3 

Mechanical Fatigue Testing of 

High-Burnup Fuel for 

Transportation Applications

Office of Nuclear Regulatory Research

NUREG/CR-7198, Rev. 1 

ORNL/TM-2016/689 
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SR-CR 08/2016 
Mechanical Fatigue Testing 

of High-Burnup Fuel for 

Transportation Applications 

Manuscript Completed: December 2016 

Date Published: October 2017 

Prepared by: Jy-An Wang and Hong Wang

Oak Ridge National Laboratory 

Managed by UT-Battelle, LLC 

Oak Ridge, TN 37831-6069 

Bruce Bevard, ORNL Program Manager

Michelle Flanagan-Bales, NRC Project Manager 

NRC Job Code N6686/N6789

Office of Nuclear Regulatory Research

NUREG/CR-7198, Rev. 1 

ORNL/TM-2016/689 

iii

ABSTRACT

Testing was conducted at Oak Ridge National Laboratory (ORNL) to determine the ability of 

high burnup (HBU) (>45 GWd/MTU) spent nuclear fuel (SNF) to maintain its integrity under 

conditions relevant to storage and transport. The cyclic integrated reversible-bending fatigue 

tester (CIRFT) is an innovative system used in these tests. The CIRFT can impose pure 

bending loads on SNF test specimens and can measure the in situ curvature of the fuel rod 

during bending. Pure bending is a condition of stress in which a bending moment is applied to a 

beam without the simultaneous presence of axial, shear, or torsional forces. The CIRFT is 

composed of a U-frame equipped with load cells. A setup with three linear variable differential 

transformers (LVDTs) was used. 

Since the initial publication of this NUREG in May 2015, several new factors were identified that 

influence the recorded curvature measurement data. Because these factors were not accounted 

for in the first publication, the initial quantitative results are not considered valid. This revision

provides updated results of the 2015 CIRFT tests. 

After the 2015 publication, a new phase of testing was implemented that used samples treated 

to induce radial hydrides. This revision includes the results of CIRFT tests on segments

subjected to hydride reorientation treatment (HRT), thus broadening the applicability of the 

research results. This first revision entirely replaces the initial May 2015 publication.

The HBU SNF rods with Zircaloy-4 (Zry-4) cladding were studied under static and dynamic 

(cyclic) bending. They survived static unidirectional bending to a maximum curvature of 

2.2–2.5 m-1, or a maximum moment of 85–87 N∙m. The maximum longitudinal cladding strain 

before failure or before reaching CIRFT displacement capacity was 1.2–1.3%. The segment 

composite structure of an HBU rod introduces numerous stress riser sites into the rod system,

often resulting in specimen fracture at the pellet-to-pellet interface regions under both static and 

dynamic CIRFT testing. Nevertheless, the static CIRFT test results on HBU rods indicate that 

fueled rods have a significantly greater flexural rigidity than the calculated flexural rigidity for 

HBU cladding alone. 

During dynamic testing, the fatigue life of HBU rods mainly depended on the level of loading. 

Under loading with moments of ±5.08 to ±35.56 N∙m, which resulted in an equivalent strain (ε)

of ±0.025 to ±0.31% strain at 5 Hz, the fatigue life (N) ranged from 5.5 × 103 to 2.3 × 106

 cycles. 

Considering the complexity and nonuniformity of the HBU fuel cladding system, it was 

significant to find that the ε-N data for the HBU were characterized by a curve that would be

expected for standard uniform materials. The ε-N curve of the HBU rods can be described by a 

power function of y = 3.839 · x−0.298, where x is the number of cycles to failure, and y is the strain 

amplitude (%). It was also significant to find that, if an endurance limit is defined by survival of 

>107

 cycles (as is typical for material fatigue endurance limits), then the HBU rods tested 

exhibited an endurance limit. The endurance limit for the HBU rods tested is likely between 

strains 0.03–0.05%. 

The maxima of the imposed curvature κ during dynamic testing ranged from ±0.07 to ±0.62 m−1

at 5 Hz. The κ-N curve of the HBU rods can be described by a power function of 

y = 6.864·x−0.283, where x is the number of cycles to failure and y is the maxima of clad tensile 

curvature |κ|max (m-1). An endurance limit is likely between 0.07–0.12 m−1 when it is defined at 

107

 cycles. 
iv

From very limited CIRFT test results of rods that underwent HRT, the following phenomena 

were observed. 

• HRT rods in static testing showed that the maximum moment decreased in comparison

to that of as-irradiated fuel rods at similar curvature levels.

• Both the HRT static and dynamic flexural rigidity are less than that of as-irradiated

baseline data.

