NUREG/CR-6760

ORNL/TM-2000/321

Study of the Effect of

Integral Burnable

Absorbers for PWR

Burnup Credit

Prepared by

C. E. Sanders and J. C. Wagner, ORNL

Oak Ridge National Laboratory

U.S. Nuclear Regulatory Commission

Office of Nuclear Regulatory Research

Washington, DC 20555-0001
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NUREG/CR– 6760

ORNL/TM–2000/321

Study of the Effect of

Integral Burnable

Absorbers for PWR

Burnup Credit

Manuscript Completed: August 2001

Date Published: March 2002

Prepared by

C. E. Sanders and J. C. Wagner, ORNL

Oak Ridge National Laboratory

Managed by UT-Battelle, LLC

Oak Ridge, TN 37831–6370

R. Y. Lee, NRC Project Manager

Prepared for

Division of Systems Analysis and Regulatory Effectiveness

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555–0001

NRC Job Code W6479

iii

ABSTRACT

The Interim Staff Guidance on burnup credit issued by the U.S. Nuclear Regulatory Commission’s Spent Fuel

Project Office recommends restricting the use of burnup credit to assemblies that have not used burnable

absorbers. This restriction eliminates a large portion of the currently discharged spent fuel assemblies from

cask loading, and thus severely limits the practical usefulness of burnup credit. This report examines the

effect of integral burnable absorbers (IBAs) on reactivity to provide technical justification for relaxing the

current restriction for dry storage and transportation, and subsequently, for developing the necessary

guidelines for relaxing the current restriction. The effect of IBAs on reactivity for various IBA designs is

shown and discussed. Further, the reactivity effect of IBAs for typical initial fuel enrichment and absorber

loadings is quantified as a function of burnup. The report concludes with a discussion of the issues for

consideration and preliminary recommendations to expand the use of burnup credit to include spent fuel

assemblies with IBAs.

v

CONTENTS

Page

ABSTRACT ............................................................................................................................................. iii

LIST OF FIGURES ................................................................................................................................. vii

LIST OF TABLES ................................................................................................................................... xi

FOREWORD .......................................................................................................................................... xiii

ACKNOWLEDGEMENTS ......................................................................................................................xv

1 INTRODUCTION ............................................................................................................................ 1

2 INTEGRAL BURNABLE ABSORBER DESIGNS CONSIDERED ................................................. 3

2.1 INTEGRAL FUEL BURNABLE ABSORBER RODS............................................................. 3

2.2 UO2-Gd2O3 INTEGRAL BURNABLE ABSORBER RODS..................................................... 3

2.3 UO2-Er2O3 INTEGRAL BURNABLE ABSORBER RODS...................................................... 4

2.4 Al2O3-B4C INTEGRAL BURNABLE ABSORBER RODS...................................................... 4

3 REACTIVITY EFFECT OF INTEGRAL BURNABLE ABSORBER RODS.................................... 5

3.1 INTRODUCTION................................................................................................................... 5

3.2 CODE DESCRIPTION ........................................................................................................... 6

3.3 CALCULATIONS................................................................................................................... 7

3.3.1 Integral Fuel Burnable Absorber Rods ......................................................................... 7

3.3.1.1 Integral Fuel Burnable Absorber Analyses for 1.57 mg 10B/inch.................... 8

3.3.1.2 Integral Fuel Burnable Absorber Analyses for 2.355 mg 10B/inch................ 17

3.3.2 UO2-Gd2O3 Integral Burnable Absorber Rods............................................................. 17

3.3.2.1 Combustion Engineering Designs ............................................................... 20

3.3.2.2 Siemens Designs ........................................................................................ 27

3.3.3 UO2-Er2O3 Integral Burnable Absorber Rods ............................................................. 36

3.3.4 Al2O3-B4C Integral Burnable Absorber Rods.............................................................. 42

3.3.5 Additional Studies ..................................................................................................... 49

3.3.5.1 Parametric Assembly Design Studies.......................................................... 49

3.3.5.1.1 Variations of gadolinia loading and initial fuel enrichment ......... 49

3.3.5.1.2 Variations of integral fuel burnable absorber loading and

initial fuel enrichment................................................................ 52

3.3.5.2 Residual Reactivity Effect of Integral Burnable Absorbers.......................... 55

3.3.5.3 Cooling Time............................................................................................. 58

3.3.5.4 Reactivity Behavior in a Poisoned Storage Cell........................................... 59

3.3.5.5 Axial Burnup............................................................................................. 62

4 SUMMARY AND IMPLICATIONS .............................................................................................. 67

5 RECOMMENDATIONS ................................................................................................................ 69

6 REFERENCES .............................................................................................................................. 71

vii

LIST OF FIGURES

Figures Page

1 Typical reactivity behavior of PWR fuel with and without IBAs (neutron poisons)................................. 6

2 HELIOS calculational model of an IFBA case containing 156 IFBA rods .............................................. 9

3 Fuel rod array (17 × 17) representing the loading pattern for 80 IFBA rods......................................... 10

4 Comparison of kinf values with IFBA rods (80 IFBA rods, 3.4 wt % 235U enriched fuel) and without

IFBA rods (3.4 wt % 235U enriched fuel) ............................................................................................. 11

5 Comparison of ∆k values as a function of burnup between assemblies with and without IFBA rods

(80 IFBA rods) for initial enrichments of 3.4 and 4.4 wt % 235U. The results are also plotted on the

enlarged scale on the right-hand-side y-axis for clarity......................................................................... 12

6 Fuel rod array (17 × 17) representing the loading pattern for 104 IFBA rods....................................... 13

7 Comparison of ∆k values as a function of burnup between assemblies with and without IFBA rods

(104 IFBA rods) for initial enrichments of 3.4 and 4.0 wt % 235U. The results are also plotted

according to the enlarged scale on the right-hand-side y-axis for clarity................................................ 14

8 Fuel rod array (17 × 17) representing the loading pattern for 156 IFBA rods....................................... 15

9 Comparison of ∆k values as a function of burnup between assemblies with and without IFBA rods

(156 IFBA rods) for initial enrichments of 4.0 and 4.4 wt % 235U. The results are also plotted on the

enlarged scale on the right-hand-side y-axis for clarity......................................................................... 16

10 Comparison of ∆k values as a function of burnup between assemblies with and without IFBA rods,

with varied number of IFBA rods and an enrichment of 4.617 wt % 235U. The results are also plotted

on the enlarged scale on the right-hand-side y-axis for clarity............................................................... 18

11 Comparison of ∆k values as a function of burnup between assemblies with and without IFBA rods

(104 IFBA rods) for initial enrichments of 4.25 and 4.617 wt % 235U. The results are also plotted

on the enlarged scale on the right-hand-side y-axis for clarity............................................................... 19

12 Fuel rod array (16 × 16) representing the loading pattern of the D and E fuel assemblies .................... 21

13 Fuel rod array (16 × 16) representing the loading pattern of the D1 and E1 fuel assemblies ................. 22

14 Fuel rod array (16 × 16) representing the loading pattern of the D2 and E2 fuel assemblies ................. 23

15 HELIOS calculational model of the D1 fuel assembly containing eight UO2-Gd2O3 rods ...................... 24
viii

LIST OF FIGURES (continued)

Figure Page

16 Comparison of ∆k values as a function of burnup between the D assembly designs with gadoliniabearing rods and the D assembly design (Figure 12) without gadolinia-bearing rods. The results

are also plotted according to the enlarged scale on the right-hand-side y-axis for clarity........................ 25

17 Comparison of ∆k values as a function of burnup between the E assembly designs with gadoliniabearing rods and the E assembly design (Figure 12) without gadolinia-bearing rods. The results

are also plotted on the enlarged scale on the right-hand-side y-axis for clarity....................................... 26

18 Fuel rod array (17 × 17) representing the M1 loading pattern.............................................................. 28

19 Fuel rod array (17 × 17) representing the M2 loading pattern.............................................................. 29

20 Fuel rod array (17 × 17) representing the M3 loading pattern.............................................................. 30

21 Fuel rod array (17 × 17) representing the M4 loading pattern.............................................................. 31

22 HELIOS calculational model of M1 fuel assembly .............................................................................. 32

23 Comparison of kinf values with and without gadolinia-bearing rods (M1 assembly).............................. 33

24 Comparison of ∆k values as a function of burnup between assemblies with and without gadoliniabearing rods (M1 and M2 assemblies). The results are also plotted on the enlarged scale on the righthand-side y-axis for clarity.................................................................................................................. 34

25 Comparison of ∆k values as a function of burnup between assemblies with (M3 and M4) and without

gadolinia-bearing rods. The results are also plotted on the enlarged scale on the right-hand-side y-axis

for clarity. .......................................................................................................................................... 35

26 HELIOS calculational model of a fuel assembly containing 60 erbia-bearing fuel rods......................... 37

27 Fuel rod array (14 × 14) representing the loading pattern for 20 UO2-Er2O3 rods................................. 38

28 Fuel rod array (14 × 14) representing the loading pattern for 60 UO2-Er2O3 rods................................. 39

29 Comparison of kinf values with erbia-bearing rods (60) present............................................................ 40

30 Comparison of ∆k values as a function of burnup between assemblies with and without UO2-Er2O3 fuel

rods. The fuel rods have a 235U enrichment of 4.3 wt %, and the UO2-Er2O3 rods contain 2.0 wt %

erbia................................................................................................................................................... 41

31 HELIOS calculational model of a fuel assembly containing 12 Al2O3-B4C rods.................................... 43
ix

LIST OF FIGURES (continued)

Figure Page

32 Fuel rod array (14 × 14) representing the loading pattern for four Al2O3-B4C rods............................... 44

33 Fuel rod array (14 × 14) representing the loading pattern for eight Al2O3-B4C rods.............................. 45

34 Fuel rod array (14 × 14) representing the loading pattern for12 Al2O3-B4C rods .................................. 46

35 Comparison of kinf values with and without Al2O3-B4C rods (12 Al2O3-B4C rods)................................ 47

36 Comparison of ∆k values as a function of burnup between assemblies with and without Al2O3-B4C

rods. The fuel rods have a 235U enrichment of 4 wt %, and the Al2O3-B4C rods have 4 wt % B4C. ....... 48

37 Comparison of ∆k values versus burnup with varying gadolinia weight percentage. The fuel

enrichments are 3.78/3.28 wt % 235U................................................................................................... 50

38 Comparison of ∆k values versus burnup with various 235U enrichments. The eight IBA rods have

6 wt % Gd2O3..................................................................................................................................... 51

39 Comparison of ∆k values versus burnup with varying IFBA loadings. The 235U enrichment is 4 wt %

for all cases. ....................................................................................................................................... 52

40 Comparison of ∆k values versus burnup with the 32 IFBA loading pattern showing varying 235U

enrichments. The results are also plotted on the enlarged scale on the right-hand-side y-axis

for clarity. .......................................................................................................................................... 53

41 Comparison of ∆k values versus burnup with the 156 IFBA loading pattern showing varying 235U

enrichments. The results are also plotted on the enlarged scale on the right-hand-side y-axis

for clarity. .......................................................................................................................................... 54

42 Reactivity worth of gadolinium isotopes versus burnup for M1 assembly............................................. 56

43 Reactivity worth of boron versus burnup for a Westinghouse 17 × 17 fuel assembly containing

64 IFBA rods ..................................................................................................................................... 57

44 Comparison of ∆k values as a function of burnup and cooling time between assemblies with and

without IFBA rods (104 IFBA rods) for and initial enrichment of 4.0 wt % 235U .................................. 58

45 Comparison of ∆k values as a function of burnup for IFBA assemblies in differing geometric

configurations..................................................................................................................................... 60
x

LIST OF FIGURES (continued)

Figure Page

46 Comparison of ∆k values as a function of burnup for assemblies with UO2-Gd2O3 rods in differing

geometric configurations..................................................................................................................... 61
xi

LIST OF TABLES

Table Page

1 Summary of parameters used for the depletion calculations ................................................................ 7

2 Westinghouse 17 × 17 fuel assembly specifications ........................................................................... 8

3 CE 16 × 16 fuel assembly specifications.......................................................................................... 20

4 Fuel assembly data for the D and E sets of assembly designs ............................................................ 20

5 Siemens 17 × 17 fuel assembly specifications .................................................................................. 27

6 Fuel assembly data for the M1–M4 assembly designs....................................................................... 27

7 CE 14 × 14 fuel assembly specifications.......................................................................................... 36

8 Summary of maximum positive ∆k values observed for IFBA cases.................................................. 55

9 Summary of maximum positive ∆k values observed for IFBA cases in a cask environment................ 61

10 Comparison of keff results in the GBC-32 cask for fuel with full-length (144 inch) IFBA rods when

the axial burnup distribution is included........................................................................................... 64

11 Comparison of keff results in the GBC-32 cask for fuel with IFBA rods when the

axial burnup distribution is included and the IFBA coating corresponds to the central 120 inches

of the axial fuel length ..................................................................................................................... 64

12 Comparison of keff results in the GBC-32 cask for fuel with IFBA rods when the

axial burnup distribution is included and the IFBA coating corresponds to the central 108 inches

of the axial fuel length ..................................................................................................................... 65

xiii

FOREWORD

In 1999 the United States Nuclear Regulatory Commission (U.S. NRC) issued initial recommended guidance

for using reactivity credit due to fuel irradiation (i.e., burnup credit) in the criticality safety analysis of spent

pressurized-water-reactor (PWR) fuel in storage and transportation packages. This guidance was issued by

the NRC Spent Fuel Project Office (SFPO) as Revision 1 to Interim Staff Guidance 8 (ISG8R1) and published

in the Standard Review Plan for Transportation Packages for Spent Nuclear Fuel, NUREG-1617 (March

2000). With this initial guidance as a basis, the NRC Office of Nuclear Regulatory Research initiated a

program to provide the SFPO with technical information that would:

• enable realistic estimates of the subcritical margin for systems with spent nuclear fuel (SNF) and an

increased understanding of the phenomena and parameters that impact the margin, and

• support the development of technical bases and recommendations for effective implementation of burnup

credit and provide realistic SNF acceptance criteria while maintaining an adequate margin of safety.

