phase II, Post-Irradiation Examination of Four Fuel Rods - One Cycle Operation in Big Rock Point Reactor, NEDC-10387, July 1971.

Nondestructive Examination of Plutonium Recycle Fuel Rods Irradiated to 30,000 mwd /mtU, NEDO-12552 (1974).

7 W. J. O'Donnell, Fracture of Cylindrical Fuel Rod Cladding Due to Plastic Instability, WAPD-TM-651, Westinghouse, April 1967.

e W. J. O'Donnell and R. F. Langer, " Fatigue Design Basis for Zircaloy Components," Nuclear Science & Engineering, y (1964), pp. 1-12.

TACO 2 - Fuel Pin Performance Analysis, BAW-10141, Babcock & Wilcox, Lynchburg, Virginia, August 1979.

1" Normal Operating Controls, BAW-10122A, Babcock & Wilcox, Lynchburg, Virginia, November 1979.

ll J. H. Taylor (B&W) to D. B. Vassallo (USNRC), Letter, December 1978.

12 J. H. Taylor (B&W) to S. A. Varga (USNRC), Letter, June 22, 1979.

OCk M COX B-1

1 I

13 L. S. Rubinstein (USNRC) to J. H. Taylor (B&W), Letter, October 18, 1979.

t i

l' CROV - Program to Determir; In-Reactor Performance of B&W Fuels, BAW-10084P-A, Rev. 3, Babcock & Wilcox, Lynchburg, Virginia, October 1980.

15 Babcock & Wilcox Version of PDQ07 User's Manual, BAW-10117, Babcock &

Wilcox, Lynchburg, Virginia, January 1977.

18 NULIF - Neutron Spectrum Generator Few-Group Constant Calculator and Fuel g

Depletion Code, BAW-10115, Babcock & Wilcox, Lynchburg, Virginia, February 5

1977.

)

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I I

iIl I

I, B-2 Babcock & Wilcox

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ML19340B637

Text

SUMMARY

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