One restriction recommended by the ISG8R1 is to preclude assemblies irradiated with burnable absorbers

from being used in a cask that implements burnup credit. Since a large portion (perhaps 50% or more) of the

discharged SNF inventory has been exposed to burnable absorbers, this restriction was identified as a

potentially unnecessary and costly restriction that could be removed with improved understanding of the

impact of burnable absorbers on SNF reactivity in transportation and storage environments. Burnable

absorbers used in PWR fuel designs can be classed into two categories: burnable poison rods that can be

separated from an assembly and integral burnable absorbers (IBAs) that are a fixed part of the assembly.

This report presents a parametric study that quantifies the changes in the SNF neutron multiplication factor

due to the presence of IBAs during fuel irradiation and discusses the behavior that causes the changes. Based

on this study and discussion, the report proposes recommendations for modifying ISG8R1 to allow loading of

burnup credit casks with assemblies exposed to IBAs.

Farouk Eltawila, Director

Division of Systems Analysis and Regulatory Effectiveness

xv

ACKNOWLEDGEMENTS

This work was supported by the Office of Nuclear Regulatory Research, U.S. Nuclear Regulatory

Commission (NRC), under Project JCN W6479, Development and Applicability of Criticality Safety Software

for Licensing Review. The authors acknowledge review and useful comments by Carl J. Withee of the Spent

Fuel Project Office and Richard Y. Lee of the Office of Nuclear Regulatory Research.

The authors gratefully acknowledge Richard J. Cacciapouti (Duke Engineering and Services), Hae-Ryong

Hwang (Korea Power Engineering Company, Inc.), and John R. Massari (Constellation Nuclear Services) for

their invaluable contributions to this report by supplying fuel data and related information. The authors

acknowledge Patrick O’Leary and Michelle Pitts (Framatome ANP) for their initial review and constructive

comments on this report. The thorough review of the manuscript by our colleagues M. D. DeHart,

I. C. Gauld, C. V. Parks, and R. M. Westfall is very much appreciated. Finally, the authors are thankful to

Carolyn Moser for her editorial comments and to Willena C. Carter for her careful preparation of this report.
-
1

1 INTRODUCTION

The concept of taking credit for the reduction in reactivity due to fuel burnup is commonly referred to as

burnup credit. The reduction in reactivity that occurs with fuel burnup is due to the change in concentration

(net reduction) of fissile nuclides and the production of actinide and fission-product neutron absorbers.

The change in the concentration of these nuclides with fuel burnup, and consequently the reduction in

reactivity, is dependent upon the depletion environment (e.g., the neutron spectrum). As a result, the

utilization of credit for fuel burnup necessitates consideration of variations in fuel designs and operating

conditions.

Continuing advancements in fuel assembly design have enabled enhanced fuel utilization, thereby increasing

the performance of reactor cores (i.e., extending core lifetimes). One characteristic of these advanced fuel

assembly designs is the expanded use of burnable absorber (neutron poison) materials, either as an integral

part of the fuel assembly or as a separate assembly used in conjunction with the fuel assembly. Burnable

absorbers may be classified into two distinct categories: (1) burnable poison rods (BPRs) and (2) integral

burnable absorbers (IBAs). BPRs are rods containing neutron-absorbing material that are inserted into the

guide tubes of a pressurized water reactor (PWR) assembly during normal operation and are commonly used

for reactivity control and enhanced fuel utilization. In contrast, IBAs are burnable poisons that are a

nonremovable, or integral, part of the fuel assembly once it is manufactured. An example of an IBA is the

Westinghouse Integral Fuel Burnable Absorber (IFBA) rod, which has a coating of zirconium diboride (ZrB2)

on the fuel pellets. Integral burnable absorbers are used extensively in many current PWR fuel assembly

designs. Although BPRs have also been commonly used in PWRs, this report focuses on the effect of IBAs

only. The effect of BPRs is addressed in Ref. 1.

The Interim Staff Guidance on burnup credit2 issued by the Nuclear Regulatory Commission’s Spent Fuel

Project Office recommends that licensees restrict the use of burnup credit to assemblies that have not used

burnable absorbers. This restriction eliminates a large portion of the currently discharged spent fuel

assemblies from cask loading and thus severely limits the practical usefulness of burnup credit. Therefore, this

report examines the effect of IBAs on reactivity for various designs and enrichment/poison loading

combinations as a function of burnup. All IBA types that have been widely used in United States (U.S.)

commercial PWRs are included in this evaluation, and to the extent possible, analyses are presented for a

realistic range of initial fuel enrichment and poison loading combinations that are representative of actual

assemblies. The effects are quantified, and trends with initial fuel enrichment and poison loading are noted.

The report concludes with a discussion of the issues for consideration and preliminary recommendations for

expanding the use of burnup credit to include spent fuel assemblies with IBAs.

3

2 INTEGRAL BURNABLE ABSORBER

DESIGNS CONSIDERED

Several different types of IBAs have been used in commercial nuclear fuel assembly designs. However, all of the

various designs are similar in that they contain thermal neutron absorbing material as an integral, nonremovable

part of the fuel assembly. Variations in the IBA material, composition, placement within rods, and rod

configurations exist among current PWR fuel assembly designs. These IBA characteristics may also be varied in

combination with the initial fuel assembly enrichment and core location to achieve core operating and fuel

management goals. For completeness, this report presents analyses for all IBA types that have been widely used

in U.S. commercial PWRs. These include Westinghouse assembly designs with IFBAs, Combustion Engineering

(CE) and Siemens assembly designs with UO2-Gd2O3 rods, CE assembly designs with UO2-Er2O3 rods, and CE

assembly designs with Al2O3-B4C rods. To the extent possible, analyses were performed for a representative,

realistic range of fuel initial enrichment and poison loading combinations. We cannot, however, confirm that all

types that have ever been used in U.S. PWRs are included in this evaluation. For clarity, each of the unique IBA

types considered in this report is described below.

The fuel assembly design data used for this analysis were collected from a variety of nonproprietary sources.

For many of the IBA types, complete detailed specifications were not openly available in any single document.

Therefore, the complete fuel design specifications required for this analysis were assembled from multiple sources

and are documented in this report for reference.

2.1 INTEGRAL FUEL BURNABLE ABSORBER RODS

Some Westinghouse fuel assembly designs include IFBA rods, which contain enriched uranium dioxide (UO2) fuel

pellets with a thin coating of ZrB2 on the outer surface. To our knowledge, IFBA rods are exclusive to

Westinghouse and have been used in Westinghouse reactor cores since about 1987. Specifications of the

assembly designs that utilize IFBA rods include boron loading in the ZrB2 coating, the number of IFBA rods, and

the placement or loading pattern of the IFBA rods within the fuel assembly. The number of IFBA rods within a

fuel assembly may vary from zero to ~60% of the total number of fuel rods. For a Westinghouse 17 × 17

assembly, which contains 264 fuel rods, loading patterns with 0, 8, 16, 32, 48, 64, 80, 104, 128, and 156 IFBA

rods are known to exist. In addition, the boron loading in the ZrB2 coating and the initial 235U enrichment are

varied to meet core management goals.

2.2 UO2-Gd2O3 INTEGRAL BURNABLE ABSORBER RODS

A number of nuclear fuel vendors — including CE, Framatome ANP (formerly Babcock & Wilcox [B&W]), and

Siemens — have manufactured gadolinia-uranium (UO2-Gd2O3) IBA rods. These UO2-Gd2O3 rods, or gadolinia

rods, are fuel rods with gadolinia (Gd2O3) as an integral part of the fuel matrix and are also used extensively in

boiling water reactors (BWRs). The weight percent or loading of Gd2O3 in each gadolinia-bearing rod and the

number of gadolinia rods within an assembly are both variable. Further, the 235U enrichment among the gadoliniabearing and nongadolinia-bearing fuel rods is varied to a small extent in some designs.
IBA Designs Considered Section 2

4

2.3 UO2-Er2O3 INTEGRAL BURNABLE ABSORBER RODS

In addition to UO2-Gd2O3 rods, CE has manufactured an IBA containing erbia (Er2O3). Similar to the UO2-Gd2O3

rods, the erbia rods include the burnable absorber Er2O3 as an integral part of the fuel matrix. The weight percent

or loading of the erbia and the number of erbia rods within an assembly are both variable, as is the 235U

enrichment.

2.4 Al2O3-B4C INTEGRAL BURNABLE ABSORBER RODS

Another IBA manufactured by CE consists of solid rods containing alumina pellets with uniformly dispersed

boron carbide particles (Al2O3-B4C) clad in zircaloy. Unlike the IFBA, UO2-Gd2O3, and UO2-Er2O3 rods, these

rods do not contain fuel. The weight percent of B4C and the number of rods per assembly are variable. These

rods are similar to BPRs but are classified herein as IBAs because they are an integral, nonremoval part of the

fuel assembly.
Section 4 Summary and Implications

5

3 REACTIVITY EFFECT OF INTEGRAL

BURNABLE ABSORBER RODS

3.1 INTRODUCTION

For PWR fuels without IBAs, reactivity decreases with burnup in a nearly linear fashion. In contrast, for

PWR fuel assembly designs that make significant use of IBAs, reactivity actually increases as fuel burnup

proceeds, reaching a maximum at a burnup where the IBA is nearly depleted, and then decreasing with burnup

in an almost linear fashion. For fuel assembly designs that make modest use of IBAs, reactivity decreases with

burnup slowly, up to the point at which the IBA is nearly depleted, and then decreases with burnup in a nearly

linear manner. The assemblies are typically designed such that the burnable absorber is effectively depleted in

the first third of the assembly life, and as a result, the assembly reactivity typically peaks within this period of

burnup. The reactivity behavior of a PWR fuel assembly with and without IBAs (neutron poisons) present as

a function of burnup is illustrated in Figure 1.

The presence of IBAs during depletion hardens the neutron spectrum, resulting in lower 235U depletion and

higher production of fissile plutonium isotopes. Enhanced plutonium production and the concurrent

diminished fission of 235U due to increased plutonium fission can increase the reactivity of the fuel at discharge

and beyond, depending on the IBA assembly design characteristics. However, as mentioned, the assemblies

are typically designed such that the burnable absorber is effectively depleted in the first third of the assembly

life, and thus is exposed to a hardened spectrum during the first third of its exposure only. Unlike BPRs,

which are inserted into assembly guide tubes, IBAs do not displace the moderator in the assembly lattice, and

thus generally have a less significant impact on the neutron spectrum.

Although a great deal of work has been performed on IBA design and development for greater fuel utilization

and core performance, few studies have been undertaken to assess the significance of IBAs on the reactivity of

discharged fuel. Recent work has provided examples of the typical magnitudes of the reactivity effects of

IBAs, including IFBA, UO2-Gd2O3, and UO2-Er2O3 rods.3 Although the analyses were limited to a single case

for each type of IBA, indications from this study are that the neutron-multiplication factor for an assembly

without IBAs is always greater (as a function of burnup) than the neutron-multiplication factor for an

assembly that use UO2-Gd2O3 or UO2-Er2O3 rods. Conversely, the neutron-multiplication factor for an

assembly without IFBA rods was found to be slightly less (~0.2% ∆k at target discharge burnup) than the

neutron-multiplication factor for an assembly with IFBA rods. The study concludes that neglecting the IBAs

yields conservative results for gadolinia- and erbia-bearing fuel and nonconservative results for IFBA fuel, and

that the reactivity effect from IBAs is generally small and well behaved.

The following sections describe the calculational methods used for this evaluation and present detailed

analyses to demonstrate the reactivity effect of IBAs as a function of burnup. The analyses include variations

in the IBA type, concentration, and initial fuel enrichment.
Reactivity Effect of IBA Rods Section 3

6

Figure 1 Typical reactivity behavior of PWR fuel with and without IBAs (neutron poisons)

3.2 CODE DESCRIPTION

The calculations presented in the following sections were performed using the HELIOS-1.6 code package,4

which primarily consists of three programs: AURORA, HELIOS, and ZENITH. HELIOS is a twodimensional (2-D), generalized-geometry transport theory code based on the method of collision probabilities

with current coupling. AURORA, the input processor, is used to define the geometry, materials, and

calculational parameters. ZENITH, the output processor, reads the results saved by HELIOS (in a binary

database) and outputs the results in text format. The HELIOS code system also contains the ORION program

for viewing and checking model geometries and materials.

HELIOS was employed for this analysis because of its capability to explicitly model the relatively

complicated, heterogeneous assembly lattices associated with IBAs. The various structures within each of the

assembly models were coupled using angular current discretization (interface currents). Unless stated

otherwise, all calculations are for an infinite radial array of fuel assemblies and utilize the 45-group neutron

cross-section library (based on ENDF/B-VI data) that is distributed with the HELIOS-1.6 code package.

0.80

0.85

0.90

0.95

1.00

1.05

1.10

1.15

1.20

1.25

1.30

0 10 20 30 40 50 60

Burnup [GWd/MTU]

kinf

un-poisoned fuel

poisoned fuel
Section 3 Reactivity Effect of IBA Rods

7

3.3 CALCULATIONS

All depletion calculations were performed with HELIOS using the properties and parameters given in Table 1.

Using the isotopic compositions from the depletion calculations, branch or restart calculations were performed

with HELIOS to determine the neutron-multiplication factor as a function of burnup for out-of-reactor

conditions (i.e., 20°C with no soluble boron present), zero cooling time, and a full nuclide set. All nuclides

that are included in the HELIOS 45-group cross-section library were included in the calculations. For each

unique IBA assembly design considered, a calculation was performed for (1) the actual assembly specification

(including the presence of the IBA) and (2) the condition in which the IBA is not included (in the calculation,

IBA rods are replaced by “normal” fuel rods with the same enrichment). Throughout the following sections,

the ∆k values between these two conditions are reported to assess the effect of IBAs on the reactivity of spent

nuclear fuel (SNF).

Table 1 Summary of parameters used for the depletion calculations

Parameter Value used in analyses

Moderator temperature (K) 600

Fuel temperature (K) 1000

Fuel density (g/cm3

) 10.44 (UO2)

Clad temperature (K) 600

Clad density (g/cm3

) 5.78 (Zr)

Power density (MW/MTU) 60*

Moderator boron concentration (ppm) 650

* Various cases were also calculated using a power density of 30 MW/MTU

(which is a more realistic value). The results showed that the ∆k values presented

in the following sections are not sensitive to variations in the power density.

The majority of the calculations were done with an infinite radial array of assemblies to gain an understanding

of behaviors and trends and so that the results would be general (i.e., not dependent on storage cell

specifications, such as poison loading). A number of the criticality calculations were repeated in Section 3.3.5

to assess the effect for poisoned storage cells, variations in cooling time, and inclusion of an axial burnup

distribution. The results of these studies indicate that the calculated effects are not sensitive to cooling time

and confirm the trends demonstrated with infinite assembly arrays.

3.3.1 Integral Fuel Burnable Absorber Rods

The IFBA, developed by Westinghouse, consists of a thin coating of ZrB2 on the outer surface of the fuel

pellets. Various IFBA loading combinations (10B concentration and number of IFBA rods) and enrichment

combinations were studied in order to establish the reactivity effect as a function of burnup. The absorber

loading and enrichment combinations considered are based on actual fuel assemblies and were selected to

encompass the range of known variations including the number of IFBA rods, 10B concentration, and initial

fuel enrichment. All of the IFBA rod analyses were performed with a Westinghouse 17 × 17 assembly.

Dimensional specifications for the fuel assembly are given in Table 2.
Reactivity Effect of IBA Rods Section 3

8

Table 2 Westinghouse 17 × 17 fuel assembly specifications

Parameter

Dimension

(cm)

Rod pitch 1.260

Assembly pitch 21.5

Cladding outside diameter 0.8898

Cladding inside diameter 0.8001

Pellet outside diameter 0.7840

Guide/instrument tube outside diameter 1.204

Guide/instrument tube inside diameter 1.124

Array size 17 × 17

Number of fuel rods 264

Number of guide/instrument tubes 25

Analyses are presented in this section for IFBA assembly designs with 80, 104, and 156 IFBA rods; two

different boron loadings (1.57 and 2.355 mg 10B/in.); and corresponding variations in initial fuel enrichment.

With the exception of the loading patterns, the specifications for the actual IFBA assembly characteristics

(i.e., number of IFBA rods, 10B loading, and initial enrichment) were obtained from Ref. 5. The IFBA loading

patterns used in these analyses are from Ref. 6. Figure 2 displays the geometry of one of the IFBA assemblies

(containing 156 IFBA rods) as modeled in HELIOS.

The presentation of the IFBA analyses is divided into two subsections, based on the boron loading. The two

subsections compare results for linear poison material (boron) loadings of 1.57 and 2.355 mg 10B/in. of pellet,

respectively.

3.3.1.1 Integral Fuel Burnable Absorber Analyses for 1.57 mg 10B/inch

The first IFBA assembly design considered is a 17 × 17 assembly with 80 IFBA rods, as shown in Figure 3.

Calculations were performed for (1) the actual assembly specification (as shown in Figure 3) and (2) the

condition in which the IFBAs rods were replaced by non-IFBA fuel rods with the same enrichment as the other

fuel rods. The kinf values as a function of burnup from the two calculations are compared in Figure 4. The ∆k

values between these two conditions are determined to assess the effect of the IFBAs on reactivity.

Separate calculations were performed with enrichments of 3.4 and 4.4 wt % 235U to cover the range of initial

enrichments for this particular IFBA loading found in the available fuel data,5 which consisted of actual fuel

assembly data from the Seabrook plant. The results (∆k as a function of burnup) are shown in Figure 5, where

it can be seen that both cases achieve a positive ∆k (i.e., the kinf value with IFBA rods present becomes greater

than the kinf value without IFBA rods present). The ∆k for the case with 3.4 wt % 235U enrichment becomes

positive at a burnup of around 20 GWd/MTU, reaching a maximum of approximately 0.18% ∆k at around

34 GWd/MTU. The ∆k for the case with 4.4 wt % 235U enrichment becomes positive at a burnup of around

25 GWd/MTU, with a maximum of 0.15% ∆k near 50 GWd/MTU. Consistent with previously published

work,3 these results indicate that the IFBA-bearing fuel kinf is not always less than the non-IFBA fuel kinf.
Section 3 Reactivity Effect of IBA Rods

9

Figure 2 HELIOS calculational model of an IFBA case containing 156 IFBA rods
Reactivity Effect of IBA Rods Section 3

10

Guide/inst. tube Non-IFBA rod

IFBA rod

Figure 3 Fuel rod array (17 × 17) representing the loading pattern for 80 IFBA rods

The second IFBA design considered contains 104 IFBA rods, as shown in Figure 6. As with all designs

considered in this subsection, the poison loading is 1.57 mg 10B/in. of pellet. Once again, the initial enrichment

was varied to span the range within the actual fuel data. The results (∆k as a function of burnup) are shown in

Figure 7, where it can be seen that both cases achieve positive ∆k values. The ∆k values for the case with

3.4 wt % 235U enrichment become positive at a burnup of about 21 GWd/MTU, reach a maximum of 0.24%

∆k at around 35 GWd/MTU, and then decrease. The ∆k values for the case with 4.0 wt % 235U enrichment

become positive at a burnup of about 24 GWd/MTU, with a maximum of 0.23% ∆k at 44 GWd/MTU.

Note that the ∆k values reach a maximum and are decreasing with increased burnup beyond that point. These

results are consistent with the results shown for 80 IFBA rods, further confirming that the reactivity of an

IFBA-bearing fuel assembly is not always less than that of a non-IFBA fuel assembly.
Section 3 Reactivity Effect of IBA Rods

11

Figure 4 Comparison of kinf values with IFBA rods (80 IFBA rods, 3.4 wt %

235U enriched fuel) and without IFBA rods (3.4 wt % 235U enriched fuel)

0.75

0.80

0.85

0.90

0.95

1.00

1.05

1.10

1.15

1.20

1.25

1.30

0 10 20 30 40 50 60

Burnup [GWd/MTU]

kinf

No IFBA rods

IFBA rods
Reactivity Effect of IBA Rods Section 3

12

-0.35

-0.30

-0.25

-0.20

-0.15

-0.10

-0.05

0.00

0.05

0 10 20 30 40 50 60

Burnup [GWd/MTU]

k [k(IFBA) - k(no_IFBA)]

-0.0100

-0.0075

-0.0050

-0.0025

0.0000

0.0025

0.0050

0.0075

0.0100

k [k(IFBA) - k(no_IFBA)]

3.4 wt % U-235

4.4 wt % U-235

Series3

Figure 5 Comparison of ∆k values as a function of burnup between assemblies

with and without IFBA rods (80 IFBA rods) for initial enrichments of 3.4 and 4.4 wt %

235U. The results are also plotted on the enlarged scale on the right-hand-side y-axis for

clarity.

∆

k

∆

k
Section 3 Reactivity Effect of IBA Rods

13

Guide/inst. tube Non-IFBA rod

IFBA rod

Figure 6 Fuel rod array (17 × 17) representing the loading pattern for 104 IFBA rods
Reactivity Effect of IBA Rods Section 3

14

Figure 7 Comparison of ∆k values as a function of burnup between assemblies

with and without IFBA rods (104 IFBA rods) for initial enrichments of 3.4 and 4.0 wt %

235U. The results are also plotted according to the enlarged scale on the right-hand-side

y-axis for clarity.

The third and final IFBA design considered for this particular 10B loading contains 156 IFBA rods, as shown

in Figure 8. Based on actual fuel data, the 235U enrichment was varied between 4.0 and 4.4 wt %. The results

(∆k as a function of burnup) are shown in Figure 9. The ∆k for the case with 4.0 wt % 235U enrichment

becomes positive at a burnup of around 25 GWd/MTU, reaching a maximum of 0.31% ∆k at around

40 GWd/MTU. The ∆k for the case with 4.4 wt % 235U enrichment becomes positive at a burnup of around

27 GWd/MTU, with a maximum of 0.29 % ∆k near 50 GWd/MTU. Comparison of Figures 7 and 9 indicates

that the maximum positive ∆k value increases with the number of IFBA rods.

-0.35

-0.30

-0.25

-0.20

-0.15

-0.10

-0.05

0.00

0.05

0 10 20 30 40 50 60

Burnup [GWd/MTU]

k [k(IFBA) - k(no_IFBA)]

-0.0100

-0.0075

-0.0050

-0.0025

0.0000

0.0025

0.0050

0.0075

0.0100

k [k(IFBA) - k(no_IFBA)]

3.4 wt % U-235

4.0 wt % U-235

Series3

∆

k

∆

k
Section 3 Reactivity Effect of IBA Rods

15

Guide/inst. tube Non-IFBA rod

IFBA rod

Figure 8 Fuel rod array (17 × 17) representing the loading pattern for 156 IFBA rods
Reactivity Effect of IBA Rods Section 3

16

Figure 9 Comparison of ∆k values as a function of burnup between assemblies

with and without IFBA rods (156 IFBA rods) for initial enrichments of 4.0 and 4.4 wt %

235U. The results are also plotted on the enlarged scale on the right-hand-side y-axis for

clarity.

-0.35

-0.30

-0.25

-0.20

-0.15

-0.10

-0.05

0.00

0.05

0 10 20 30 40 50 60

Burnup [GWd/MTU]

k [k(IFBA) - k(no_IFBA)]

-0.0100

-0.0075

-0.0050

-0.0025

0.0000

0.0025

0.0050

0.0075

0.0100

k [k(IFBA) - k(no_IFBA)]

4.0 wt % U-235

4.4 wt % U-235

Series3

∆k[k(IFBA) – k(no_IFBA)]

∆k[k(IFBA) – k(no_IFBA)]
Section 3 Reactivity Effect of IBA Rods

17

3.3.1.2 Integral Fuel Burnable Absorber Analyses for 2.355 mg 10B/inch

The three IFBA loading patterns considered in the previous subsection (80, 104 and 156 IFBA rods) were also

studied with a poison loading of 2.355 mg 10B/in., for corresponding enrichments based on actual fuel data.

The ∆k values for the three IFBA rod configurations are shown in Figure 10. These results confirm that for a

given fuel enrichment, the positive reactivity effect increases with poison loading (i.e., number of IFBA rods).

An assembly design containing 104 IFBA rods was also evaluated with an enrichment of 4.25 wt % 235U.

The ∆k values are shown in Figure 11, along with the reactivity differences from the previous calculation with

104 IFBA rods and 4.617 wt % 235U enrichment. The results indicate that the maximum positive ∆k value

increases slightly with decreasing initial fuel enrichment.

These results are consistent with the results shown in the previous subsection, demonstrating that the kinf

values for fuel containing IFBA rods are not always less than the kinf values for fuel without IFBA rods. For

burnups greater than ~20 GWd/MTU, the IFBA rods are shown to have a small positive reactivity effect as

compared to assemblies without IFBA rods. This transition point moves up with increasing fuel enrichment

and increasing number of IFBA rods. For all of the IFBA cases considered, the maximum positive reactivity

effect was determined to be 0.41%. However, it should be pointed out that this maximum value corresponds

to the most heavily poisoned assembly design (156 IFBA rods with 2.355 mg 10B/in.) and, judging from the

available data, is not representative of typical IFBA assemblies.

For fixed initial fuel enrichment, the positive reactivity effect has been shown to increase with increasing

poison loading (i.e., both increasing 10B loading and increasing number of IFBA rods). Finally, the results in

this and the previous subsection make clear that for a fixed poison loading the positive reactivity effect

increases slightly with decreasing fuel enrichment.

3.3.2 UO2-Gd2O3 Integral Burnable Absorber Rods

A number of nuclear fuel vendors — including CE, Framatome ANP (formerly B&W), and Siemens — have

manufactured gadolinia-uranium (UO2-Gd2O3) IBA rods. These UO2-Gd2O3 rods, or gadolinia rods, are fuel

rods with Gd2O3 as an integral part of the fuel matrix. The weight percent or loading of Gd2O3 in each

gadolinia-bearing rod and the number of gadolinia rods within an assembly are both variable. Further, the 235U

enrichment among the gadolinia-bearing and nongadolinia-bearing fuel rods is varied to a small extent in some

designs, with the gadolinia-bearing fuel rods typically having lower 235U enrichment than the nongadoliniabearing fuel rods. (Because of the lower heat conductivity of the UO2-Gd2O3 fuel as compared to UO2 fuel the

235U enrichment in the UO2-Gd2O3 fuel rod is often reduced to meet the design criterion for maximum fuel

temperature.7

) A known inherent penalty associated with gadolinia rods is that gadolinia displaces uranium in

the fuel matrix, resulting in a reduced heavy-metal mass.

Various gadolinia loadings (wt % Gd2O3 and number of gadolinia-bearing rods) and enrichment combinations

were studied in order to establish the reactivity effect as a function of burnup. The absorber loading and

enrichment combinations considered are based on actual fuel assemblies and were selected to encompass the

range of known variations.

Analyses are presented in the following subsections for two distinct fuel assembly designs that employ

gadolinia-bearing rods: the CE 16 × 16 assembly design (which includes large water holes) and the Siemens

17 × 17 design. The Siemens assembly design does not include oversized water holes, and thus is expected
Reactivity Effect of IBA Rods Section 3

18

-0.35

-0.30

-0.25

-0.20

-0.15

-0.10

-0.05

0.00

0.05

0 10 20 30 40 50 60

Burnup [GWd/MTU]

∆k [k(IFBA) - k(no_IFBA)]

-0.0100

-0.0075

-0.0050

-0.0025

0.0000

0.0025

0.0050

0.0075

0.0100

∆k [k(IFBA) - k(no_IFBA)]

80 IFBA rods

104 IFBA rods

156 IFBA rods

Series4

Figure 10 Comparison of ∆k values as a function of burnup between assemblies

with and without IFBA rods, with varied number of IFBA rods and an enrichment of

4.617 wt % 235U. The results are also plotted on the enlarged scale on the right-handside y-axis for clarity.
Section 3 Reactivity Effect of IBA Rods

19

Figure 11 Comparison of ∆k values as a function of burnup between assemblies

with and without IFBA rods (104 IFBA rods) for initial enrichments of 4.25 and

4.617 wt % 235U. The results are also plotted on the enlarged scale on the right-handside y-axis for clarity.

-0.35

-0.30

-0.25

-0.20

-0.15

-0.10

-0.05

0.00

0.05

0 10 20 30 40 50 60

Burnup [GWd/MTU]

k [k(IFBA) - k(no_IFBA)]

-0.0100

-0.0075

-0.0050

-0.0025

0.0000

0.0025

0.0050

0.0075

0.0100

k [k(IFBA) - k(no_IFBA)]

4.25 wt % U-235

4.617 wt % U-235

Series3

∆k[k(IFBA) – k(no_IFBA)]

∆k[k(IFBA) – k(no_IFBA)]
Reactivity Effect of IBA Rods Section 3

20

to be representative of other similar fuel assembly designs (e.g., Framatome ANP and Westinghouse designs

that employ gadolinia-bearing rods).

3.3.2.1 Combustion Engineering Designs

The specifications for the CE gadolinia-bearing fuel assembly designs were obtained from Ref. 8. The fuel

dimension specifications for these CE 16 × 16 assemblies are listed in Table 3. To facilitate the discussion,

the assembly designs are classified into two sets, based on their initial pin enrichments. The assembly designs

within the two designated sets, the D sets (D, D1 and D2) and the E sets (E, E1 and E2), have variable

numbers of gadolinia-bearing rods, as defined in Table 4. The corresponding fuel loading diagrams are

illustrated in Figures 12 through 14. Figure 15 displays the D1 geometry as modeled in HELIOS.

Table 3 CE 16 × 16 fuel assembly specifications

Parameter

Dimension

(cm)

Rod pitch 1.285

Assembly pitch 20.78

Cladding outside diameter 0.97028

Cladding inside diameter 0.84328

Pellet outside diameter 0.82550

Water hole outside diameter 2.4892

Water hole inside diameter 2.286

Array size 16 × 16

Number of fuel rods 236

Number of water holes 5

Table 4 Fuel assembly data for the D and E sets of assembly designs

Fuel

assembly

UO2 fuel rod

enrichment

No. of UO2 fuel

rods/assembly

No. of

UO2-Gd2O3

rods/assembly

Gd2O3 /

235U wt %

for UO2-Gd2O3 rods

D 4.42/3.92 184/52 0 N/A

D1 4.42/3.92 176/52 8 6.0/4.42*

D2 4.42/3.92 172/52 12 6.0/4.42

E 4.60/4.10 184/52 0 N/A

E1 4.60/4.10 176/52 8 8.0/4.60

E2 4.60/4.10 172/52 12 8.0/4.60

*Read as 6.0 wt % Gd2O3 and 4.42 wt % 235U enrichment in UO2-Gd2O3 rods.
Section 3 Reactivity Effect of IBA Rods

21

Water hole Higher enr. fuel rod

Lower enr. fuel rod

Figure 12 Fuel rod array (16 × 16) representing the loading pattern of the D and E fuel

assemblies
Reactivity Effect of IBA Rods Section 3

22

Water hole Lower enr. fuel rod

Gd-bearing fuel rod

Higher enr. fuel rod

Figure 13 Fuel rod array (16 × 16) representing the loading pattern of the D1

and E1 fuel assemblies
Section 3 Reactivity Effect of IBA Rods

23

Water hole Lower enr. fuel rod

Gd-bearing fuel rod

Higher enr. fuel rod

Figure 14 Fuel rod array (16 × 16) representing the loading pattern of the D2

and E2 fuel assemblies
Reactivity Effect of IBA Rods Section 3

24

Figure 15 HELIOS calculational model of the D1 fuel assembly containing eight UO2-Gd2O3 rods
Section 3 Reactivity Effect of IBA Rods

25

The D and E fuel designs were used to study the effect of various numbers of gadolinia-bearing rods. The D,

D1, and D2 fuel assemblies were modeled according to Tables 3 and 4. The resulting ∆k values between the

gadolinia-bearing (D1 and D2) and nongadolinia-bearing fuel (D) as a function of burnup are shown in

Figure 16. The ∆k values do not become positive for either of the gadolinia-bearing fuel assembly designs.

In other words, the kinf values for the gadolinia-bearing fuel remain less than the kinf values for the

nongadolinia fuel. Also note that the D2 fuel assembly, with 12 gadolinia-bearing rods, has slightly more

negative ∆k values than the D1 fuel assembly, which has only 8 gadolinia-bearing rods.

-0.10

-0.09

-0.08

-0.07

-0.06

-0.05

-0.04

-0.03

-0.02

-0.01

0.00

0 10 20 30 40 50 60

Burnup [GWd/MTU]

∆k [k(Gd2O3) - k(no_Gd2O3)]

-0.010

-0.009

-0.008

-0.007

-0.006

-0.005

-0.004

-0.003

-0.002

-0.001

0.000

∆k [k(Gd2O3) - k(no_Gd2O3)]

8 Gd rods - D1

12 Gd rods - D2

Series3

Figure 16 Comparison of ∆k values as a function of burnup between the

D assembly designs with gadolinia-bearing rods and the D assembly design (Figure 12)

without gadolinia-bearing rods. The results are also plotted according to the enlarged

scale on the right-hand-side y-axis for clarity.
Reactivity Effect of IBA Rods Section 3

26

The effect of the gadolinia-bearing rods was also studied for the E fuel assembly designs, modeled according

to Tables 3 and 4. The resulting ∆k values as a function of burnup are shown in Figure 17. Again, the ∆k

values for both of the gadolinia-bearing fuel assemblies (E1 and E2) remain negative throughout the entire

burnup period. Also note that ∆k values are slightly more negative for the higher 235U enriched/higher

gadolinia-loading E fuel assembly designs than for the lower 235U enriched/lower gadolinia-loading D fuel

assembly designs.

The results from both the D and E fuel assembly designs indicate that the gadolinia-bearing fuel kinf values are

less than the non-gadolinia-bearing fuel kinf values throughout burnup. The extent by which the reactivity of

the gadolinia-bearing fuel assembly is reduced increases with increasing gadolinia loading (wt % Gd2O3 and

the number of gadolinia-bearing rods).

Figure 17 Comparison of ∆k values as a function of burnup between the

E assembly designs with gadolinia-bearing rods and the E assembly design (Figure 12)

without gadolinia-bearing rods. The results are also plotted on the enlarged scale on the

right-hand-side y-axis for clarity.

-0.10

-0.09

-0.08

-0.07

-0.06

-0.05

-0.04

-0.03

-0.02

-0.01

0.00

0 10 20 30 40 50 60

Burnup [GWd/MTU]

k [k(Gd2O3) - k(no_Gd2O3)]

-0.010

-0.009

-0.008

-0.007

-0.006

-0.005

-0.004

-0.003

-0.002

-0.001

0.000

k [k(Gd2O3) - k(no_Gd2O3)]

8 Gd rods - E1

12 Gd rods - E2

Series3

∆

k

∆

k
Section 3 Reactivity Effect of IBA Rods

27

3.3.2.2 Siemens Designs

In addition to the CE assembly designs analyzed in the previous subsection, Siemens assembly designs with

gadolinia-bearing rods were analyzed. Four different 17 × 17 fuel assembly designs with various gadolinia

loadings were considered. The common assembly specifications are listed in Table 5.

Table 5 Siemens 17 × 17 fuel assembly specifications

Parameter

Dimension

(cm)

Rod pitch 1.260

Assembly pitch 21.5

Cladding outside diameter 0.95504

Cladding inside diameter 0.83312

Pellet outside diameter 0.81661

Guide/instrument tube outside diameter 1.2192

Guide/instrument tube inside diameter 1.1379

Array size 17 × 17

Number of fuel rods 264

Number of guide/instrument tubes 25

The assembly designs considered (designated M1–M4) are summarized in Table 6 and illustrated in

Figures 18–21. Unpoisoned, equivalent enrichment reference cases (corresponding to M1–M4, respectively)

were also analyzed and used for comparison. These unpoisoned cases were exactly the same as the actual

cases, with the exception that the UO2-Gd2O3 rods were replaced with equivalent enrichment UO2 rods.

The HELIOS calculational model for the M1 assembly design is shown in Figure 22.

Table 6 Fuel assembly data for the M1–M4 assembly designs

Fuel

assembly

UO2 fuel rod

enrichment

No. of UO2 fuel

rods/assembly

No. of UO2-Gd2O3

rods/assembly

Gd2O3 /

235U wt % for

UO2-Gd2O3 rods

M1 4.25 236 16

12

8.0/3.91*

4.0/4.08

M2 4.25 240 16

8

8.0/3.91

4.0/4.08

M3 4.25 244 16

4

6.0/3.99

2.0/4.16

M4 4.25 260 4 2.0/4.16

*Read as 8.0 wt % Gd2O3 and 3.91 wt % 235U enrichment in UO2-Gd2O3 rods.
Reactivity Effect of IBA Rods Section 3

28

4 4

8 8

8

8 8

4 4

8 4 8

8 4 4 8

8 4 8

4 4

8 8

8

8 8

4 4

Guide/inst. tube Fuel rod

4 Gd-bearing fuel rod with 4 wt % Gd2O3

8 Gd-bearing fuel rod with 8 wt % Gd2O3

Figure 18 Fuel rod array (17 × 17) representing the M1 loading pattern
Section 3 Reactivity Effect of IBA Rods

29

8 8

4 4

8 8 8 8

8 8

4 4

4 4

8 8

8 8 8 8

4 4

8 8

Guide/inst. tube Fuel rod

4 Gd-bearing fuel rod with 4 wt % Gd2O3

8 Gd-bearing fuel rod with 8 wt % Gd2O3

Figure 19 Fuel rod array (17 × 17) representing the M2 loading pattern
Reactivity Effect of IBA Rods Section 3

30

6 6

6 6

6 6

2 2

6 6

6 6

2 2

6 6

6 6

6 6

Guide/inst. tube Fuel rod

2 Gd-bearing fuel rod with 2 wt % Gd2O3

6 Gd-bearing fuel rod with 6 wt % Gd2O3

Figure 20 Fuel rod array (17 × 17) representing the M3 loading pattern
Section 3 Reactivity Effect of IBA Rods

31

2 2

2 2

Guide/inst. tube Fuel rod

2 Gd-bearing fuel rod with 2 wt % Gd2O3

Figure 21 Fuel rod array (17 × 17) representing the M4 loading pattern
Reactivity Effect of IBA Rods Section 3

32

Figure 22 HELIOS calculational model of M1 fuel assembly
Section 3 Reactivity Effect of IBA Rods

33

The kinf values as a function of burnup for the unpoisoned reference case (corresponding to M1) and the

poisoned M1 case are compared in Figure 23. The ∆k values as a function of burnup for the M1–M4

assemblies are shown in Figures 24 and 25, where it can be seen that all of the gadolinia-bearing fuel assembly

designs yield negative ∆k values. As mentioned earlier, each of the four cases has a separate reference case

that does not contain any gadolinia.

Figure 23 Comparison of kinf values with and without gadolinia-bearing rods

(M1 assembly)

These results are consistent with the results shown in the previous subsection, confirming that the kinf values

for fuel assemblies with gadolinia-bearing rods are always less than the kinf values for fuel assemblies without

gadolinia-bearing rods. Further, the amount by which the kinf values are lower increases with increasing

gadolinia loading. The reason for this negative reactivity with gadolinia-bearing rods is discussed in

Section 3.3.5 (Subsection 3.3.5.2).

0.80

0.85

0.90

0.95

1.00

1.05

1.10

1.15

1.20

1.25

1.30

0 10 20 30 40 50 60

Burnup [GWd/MTU]

kinf

No Gd-bearing rods

Gd-bearing rods
Reactivity Effect of IBA Rods Section 3

34

Figure 24 Comparison of ∆k values as a function of burnup between assemblies

with and without gadolinia-bearing rods (M1 and M2 assemblies). The results are also

plotted on the enlarged scale on the right-hand-side y-axis for clarity.

-0.25

-0.20

-0.15

-0.10

-0.05

0.00

0 10 20 30 40 50 60

Burnup [GWd/MTU]

k [k(Gd2O3) - k(no_Gd2O3)]

-0.015

-0.012

-0.009

-0.006

-0.003

0.000

k [k(Gd2O3) - k(no_Gd2O3)]

M1

M2

Series3

∆

k

∆

k
Section 3 Reactivity Effect of IBA Rods

35

Figure 25 Comparison of ∆k values as a function of burnup between assemblies

with (M3 and M4) and without gadolinia-bearing rods. The results are also plotted on

the enlarged scale on the right-hand-side y-axis for clarity.

-0.35

-0.30

-0.25

-0.20

-0.15

-0.10

-0.05

0.00

0.05

0 10 20 30 40 50 60

Burnup [GWd/MTU]

k [k(Gd2O3) - k(no_Gd2O3)]

-0.0150

-0.0125

-0.0100

-0.0075

-0.0050

-0.0025

0.0000

0.0025

0.0050

k [k(Gd2O3) - k(no_Gd2O3)]

M3

M4

Series3

∆

k

∆

k
Reactivity Effect of IBA Rods Section 3

36

3.3.3 UO2-Er2O3 Integral Burnable Absorber Rods

In addition to UO2-Gd2O3 rods, CE has manufactured an IBA containing erbia (Er2O3). Assemblies with

erbia-bearing rods are far less prevalent than assemblies with gadolinia-bearing rods. Similar to the

UO2-Gd2O3 rods, the erbia-bearing rods include the burnable absorber (Er2O3) as an integral part of the fuel

matrix. The weight percent of the erbia and the number of erbia-bearing rods within an assembly are both

variable, as well as the 235U enrichment.

Common specifications for the CE 14 × 14 erbia-bearing fuel assembly design9 are listed in Table 7.

Figure 26 shows the geometry of one of the fuel assemblies considered (containing 60 erbia rods). The two

unique erbia-bearing fuel assembly lattices considered in this analysis are shown in Figures 27 and 28.

The enrichment of all rods is 4.3 wt % 235U, and the weight percent of Er2O3 in the erbia-bearing rods is 2.0.

Calculations were performed for (1) the actual assembly specifications (as shown in Figures 27 and 28) and

(2) a case in which the UO2-Er2O3 rods were replaced with equivalent enrichment (4.3 wt % 235U) UO2 rods.

The kinf values as a function of burnup for the two conditions are compared in Figure 29. The results (∆k

values) from the calculations are shown in Figure 30. It can be seen that the kinf values for the non-erbiabearing fuel assembly remain higher than the kinf values corresponding to the erbia-bearing fuel assemblies

(i.e., the ∆k values are always negative). Further, the ∆k values become more negative with increasing erbia

loading.

Table 7 CE 14 × 14 fuel assembly specifications

Parameter

Dimension

(cm)

Rod pitch 1.47

Assembly pitch 20.8

Cladding outside diameter 1.1176

Cladding inside diameter 0.97536

Pellet outside diameter 0.95631

Water hole outside diameter 2.4079

Water hole inside diameter 2.3063

Array size 14 × 14

Number of fuel rods 176

Number of water holes 5
Section 3 Reactivity Effect of IBA Rods

37

Figure 26 HELIOS calculational model of a fuel assembly containing 60 erbia-bearing fuel rods
Reactivity Effect of IBA Rods Section 3

38

Water hole Fuel rod

Er-bearing fuel rod

Figure 27 Fuel rod array (14 × 14) representing the loading pattern for 20 UO2-Er2O3 rods
Section 3 Reactivity Effect of IBA Rods

39

Water hole Fuel rod

Er-bearing fuel rod

Figure 28 Fuel rod array (14 × 14) representing the loading pattern for 60 UO2-Er2O3 rods
Reactivity Effect of IBA Rods Section 3

40

Figure 29 Comparison of kinf values with erbia-bearing rods (60) present

0.80

0.85

0.90

0.95

1.00

1.05

1.10

1.15

1.20

1.25

1.30

0 10 20 30 40 50 60

Burnup [GWd/MTU]

kinf

No Er-bearing rods

Er-bearing rods
Section 3 Reactivity Effect of IBA Rods

41

Figure 30 Comparison of ∆k values as a function of burnup between assemblies

with and without UO2-Er2O3 fuel rods. The fuel rods have a 235U enrichment of

4.3 wt %, and the UO2-Er2O3 rods contain 2.0 wt % erbia.

-0.16

-0.14

-0.12

-0.10

-0.08

-0.06

-0.04

-0.02

0.00

0 10 20 30 40 50 60

Burnup [GWd/MTU]

∆ k [k (Er2O 3) - k (n o_Er2O 3)]

20 Erbia rods

60 Erbia rods
Reactivity Effect of IBA Rods Section 3

42

3.3.4 Al2O3-B4C Integral Burnable Absorber Rods

Another IBA manufactured by CE consists of solid rods containing aluminum pellets with uniformly dispersed

boron carbide particles (Al2O3-B4C), clad in zircaloy. Unlike the IFBA, UO2-Gd2O3, and UO2-Er2O3 rods,

these rods do not contain fuel and hence are often referred to as burnable poison rods. However, because the

B4C rods are an integral, nonremovable part of the fuel assembly, they are included in this study. The weight

percent of B4C and the number of Al2O3-B4C rods per assembly are variable.

The limited specifications for actual CE fuel assembly designs with Al2O3-B4C rods were obtained from

Ref. 9. Figure 31 displays the geometry of one of the CE fuel assembly designs considered, containing

12 Al2O3-B4C rods. The fuel dimension specifications for these CE 14 × 14 assemblies are listed in Table 7,

and the assembly lattices are shown in Figures 32–34. The Al2O3-B4C rods contain 4.0 wt % B4C and the

UO2 rods have an initial enrichment of 4.0 wt % 235U.

For the unpoisoned reference cases, the Al2O3-B4C rods were replaced with equivalent enrichment (4.0 wt %

235U) fuel rods. The kinf values as a function of burnup with and without Al2O3-B4C rods present are compared

in Figure 35. The ∆k values as a function of burnup are shown in Figure 36, where it can be seen that the

cases with Al2O3-B4C rods maintain negative ∆k values throughout the entire burnup period.
Section 3 Reactivity Effect of IBA Rods

43

Figure 31 HELIOS calculational model of a fuel assembly containing 12 Al2O3-B4C rods
Reactivity Effect of IBA Rods Section 3

44

Water hole Fuel rod

Al2O3-B4C rod

Figure 32 Fuel rod array (14 × 14) representing the loading pattern for four Al2O3-B4C rods
Section 3 Reactivity Effect of IBA Rods

45

Water hole Fuel rod

Al2O3-B4C rod

Figure 33 Fuel rod array (14 × 14) representing the loading pattern for eight Al2O3-B4C rods
Reactivity Effect of IBA Rods Section 3

46

Water hole Fuel rod

Al2O3-B4C rod

Figure 34 Fuel rod array (14 × 14) representing the loading pattern for12 Al2O3-B4C rods
Section 3 Reactivity Effect of IBA Rods

47

Figure 35 Comparison of kinf values with and without Al2O3-B4C rods (12 Al2O3-B4C rods)

0.75

0.80

0.85

0.90

0.95

1.00

1.05

1.10

1.15

1.20

1.25

1.30

0 10 20 30 40 50 60

Burnup [GWd/MTU]

kinf

Without Al2O3-B4C rods

With Al2O3-B4C rods
Reactivity Effect of IBA Rods Section 3

48

Figure 36 Comparison of ∆k values as a function of burnup between assemblies

with and without Al2O3-B4C rods. The fuel rods have a 235U enrichment of 4 wt %, and the

Al2O3-B4C rods have 4 wt % B4C.

-0.20

-0.18

-0.16

-0.14

-0.12

-0.10

-0.08

-0.06

-0.04

-0.02

0.00

0 5 10 15 20 25 30 35 40 45 50 55 60

Burnup [GWd/MTU]

k [k(Al2O3-B4C) - k(no_Al2O3-B4C)]

4 Al2O3-B4C rods

8 Al2O3-B4C rods

12 Al2O3-B4C rods

∆

k
Section 3 Reactivity Effect of IBA Rods

49

3.3.5 Additional Studies

Because this study was performed in support of burnup credit, several of the calculations presented in the

previous sections were repeated with modeling assumptions and conditions common to burnup credit analyses

to assess their impact on the results. This section presents infinite assembly array calculations to study the

effect of variations in assembly designs (i.e., poison loading and initial fuel enrichment) and variations in

cooling time, and to investigate the reactivity worth of residual IBA materials. The section also presents

studies of reactivity behavior in a poisoned storage cell and the effect of including an axial burnup distribution

in a cask model.

3.3.5.1 Parametric Assembly Design Studies

All of the analyses presented in the previous sections are based on actual fuel assembly designs obtained from

plant data. The additional studies presented in this section are an attempt to better understand the impact of

IBA rods for variations outside those shown in the previous sections. These studies are intended to clarify

trends and/or relationships between the variable characteristics (e.g., poison loading and initial fuel

enrichment). Note that although these studies are based on actual fuel assembly designs, the parameters are

varied outside of the known range. Consequently, the results in this section should be used to enhance

understanding, and not for quantifying reactivity effects.

3.3.5.1.1 Variations of gadolinia loading and initial fuel enrichment

The first additional study involves variations in poison loading and fuel enrichment in a CE fuel assembly

design with UO2-Gd2O3 rods. The reference case consists of the 16 × 16 fuel assembly with no burnable

absorber rods, as shown in Figure 12. The dimensions and specification of the fuel assembly are provided in

Tables 3 and 4. For the cases with IBA rods present, the assembly layout is as shown in Figure 13.

Calculations were performed using varying weight percentages of gadolinia in the UO2-Gd2O3 rods to

investigate the sensitivity to gadolinia loading. Gadolinia enrichments of 4, 6, and 8 wt % were considered,

while the 235U enrichment was fixed at 3.78/3.28. The ∆k values for the various gadolinia enrichment cases as

function of burnup are shown in Figure 37. These results are consistent with earlier results for actual fuel

assembly designs, showing that the ∆k values become more negative with increasing gadolinia loading.

The calculational model with 6 wt % gadolinia in the UO2-Gd2O3 rods was further investigated by considering

cases with uniform fuel rod enrichments of 3, 4, and 5 wt % 235U. The ∆k values for the various enrichments

are shown in Figure 38. All of the cases maintain negative ∆k values, and the initial 235U enrichment is shown

to have a relatively small effect.
Reactivity Effect of IBA Rods Section 3

50

Figure 37 Comparison of ∆k values versus burnup with varying gadolinia

weight percentage. The fuel enrichments are 3.78/3.28 wt % 235U.

-0.040

-0.035

-0.030

-0.025

-0.020

-0.015

-0.010

-0.005

0.000

0 5 10 15 20 25 30 35 40 45 50 55 60

Burnup [GWd/MTU]

k [k(Gd2O3) - k(no_Gd2O3)]

4 w t % Gd2O3

6 w t % Gd2O3

8 w t % Gd2O3

∆

k
Section 3 Reactivity Effect of IBA Rods

51

Figure 38 Comparison of ∆k values versus burnup with various 235U enrichments.

The eight IBA rods have 6 wt % Gd2O3.

-0.040

-0.035

-0.030

-0.025

-0.020

-0.015

-0.010

-0.005

0.000

0 5 10 15 20 25 30 35 40 45 50 55 60

Burnup [GWd/MTU]

k [k(Gd2O3) - k(no_Gd2O3)]

3 w t % U-235

4 w t % U-235

5 w t % U-235

∆k[k(Gd2O3) – k(no_Gd2O3)]
Reactivity Effect of IBA Rods Section 3

52

3.3.5.1.2 Variations of integral fuel burnable absorber loading and initial fuel enrichment

The second additional study involves variations in IFBA loading and fuel enrichment in a 17 × 17 assembly.

The dimensions and fuel assembly specifications can be found in Table 2. Analyses were performed for IFBA

assembly designs10 with 32, 64, 80, 104, 128, and 156 IFBA rods and 235U enrichments of 3.0, 4.0, and

5.0 wt % with a boron loading of 1.57 mg 10B/in. The ∆k values for the various IFBA loading patterns as a

function of burnup for initial 235U enrichment of 4.0 wt % are shown in Figure 39. These results are consistent

with those shown in Section 3.3.1 because the maximum positive ∆k values increase as the number of IFBA

rods increase.

Figure 39 Comparison of ∆k values versus burnup with varying IFBA loadings.

The 235U enrichment is 4 wt % for all cases.

To illustrate the effect of initial fuel enrichment, ∆k values for 3.0, 4.0, and 5.0 wt % 235U are shown in

Figures 40 and 41 for IFBA loading patterns of 32 and 156, respectively. As is consistent with the results

shown in Section 3.3.1, the positive ∆k values increase with decreasing enrichment. Note that the 3.0 and

4.0 wt % 235U cases reach a maximum ∆k and then begin to decrease. The maximum positive ∆k values for

the various IFBA cases are summarized in Table 8.

-0.010

-0.008

-0.006

-0.004

-0.002

0.000

0.002

0.004

0 10 20 30 40 50 60

Burnup [GWd/MTU]

∆k [k(IFBA) - k(no_IFBA)]

32 IFBA

64 IFBA

80 IFBA

104 IFBA

128 IFBA

156 IFBA
Section 3 Reactivity Effect of IBA Rods

53

-0.05

-0.04

-0.03

-0.02

-0.01

0.00

0.01

0 10 20 30 40 50 60

Burnup [GWd/MTU]

∆k [k(IFBA) - k(no_IFBA)]

-0.0020

-0.0013

-0.0007

0.0000

0.0007

0.0013

0.0020

∆k [k(IFBA) - k(no_IFBA)]

3 wt % U-235

4 wt % U-235

5 wt % U-235

Series4

Figure 40 Comparison of ∆k values versus burnup with the 32 IFBA loading

pattern showing varying 235U enrichments. The results are also plotted on the enlarged

scale on the right-hand-side y-axis for clarity.

∆

k

∆

k
Reactivity Effect of IBA Rods Section 3

54

Figure 41 Comparison of ∆k values versus burnup with the 156 IFBA loading

pattern showing varying 235U enrichments. The results are also plotted on the enlarged

scale on the right-hand-side y-axis for clarity.

-0.14

-0.13

-0.11

-0.10

-0.08

-0.07

-0.05

-0.04

-0.02

-0.01

0.01

0 10 20 30 40 50 60

Burnup [GWd/MTU]

k [k(IFBA) - k(no_IFBA)]

-0.005

-0.004

-0.003

-0.002

-0.001

0.000

0.001

0.002

0.003

0.004

0.005

k [k(IFBA) - k(no_IFBA)]

3 wt % U-235

4 wt % U-235

5 wt % U-235

Series4

∆

k

∆

k
Section 3 Reactivity Effect of IBA Rods

55

Table 8 Summary of maximum positive ∆k values observed for IFBA cases

Enrichment

(wt % 235U)

IFBA loading pattern 3.0 4.0 5.0

32 0.0009 0.0008 0.0007

64 0.0018 0.0015 0.0013

80 0.0020 0.0016 0.0014

104 0.0026 0.0021 0.0018

128 0.0031 0.0026 0.0022

156 0.0038 0.0031 0.0027

3.3.5.2 Residual Reactivity Effect of Integral Burnable Absorbers

Sections 3.3.1 and 3.3.2 demonstrate that the ∆k values become positive for fuel assembly designs containing

IFBA rods but remain negative for gadolinia-bearing fuel assembly designs. To understand this different

behavior, a study looked at the residual effects of the gadolinium in the UO2-Gd2O3 rods, as well as the boron

poison in the IFBA rods. The analyses for the residual effects of gadolinium were performed for the Siemens

17 × 17 fuel design referred to as M1 (see Table 5 for fuel assembly data).

In order to observe the residual effects of gadolinium, the reactivity worth (∆k/k) was calculated by artificially

setting various compositions to zero in the criticality calculation model. Two additional calculations were

performed for comparison to the unaltered condition. The first calculation featured the M1 fuel design with

the non-fission product (FP) gadolinium artificially set to zero. The second calculation, performed for the M1

fuel design, artificially set the non-FP minor gadolinium isotopes (152Gd, 154Gd, 156Gd, 158Gd, and 160Gd) to

zero. The results of these calculations and the results from the unaltered condition were used to calculate the

reactivity worth of the two sets of gadolinium isotopes, which are displayed in Figure 42 as a function of

burnup. It can be seen that there is a negative residual effect associated with the presence of the gadolinium

isotopes.

Analyses were also performed for the Westinghouse IFBA design containing 64 IFBA rods, 1.57 10B/in.

boron loading, and 4.0 wt % initial fuel enrichment to study the residual effects of 10B and 11B. The reactivity

worth was calculated by setting 10B and 11B to zero during the calculation. The results, plotted versus burnup

in Figure 43, demonstrate that there are no significant residual reactivity effects with IFBA rods.

These analyses show that there is a negative residual effect for gadolinia-bearing fuel but no such effect for

fuel designs with IFBA rods. The main difference between the UO2-Gd2O3 rods and the IFBA rods is that

gadolinium is an integral part of the fuel matrix in the UO2-Gd2O3 rod while the boron is placed as a thin

coating on the outer surface of the fuel pellets in the IFBA rod. The gadolinium isotopes displace uranium in

the UO2-Gd2O3 rod, resulting in reduced reactivity (due to the reduction in heavy metal mass). The IFBA

coating does not displace uranium, and thus, there is no negative residual effect.
Reactivity Effect of IBA Rods Section 3

56

Figure 42 Reactivity worth of gadolinium isotopes versus burnup for M1 assembly

-0.20

-0.18

-0.16

-0.14

-0.12

-0.10

-0.08

-0.06

-0.04

-0.02

0.00

0 10 20 30 40 50 60

Burnup [GWd/MTU]

k/k

reactivity worth of minor Gd isotopes

(Gd-152, Gd-154, Gd-156, Gd-158, Gd-160)

reactivity worth of Gd absorber

(all non-fission-product Gd)
Section 3 Reactivity Effect of IBA Rods

57

3.3.5.3

3.3.5.4

3.3.5.5

3.3.5.6

3.3.5.7

3.3.5.8

3.3.5.9

3.3.5.10

3.3.5.11

Figure 43 Reactivity worth of boron versus burnup for a Westinghouse 17 × 17 fuel assembly

containing 64 IFBA rods

-0.08

-0.07

-0.06

-0.05

-0.04

-0.03

-0.02

-0.01

0.00

0 10 20 30 40 50 60

Burnup [GWd/MTU]

k/k

reactivity worth of B-10

reactivity worth of B-10 and B-11
Reactivity Effect of IBA Rods Section 3

58

3.3.5.3 Cooling Time

To study the effects of cooling time, a few calculations were performed at cooling times other than zero.

The fuel assembly employed for the study was the design containing 104 IFBA rods with a poison loading of

1.57 mg 10B/in. and 4.0 wt % 235U enrichment. Infinite assembly array calculations were performed for

cooling times of 5, 20, and 40 years. The results (∆k as a function of burnup) are displayed in Figure 44.

It can be seen that cooling time has little impact on the ∆k values, and thus, the results at zero cooling time are

expected to be applicable within the time frame relevant to cask storage and transportation.

Figure 44 Comparison of ∆k values as a function of burnup and cooling time between

assemblies with and without IFBA rods (104 IFBA rods) for and initial enrichment of

4.0 wt % 235U

-0.04

-0.035

-0.03

-0.025

-0.02

-0.015

-0.01

-0.005

0

0.005

0 10 20 30 40 50 60

Burnup [GWd/MTU]

∆k [k(IFBA) - k(no_IFBA)]

0 year cooling time

5 year cooling time

20 year cooling time

40 year cooing time
Section 3 Reactivity Effect of IBA Rods

59

3.3.5.4 Reactivity Behavior in a Poisoned Storage Cell

The presence of fixed absorbers (e.g., boral panels), which are commonly used in SNF storage cells, affect the

neutron spectrum; and thus, their presence can be an important consideration in burnup credit analyses. This

is particularly true when estimating the reactivity worth of thermal neutron absorbers (e.g., fission products)

because the absorbers compete for neutrons with the poison (e.g., boron) in the fixed absorber panels. For

example, it has been shown that the reactivity worth of fission products is reduced by the presence of fixed

absorber panels,11 as compared to estimates of the reactivity worth of fission products in a configuration

without fixed absorbers panels (e.g., an infinite pin lattice). To evaluate the impact of fixed absorbers, a

number of the HELIOS calculations were repeated using an infinite array of poisoned storage cells from the

GBC-32 cask, which is a generic 32-PWR assembly cask developed to be representative of actual burnup

credit casks designed by industry. The boron loading in the boral panels in the GBC-32 cask is 0.225 g

10B/cm2

; detailed specifications for the GBC-32 cask are available in Ref. 12.

Calculations were performed for the Westinghouse 17 × 17 design with various IFBA loading and initial fuel

enrichments, and the ∆k values were compared to the ∆k values from the IFBA calculations, which were based

on an infinite array of assemblies. The results (∆k as a function of burnup) are shown in Figure 45, where it

can be seen that the reactivity effects are nearly the same in both environments. The cask environment

produces slightly higher ∆k, especially early in burnup, because the residual IBA material competes with the

poison in the absorber panels for thermal neutrons, thereby reducing the relative reactivity worth of the fixed

absorber panels. The maximum positive ∆k values for the various IFBA cases modeled in a cask environment

are summarized in Table 9.

The Siemens M1 and M3 designs were also modeled in the GBC-32 cell. The results (∆k as a function of

burnup) are shown in Figure 46, where it can be seen that the ∆k values are less negative in the storage cell

geometry. However, the ∆k values do remain negative throughout burnup.

While only the results for the IFBA and UO2-Gd2O3 IBA designs are presented here, studies for all the IBA

designs were performed in the poisoned storage cell. The results from these calculations confirm the validity

of the infinite assembly array calculations. However, it is noted that the positive ∆k effect of IFBAs is slightly

larger in the poisoned storage cell.
Reactivity Effect of IBA Rods Section 3

60

Figure 45 Comparison of ∆k values as a function of burnup for IFBA assemblies in differing

geometric configurations

-0.040

-0.035

-0.030

-0.025

-0.020

-0.015

-0.010

-0.005

0.000

0.005

0 10 20 30 40 50 60

Burnup [GWd/MTU]

∆k [k(IFBA) - k(no_IFBA)]

32 IFBA, cask storage cell array

32 IFBA, infinite assemblyarray

104 IFBA, cask storage cell array

104 IFBA, infinite assembly array

156 IFBA, cask storage cell array

156 IFBA, infinite assembly array
Section 3 Reactivity Effect of IBA Rods

61

Table 9 Summary of maximum positive ∆k values observed for IFBA cases in a cask environment

Enrichment

(wt % 235U)

IFBA loading pattern 3.0 4.0 5.0

32 0.0014 0.0013 0.0012

80 0.0024 0.0020 0.0018

104 0.0029 0.0025 0.0021

156 0.0054 0.0049 0.0044

Figure 46 Comparison of ∆k values as a function of burnup for assemblies with UO2-Gd2O3 rods

in differing geometric configurations

-0.10

-0.09

-0.08

-0.07

-0.06

-0.05

-0.04

-0.03

-0.02

-0.01

0.00

0 5 10 15 20 25 30 35 40 45 50 55 60

Burnup [GWd/MTU]

k [k(Gd2O3) - k(no_Gd2O3)]

M1, cask storage cell array

M1, infinite assembly array

M3, cask storage cell array

M3, infinite assembly array

∆

k
Reactivity Effect of IBA Rods Section 3

62

3.3.5.5 Axial Burnup

To evaluate the effect of IBAs with an axial burnup distribution present, a few 3-D criticality calculations

were performed using the KENO V.a module of the SCALE system,13 employing spent fuel isotopics

calculated by HELIOS. The KENO V.a criticality calculations used the 238-group cross-section library,

which is based primarily on ENDF/B-V data. For this analysis, the GBC-32 cask and Westinghouse 17 × 17

assembly with 4.0 wt % 235U initial enrichment were used. Calculations were performed for (1) a fuel

assembly with 104 IFBA rods (poison loading of 1.57 mg 10B/in.) and (2) a fuel assembly with no IFBA rods.

The active fuel length of the assemblies was divided into 18 equal-length axial regions to facilitate the axial

variation in SNF and IFBA coating compositions due to the axial burnup distribution. The axial profile

suggested in Ref. 14 for PWR fuel with assembly-average discharge burnup greater than 30 GWd/MTU was

used for this study.

While individual fuel pin (fuel and IFBA coating) compositions are tracked and utilized in the HELIOS

depletion and criticality calculations, lattice-average compositions (for both fuel and IFBA coating) were used

in the KENO V.a calculations. With this approximation, the KENO V.a calculations include lattice-average

fuel and IFBA coating compositions for each of the 18 axial fuel regions modeled [(1 UO2 fuel pin

composition) + (1 IFBA coating composition) × (18 axial regions) = 36 fuel-related compositions], as opposed

to separate fuel and IFBA compositions for each axial region of each fuel pin [neglecting symmetry, (264 UO2

fuel pin compositions) + (104 IFBA coating compositions) × (18 axial regions) = 6624 fuel-related

compositions]. Therefore, the following KENO V.a calculations and comparisons contain an approximation

and associated uncertainty that is not present in the previous HELIOS comparisons.

The KENO criticality calculations were performed with subsets of the available nuclides compositions from

HELIOS. The use of a subset of actinides in burnup credit calculations is referred to as “actinide-only”

burnup credit. The nuclides used here for actinide-only calculations are consistent with those specified in a

Department of Energy (DOE) Topical Report on burnup credit,14 with the exception that 236U and 237Np are

also included. The use of a subset of actinides and fission products is referred to herein as “actinide + fission

product” burnup credit. The fission product nuclides used here for actinide + fission product calculations are

consistent with those identified in Table 2 of Ref. 15 as being the most important for criticality calculations.

Table 10 displays the results from calculations assuming the IFBA coating extends the entire length of the fuel

(i.e., 144 inches), and shows that the cases with IFBA rods present yield a slightly lower effective neutron

multiplication factor, keff, than the cases without IFBA rods. This is due to the residual IBA material in the

underburned end regions, which dominate the neutron multiplication in the SNF. Hence, the difference

between cases with and without IFBAs decreases with burnup; for high burnups (e.g., 60 GWd/MTU) the

difference is very small.

A review of the relevant literature, however, indicates that the IFBA coating seldom (if ever) extends over the

entire active fuel length. Rather, the IFBA coating may vary in axial location and length depending on

particular core design requirements. Reference 16 indicates that the typical IFBA coating (or stack) length is

in the range of 116 to 134 inches, while Ref. 17 cites a minimum IFBA coating length of 108 inches.

Therefore additional calculations were performed for reduced IFBA coating lengths of 120 and 108 inches,

assuming the IFBA coating is centered with respect to the active fuel length. The results for the cases

corresponding to 120 and 108-inch IFBA coating lengths are given in Tables 11 and 12, respectively. The

results in Tables 11 and 12 show that for these shorter IFBA coating lengths, a case with IFBA rods can yield

higher keff values than a corresponding case without IFBA rods. Further, the results in Tables 10–12

demonstrate that that as the IFBA length decreases, the difference (∆k) between cases with and without IFBA
Section 3 Reactivity Effect of IBA Rods

63

rods present becomes positive earlier in burnup. This is because the lower-burnup end regions no longer

contain any residual IBA material due to the shorter IFBA coating lengths.

These results are important because they show that the small positive effect for fuel with IFBA rods that was

shown with 2-D calculations (infinite radial assembly arrays and infinite radial arrays of assemblies in

poisoned storage cells) does not exist in a 3-D cask model that includes the axial burnup distribution and fulllength axial poison loading. However, the results also show that the effect of the IFBA rods is dependent on

the axial length of the poison loading; for typical IFBA coating lengths (i.e., 116 to 134 inches) there is a small

positive effect associated with the IFBA rods. Finally, the results support the expectation that assemblies with

the other types of IBAs over the full axial length of the fuel are less reactive, as compared to assemblies

without IBAs, than what was shown with the 2-D calculations.
Reactivity Effect of IBA Rods Section 3

64

Table 10 Comparison of keff results in the GBC-32 cask for fuel with full-length (144 inch) IFBA rods

when the axial burnup distribution is included

keff ±1-σ

Burnup

(GWd/MTU)

Reference case

(no IFBAs present)

IFBA case

(104 IFBAs present)

∆k

(kIFBA – kno_IFBA)

Actinide-only

15 1.0621 ±0.0006 1.0518 ±0.0006 −0.0103 ±0.0008

30 0.9933 ±0.0006 0.9844 ±0.0005 −0.0089 ±0.0008

45 0.9419 ±0.0005 0.9350 ±0.0005 −0.0069 ±0.0007

60 0.8959 ±0.0006 0.8934 ±0.0006 −0.0025 ±0.0008

Actinides and fission products

15 1.0235 ±0.0006 1.0137 ±0.0005 −0.0098 ±0.0008

30 0.9406 ±0.0005 0.9336 ±0.0006 −0.0070 ±0.0008

45 0.8782 ±0.0006 0.8729 ±0.0006 −0.0053 ±0.0009

60 0.8223 ±0.0007 0.8209 ±0.0007 −0.0014 ±0.0010

Table 11 Comparison of keff results in the GBC-32 cask for fuel with IFBA rods when the

axial burnup distribution is included and the IFBA coating corresponds to the central 120 inches

of the axial fuel length

keff ±1-σ

Burnup

(GWd/MTU)

Reference case

(no IFBAs present)

IFBA case

(104 IFBAs present)

∆k

(kIFBA-kno_IFBA)

Actinide-only

15 1.0621 ±0.0006 1.0530 ±0.0006 −0.0091 ±0.0008

30 0.9933 ±0.0006 0.9872 ±0.0006 −0.0061 ±0.0008

45 0.9419 ±0.0005 0.9496 ±0.0006 0.0077 ±0.0008

60 0.8959 ±0.0006 0.9045 ±0.0006 0.0086 ±0.0008

Actinides and fission products

15 1.0235 ±0.0006 1.0155 ±0.0005 −0.0080 ±0.0008

30 0.9406 ±0.0005 0.9357 ±0.0006 −0.0049 ±0.0008

45 0.8782 ±0.0006 0.8843 ±0.0006 0.0061 ±0.0008

60 0.8223 ±0.0007 0.8295 ±0.0006 0.0072 ±0.0008
Section 3 Reactivity Effect of IBA Rods

65

Table 12 Comparison of keff results in the GBC-32 cask for fuel with IFBA rods when the

axial burnup distribution is included and the IFBA coating corresponds to the central 108 inches

of the axial fuel length

keff ±1-σ

Burnup

(GWd/MTU)

Reference case

(no IFBAs present)

IFBA case

(104 IFBAs present)

∆k

(kIFBA-kno_IFBA)

Actinide-only

15 1.0621 ±0.0006 1.0566 ±0.0006 −0.0055 ±0.0008

30 0.9933 ±0.0006 0.9996 ±0.0006 0.0063 ±0.0008

45 0.9419 ±0.0005 0.9501 ±0.0006 0.0082 ±0.0008

60 0.8959 ±0.0006 0.9048 ±0.0006 0.0089 ±0.0008

Actinides and fission products

15 1.0235 ±0.0006 1.0213 ±0.0007 −0.0022 ±0.0008

30 0.9406 ±0.0005 0.9476 ±0.0006 0.0070 ±0.0008

45 0.8782 ±0.0006 0.8857 ±0.0005 0.0075 ±0.0008

60 0.8223 ±0.0007 0.8304 ±0.0006 0.0081 ±0.0008

Section 4 Summary and Implications

67

4 SUMMARY AND IMPLICATIONS

Numerous IBA types have been used in commercial nuclear fuel assembly designs to suppress initial

reactivity. Variations in the IBA poison material, composition, placement within rods, and rod configurations

exist among current PWR IBA fuel assembly designs. These IBA characteristics are varied in combination

with the initial fuel assembly enrichment and core location to achieve core operating and fuel management

goals. To assess the potential impact of these IBAs on the reactivity of SNF, analyses have been presented in

this report for all IBA types widely used in U.S. PWRs. These IBA types include IFBA rods, UO2-Gd2O3

rods, UO2-Er2O3 rods, and Al2O3-B4C rods. Based on the available data, analyses were performed for a

realistic range of fuel initial enrichment and poison loading combinations based on actual assemblies.

Assemblies with IBAs are designed such that the poison material is effectively depleted during the first third

of the assembly life. As a result, depending on the IBA type and loading, the assembly reactivity may actually

increase with burnup to a maximum or peak at which the IBA is essentially depleted. At this point

(or beyond), the reactivity of an assembly with IBAs may equal or exceed the reactivity of a nonpoisoned

equivalent enrichment assembly. To determine if any of the assemblies with IBAs exceed the reactivity of

nonpoisoned equivalent enrichment assemblies, this evaluation compared kinf values, throughout burnup,

between fuel assemblies with and without IBAs. The premise was that if assemblies with IBAs yield lower kinf

values (throughout burnup) than assemblies without IBAs, burnup credit criticality safety analyses may

conservatively neglect the presence of IBAs. Thus, for each unique IBA assembly design considered, a

calculation was performed for (1) the actual assembly specification (including the presence of the IBA) and

(2) an artificial condition in which the IBA was neglected (nonpoisoned equivalent enrichment case).

The difference between the calculated kinf values (∆k) from these two cases was used to assess the reactivity

effect of each of the IBA types.

First, 2-D analyses were performed for Westinghouse assembly designs with IFBA rods, including varying

numbers of IFBA rods, varying boron loadings, and various initial fuel enrichments. The results (as a function

of burnup) indicate that the kinf values for fuel assemblies containing IFBA rods may exceed the kinf values for

fuel assemblies without IFBA rods. In fact, all IFBA assembly designs considered were shown to have a

small, positive reactivity effect, as compared to assemblies without IFBA rods. For a fixed initial fuel

enrichment, the positive reactivity effect was shown to increase with increasing poison loading (i.e., both

increasing 10B loading and increasing number of IFBA rods). Additionally, for a fixed poison loading, the

positive reactivity effect was shown to increase with decreasing initial fuel enrichment. For all of the IFBA

cases considered, the maximum positive ∆k value was found to be ~0.5%. However, this maximum value

corresponds to a very heavily poisoned, atypical assembly design and, based on the available data, is not

representative of typical IFBA assemblies. The positive reactivity effect for typical IFBA assembly designs

(designs that are used in significant numbers) appears to be less than ~0.3%.

Second, analyses were performed for assembly designs with UO2-Gd2O3 rods, including variations in gadolinia

loadings (wt % Gd2O3 and number of gadolinia-bearing rods) and initial fuel enrichment. The calculations

showed that, throughout burnup, the kinf values for fuel assemblies containing UO2-Gd2O3 rods remain less

than the kinf values for fuel assemblies without UO2-Gd2O3 rods. Therefore, it is concluded that gadoliniabearing rods yield a negative reactivity effect. The negative reactivity effect was found to increase with

increasing gadolinia loading (wt % Gd2O3 and number of gadolinia-bearing rods) and increasing initial fuel

enrichment. For all of the actual fuel assembly designs considered, the kinf values for fuel assemblies with

gadolinia-bearing rods were always less (by as much as 0.5%) than the kinf values for fuel assemblies without

gadolinia-bearing rods.
Summary and Implications Section 4

68

Third, analyses were performed for assembly designs with UO2-Er2O3, including variations in the number of

UO2-Er2O3 rods. For the assembly designs considered, the kinf values for the non-erbia-bearing fuel assemblies

consistently remained higher than the kinf values for the erbia-bearing fuel assemblies. The amount by which

the kinf values for the erbia-bearing fuel were less than the kinf values for the non-erbia-bearing fuel increased

with increasing numbers of UO2-Er2O3 rods. Therefore, it is concluded that erbia-bearing rods yield a negative

reactivity effect.

Finally, analyses were performed for assembly designs with Al2O3-B4C rods, including variations in the

number Al2O3-B4C rods. For the assembly designs considered, the kinf values for the fuel assemblies with

Al2O3-B4C rods consistently remained lower than the kinf values for the fuel assemblies without Al2O3-B4C

rods. The amount by which the kinf values for the fuel assemblies with Al2O3-B4C rods were less than the kinf

values for the fuel assemblies without Al2O3-B4C rods increased with increasing numbers of Al2O3-B4C rods.

Therefore, it is concluded that Al2O3-B4C rods yield a negative reactivity effect.

In summary, the analyses presented in Sections 3.3.1–3.3.4 demonstrate that the neutron multiplication factor

for an assembly without IBAs is always greater (as a function of burnup) than the neutron multiplication

factor for an assembly using any of the following IBA types: UO2-Gd2O3, UO2-Er2O3, or Al2O3-B4C rods.

Conversely, the neutron multiplication factor for an assembly with IFBA rods present was found to exceed

(maximum of 0.4% ∆k) the neutron multiplication factor for an assembly without IFBA rods. Therefore,

neglecting the IBAs in a burnup-credit criticality safety analysis will yield conservative results for assembly

designs with UO2-Gd2O3, UO2-Er2O3, or Al2O3-B4C IBA rods and nonconservative results for assembly

designs with IFBA rods. In all cases, for burnups characteristic of discharge, the reactivity effect of IBAs is

relatively small (less than ~1.0% ∆k) and generally well behaved.

These results are consistent with previous work, which provided illustrative examples of the reactivity effects

of several of the IBA types based on a 2-D analysis of a single case for each type.3 However, 3-D cask

calculations performed for this study (Section 3.3.5.5) demonstrated that when the axial burnup distribution is

included, assemblies with IFBA rods are actually less reactive than corresponding assemblies without IFBA

rods because of the residual absorber in the low-burnup end regions. The amount by which the IFBA

assemblies are less reactive decreases with burnup, as the residual absorber in the end regions is depleted.

This finding supports the expectation that assemblies with the other types of IBAs over the full axial length are

also less reactive than what was shown with the 2-D calculations. However, the results also show that the

effects of the IFBA rods is dependent on the axial length of the poison loading and that for typical IFBA

coating lengths (i.e., 116 to 134 inches) there is a small positive effect associated with the IGBA rods.

These results are important to burnup credit because they demonstrate that assembly designs with IFBA

UO2-Gd2O3, UO2-Er2O3, or Al2O3-B4C IBA rods are less reactive throughout burnup than their corresponding

designs without the IBA rods (i.e., nonpoisoned equivalent enrichment). Consequently, with the notable

exception of assemblies with IFBA rods, neglecting the presence of IBAs in a burnup credit criticality safety

evaluation will yield slightly conservative results. For assembly designs with IFBA rods, the positive

reactivity effect is small and dependent on the axial length of the IFBA coating.
Section 5 Recommendations

69

5 RECOMMENDATIONS

The analyses presented in this report provide a technical basis for revising NRC’s Interim Staff Guidance

(Ref. 2) to allow burnup credit for assembly designs with IBAs. Although the analyses do not address the

issue of validation of depletion methods for assembly designs with IBAs, they do demonstrate that the effect of

the IBAs is relatively small (at or near target discharge burnups) and generally well behaved. Furthermore, the

recommended approaches for addressing fuel assemblies with IBAs, as described below, do not involve

explicit analyses with IBAs present and thus do not necessitate validation of the depletion methods for

assembly designs with IBAs. Therefore, it is concluded that the effect of the various IBA types may be

adequately calculated and that the current restriction on assemblies with IBAs should be eliminated.

The analyses described in this report conclusively demonstrate that, with the exception of the Westinghouse

IFBA rods, the neutron multiplication factor for an assembly without IBAs is always greater (throughout

burnup) than the neutron multiplication factor for an assembly with IBAs, including UO2-Gd2O3, UO2-Er2O3,

and Al2O3-B4C rods. Therefore, for those IBAs other than IFBAs, burnup credit criticality safety analyses

may simply and conservatively neglect the presence of the IBAs by assuming nonpoisoned equivalent

enrichment fuel. Considering the variations in IBA assembly designs, neglecting the presence of the IBAs is an

important simplifying assumption that does not add significant unnecessary conservatism.

For assembly designs with IFBA rods, 2-D (radially infinite) calculations have demonstrated that the neutron

multiplication factor is slightly greater (maximum of 0.4% ∆k) than the neutron multiplication factor for

assembly designs without IFBA rods. Three-dimensional cask calculations showed that when the axial burnup

distribution is included, assemblies with full-axial length IFBA coatings are less reactive than corresponding

assemblies without IFBA rods, because of the residual absorber in the low-burnup end regions. However, the

results also indicated that the effect of the IFBA rods is dependent on the axial length of the poison loading and

that for typical IFBA coating lengths (i.e., 116 to 134 inches), there is a small positive effect associated with

the IFBA rods. Consequently, the positive reactivity effect due to the presence of IFBA rods should be

considered in any burnup-credit criticality safety analysis seeking to qualify IFBA assemblies as acceptable

contents. Due to the significant variations in IFBA assembly designs, simple strategies for addressing the

positive reactivity effect are desirable. Two possible strategies for consideration include: (1) the inclusion of

a small reactivity bias to bound the effect of the IFBA rods, or (2) demonstration that the effect of the IFBA

rods is bounded by the effect of other modeling assumptions (e.g., BPR exposure). The use of a reactivity bias

will require justification for the value of the bias. Alternatively, it may be simpler and less burdensome to

demonstrate that the effect of the IFBA rods is bounded by the BPR modeling approach. Comparison of the

reactivity effect of IFBA rods (shown in this report) to the reactivity effect of BPRs (as quantified in Ref. 1),

clearly demonstrates that the reactivity effect of the IFBA rods is significantly less than the reactivity effect

due to BPRs. Furthermore, considering the fact that BPRs are seldom used within assemblies that have IFBA

rods, and when used, are employed in a limited way (e.g., a small number of BPRs may be used in conjunction

with an assembly that has a relatively light IFBA loading), reliance on the BPR modeling to account for the

effect of IFBA rods is justified. However, this approach would only be applicable to analyses that consider

BPR exposure.

71

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73

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94. L. Markova, Ustav jaderneho vyzkumu Rez, Theoretical Reactor Physics, Nuclear Research

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95. Daniel Marloye, Belgonucléaire, Av. Ariane 4, B-1200, Brussels, Belgium

96. J. R. Massari, Constellation Nuclear Services, 2200 Defense Highway, Suite 405, Crofton, MD

21114-2404

97. C. W. Mays, Framatome ANP, 3315 Old Forest Road, P.O. Box 10935, Lynchburg, VA 24506-

0935

98. J. N. McKamy, U.S. Department of Energy, Office of Engineering Assistance and Site Interface,

EH-34, 19901 Germantown Rd., Germantown, MD 20874

99. N. B. McLeod, JAI Corporation, 4103 Chain Bridge Road, Suite 200, Fairfax, VA 22030

100. D. Mennerdahl, E. Mennerdahl Systems, Starvägen 12, S-183 57 Täby, Sweden

101. R. L. Murray, 8701 Murray Hill Drive, Raleigh, NC 27615

102. J. A. Myers, U. S. Nuclear Regulatory Commission, NMSS/SFPO/TRD, MS O13-D13,

Washington, DC 20555-0001

103. K. A. Neimer, Duke Engineering & Services, 400 S. Tyron St., WC26B, P.O. Box 1004,

Charlotte, NC 28201-1004

104. P. Noel, Bechtel SAIC Company, LLC, 1261 Town Center Drive, Las Vegas, Nevada 89134

105. I. Nojiri, Japan Nuclear Cycle Development Institute, Environment and Safety Division, Tokai

Works, Muramatsu Tokai-mura, Naka-gun Ibaraki-ken 319-1194, Japan

106. J. C. Neuber, SIEMENS AG, KWU NS-B, Berliner Str. 295-303, D-63067 OFFENBACH AM

MAIN, Germany

107. A. Nouri, OECD/NEA Data Bank, Le Seine-Saint Germain, 12 Boulevard des Iles, F-92130

Issy-les-Moulineaux, France

108–109. Office of Scientific and Technical Information, U.S. Department of Energy, P.O. Box 62,

Oak Ridge, TN 37831

110. Office of the Assistant Manager for Energy Research and Development, Department of Energy

Oak Ridge Operations (DOE-ORO), P.O. Box 2008, Oak Ridge, TN 37831

111. H. Okuno, Japan Atomic Energy Research Institute, Department of Fuel Cycle, Safety Research,

2-4 Shirakata-Shirane, 319-1195 Tokai-mura, Naka-Gun, Ibaraki-ken, Japan

112. P. M. O’Leary, Framatome ANP, 3315 Old Forest Road, P.O. Box 10935, Lynchburg, VA

24506-0935

113. J. R. Massari, Constellation Nuclear Services, 2200 Defense Highway, Suite 405, Crofton, MD

21114-2404

114. N. L. Osgood, U.S. Nuclear Regulatory Commission, Office of Nuclear Materials Safety and

Safeguards, MS O13-D13, Washington, DC 20555-0001

115. T. Parish, Department of Nuclear Engineering, Texas A & M University, College Station, TX

77843-3313
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116. V. A. Perin, U.S. Nuclear Regulatory Commission, NMSS/FCSS/SPB, MS T8-A33,

Washington, DC 20555-0001

117. B. Petrovic, Westinghouse Electric Company, Science and Technology Department, 1344 Beulah

Road, Pittsburgh, PA 15235

118. J. S. Philbin, Sandia National Laboratory, PO Box 5800, Mail Stop 1143, Albuquerque,

New Mexico 87185-1143

119. M. L. Pitts, Framatome ANP, 3315 Old Forest Road, PO Box 10935, Lynchburg, VA 24506-

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120. M. Rahimi, U.S. Nuclear Regulatory Commission, NMSS/DWM/HLWB, MS T7-F3,

Washington, DC 20555-0001

121. E. L. Redmond II, Holtec International, 555 Lincoln Drive West, Marlton, NJ 08053

122. Frederik Reitsma, Reactor Theory, Building 1900, Pelindaba, NECSA, PO Box 582,

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123. C. Rombough, CTR Technical Services, Inc., 5619 Misty Crest Dr., Arlington, TX 76017-4147

124. J. E. Rosenthal, U.S. Nuclear Regulatory Commission, RES/DSARE/SMSAB, MS T10-K8,

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125. D. Salmon, Bechtel SAIC Company, LLC, 1261 Town Center Drive, Las Vegas, Nevada 89134

126. A. Santamarina, Commissariat A L’Energie Atomique, Nuclear Reactor Division, Reactor

Studies Department, Reactor and Cycle Physics Service,

CEA/CADARACHE/DRN/DER/SPRC Bat. 230, 13108 Saint-Paul-Lez-Durance, Cedex,

France

127. E. Sartori, OECD/NEA Data Bank, Le Seine-Saint Germain, 12 Boulevard des Iles, F-92130

Issy-les-Moulineaux, France

128. J. J. Sapyta, Framatome ANP, 3315 Old Forest Road, P.O. Box 10935, Lynchburg, Virginia

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129. D. Schrire, Hot Cell Laboratory, Studsvik Nuclear, Nyköping 61182, Sweden

130. H. H. Schweer, Bundesamt fuer Strahlenschutz, Willi Brandt Str. 5, D-38226 SALZGITTER,

Germany

131. M. Smith, Virginia Power Co., P.O. Box 2666, Richmond, VA 23261

132. N. R. Smith, AEA Technology, A32 Winfrith, Dorchester, Dorset DT2 8DH, United Kingdom

133. Rudi Stamm'ler, Studsvik Scandpower AS, Gasevikveien 2, P O Box 15, NO-2027 Kjeller

Norway

134. J. T. Stewart, Department of Environment, Transport, and Re, RMTD, 4/18, GMH, 76

Marsham Street, London SW1P 4DR, United Kingdom

135. T. Suto, Power Reactor and Nuclear Fuel Development Corporation, Technical Service Division,

Tokai Reprocessing Plant, Tokai Works, Tokai-Mura, Naka-gun, Ibaraki-ken, Japan

136. H. Taniuchi, Kobe Steel, Ltd., 2-3-1 Shinhama, Arai-Cho, Takasago, 676 Japan

137. D. A. Thomas, Bechtel SAIC Company, LLC, 1261 Town Center Drive, Las Vegas, Nevada

89134

138. P. R. Thorne, British Nuclear Fuels plc (BNFL), Nuclear and Radiological Safety,

R101 Rutherford House, Risley Warrington WA3 6AS, United Kingdom

139. J. R. Thornton, Duke Engineering & Services, 230 S. Tyron St., P.O. Box 1004, Charlotte, NC

28201-1004

140. S. E. Turner, Holtec International, 230 Normandy Circle East, Palm Harbor, FL 34683

141. A. P. Ulses, U.S. Nuclear Regulatory Commission, NRR/DSSA/SRXB, MS O10-B3,

Washington, DC 20555-0001
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142. M. E. Wangler, U.S. Department of Energy, EH-33.2, Washington, DC 20585-0002

143. M. D. Waters, U.S. Nuclear Regulatory Commission, NMSS/SFPO/SLID, MS O13-D13,

Washington, DC 20555-0001

144. A. Wells, 2846 Peachtree Walk, Duluth, GA 30136

145. S. A. Whaley, U.S. Nuclear Regulatory Commission, NMSS/SFPO/TRD, MS O13-D13,

Washington, DC 20555-0001

146. B. H. White, U.S. Nuclear Regulatory Commission, NMSS/SFPO/TRD, MS O13-D13,

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147. R. Wilson, Rocky Flats Field Office, USDOE, 10808 Highway 93, Golden, CO 80403-8200

148. C. J. Withee, U.S. Nuclear Regulatory Commission, NMSS/SFPO/TRD, MS O13-D13,

Washington, DC 20555-0001
NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION

(2-89)

NRCM 1102 BIBLIOGRAPHIC DATA SHEET

3201, 3202 ( See instructions on the reverse)

1. REPORT NUMBER

(Assigned by NRC, Add Vol., Supp.,

Rev., and Addendum Numbers,

if any.)

NUREG/CR-6760

ORNL/TM-2000/321 2. TITLE AND SUBTITLE

Study of the Effect of Integral Burnable Absorbers for PWR Burnup Credit

3. DATE REPORT PUBLISHED

MONTH

March

YEAR

2002

4. FIN OR GRANT NUMBER

W6479

5. AUTHOR(S)

C. E. Sanders and J. C. Wagner

6. TYPE OF REPORT

Technical

7. PERIOD COVERED (Inclusive Dates)

8. PERFORMING ORGANIZATION — NAME AND ADDRESS (If NRC, provide Division, Office or Region, U.S. Nuclear Regulatory Commission, and mailing address;

if contractor, provide name and mailing address.)

Oak Ridge National Laboratory

Managed by UT-Battelle, LLC

Oak Ridge, TN 37831-6370

9. SPONSORING ORGANIZATION — NAME AND ADDRESS (If NRC, type “Same as above”; if contractor, provide NRC Division, Office or Region, U.S. Regulatory

Commission, and mailing address.)

Division of Systems Analysis and Regulatory Effectiveness

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

10. SUPPLEMENTARY NOTES

R. Y. Lee, NRC Project Manager

11. ABSTRACT (200 words or less)

The Interim Staff Guidance on burnup credit issued by the U.S. Nuclear Regulatory Commission’s Spent Fuel Project Office

recommends restricting the use of burnup credit to assemblies that have not used burnable absorbers. This restriction

eliminates a large portion of the currently discharged spent fuel assemblies from cask loading, and thus severely limits the

practical usefulness of burnup credit. This report examines the effect of integral burnable absorbers (IBAs) on reactivity

to provide technical justification for relaxing the current restriction for dry storage and transportation, and subsequently, for

developing the necessary guidelines for relaxing the current restriction. The effect of IBAs on reactivity for various IBA

designs is shown and discussed. Further, the reactivity effect of IBAs for typical initial fuel enrichment and absorber

loadings is quantified as a function of burnup. The report concludes with a discussion of the issues for consideration and

preliminary recommendations to expand the use of burnup credit to include spent fuel assemblies with IBAs.

12. KEY WORDS/DESCRIPTORS (List words or phrases that will assist researchers in locating the report.)

criticality safety, burnup credit, transportation, dry storage, integral burnable absorbers

13. AVAILABILITY STATEMENT

unlimited

14. SECURITY CLASSIFICATION

(This Page) unclassified

(This Report) unclassified

15. NUMBER OF PAGES

16. PRICE

 NRC FORM 335 (2-89)