• The load amplitude vs. failure frequency trend indicated that HRT rods had slightly less

fatigue life than those indicated in the as-irradiated baseline CIRFT data.
v 

FOREWORD

To obtain certification for a spent nuclear fuel (SNF) storage cask and transportation package, 

the US Nuclear Regulatory Commission (NRC) requires evaluation of nuclear fuel performance 

under normal and accident conditions as specified in 10 Code of Federal Regulations (CFR) 

Parts 71 and 72. The safety and security of SNF storage and transportation systems must be 

demonstrated to obtain certification. As part of the requirement to demonstrate safety, the 

applicant must assess the condition and performance of SNF in the storage and transportation 

system during operations and design-basis accidents. 

In the United States, SNF is expected to be transported off site to at least one storage facility 

before permanent disposal. While the SNF is in transport or in dry storage, it will experience 

various localized stresses. As nuclear fuel is now regularly irradiated to high-burnup (HBU) 

values (>45 GWd/MTU), any phenomena associated with HBU SNF that could impact fatigue 

resistance (e.g., hydride reorientation) should be evaluated under storage and transportation

conditions. 

For these reasons, the NRC tasked the Oak Ridge National Laboratory (ORNL) to investigate 

the flexural rigidity and fatigue life of HBU SNF. Testing was designed to evaluate the fuel rod 

as a composite system, including the presence of fuel inside the cladding and any 

pellet/cladding bonding effects. The project proceeded in two phases. Phase I involved testing 

HBU SNF in the as-irradiated state, in which hydrides are expected to be predominantly in the 

circumferential-axial orientation. Phase II involved testing HBU SNF segments following a 

treatment designed to reorient the hydrides in the cladding to be predominantly in the radialaxial orientation. The properties measured in this testing program should be observed to

evaluate SNF integrity under storage and transportation conditions, including the details of SNF 

package or cask design and expected loads. 

This first revision entirely replaces the initial May 2015 publication. It includes revised results of 

previous CIRFT tests, as well as new results of testing on the samples treated to induce radial 

hydrides. 

vii

TABLE OF CONTENTS 

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

FOREWORD...............................................................................................................................v

TABLE OF CONTENTS............................................................................................................ vii

LIST OF FIGURES ....................................................................................................................ix

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

EXECUTIVE SUMMARY .........................................................................................................xiii

ACKNOWLEDGMENTS............................................................................................................xv

LIST OF ACRONYMS, ABBREVIATIONS, AND UNITS........................................................xvii

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

2 APPROACH.........................................................................................................................3

3 EXPERIMENTAL METHOD.............................................................................................. 7

3.1 The U-Frame Setup .................................................................................................... 7

3.2 Grip Design................................................................................................................. 9

3.3 LVDT Measurements and Corrections ......................................................................13

3.3.1 Correction for Disk-Type LVDT Head Bias.....................................................13

3.3.2 Curvature Adjustment for LVDT Stem Dynamic Stability................................16

3.4 Test Segment Selection and Preparation ..................................................................17

3.4.1 Hydride-Reorientation Treatment (HRT) Procedure.......................................19

3.5 Test Protocol.............................................................................................................20

3.5.1 Static Testing................................................................................................. 20

3.5.2 Dynamic Testing............................................................................................21

3.6 Data Processing........................................................................................................22

3.6.1 Flexural Rigidity Calculation...........................................................................23

4 RESULTS..........................................................................................................................26

4.1 Static Tests ...............................................................................................................26

4.1.1 Characteristics of Moment-Curvature Curve ..................................................32

4.1.2 Comparison of Static Results with PNNL Cladding Data................................34

4.2 Dynamic Testing .......................................................................................................36

4.2.1 κ-N Curve ......................................................................................................50

4.3 Post-Irradiation Examination .....................................................................................55

4.3.1 As-Irradiated Samples PIE.............................................................................55

4.3.2 HRT Samples PIE .........................................................................................58

4.4 Cladding Oxide Effect Study......................................................................................59

4.5 HRT CIRFT Test Data Evaluation .............................................................................60

5 CONCLUSIONS ................................................................................................................63
viii

6 REFERENCES ...................................................................................................................65

APPENDIX A ROD SEGMENTS AND CUTTING PLAN .....................................................A-1

APPENDIX B METROLOGY OF ROD SEGMENTS...........................................................B-1

APPENDIX C BENCHMARKING OF TEST SYSTEM.........................................................C-1

APPENDIX D STATIC TEST RESULTS .............................................................................D-1

APPENDIX E DYNAMIC TEST RESULTS...........................................................................E-1

APPENDIX F POST-IRRADIATION EXAMINATION (PIE) ................................................ F-1

APPENDIX G VERIFICATION OF CURVATURE MEASUREMENT ..................................G-1

APPENDIX H HYDRIDE REORIENTATION SAMPLE PREPARATION............................H-1
ix

LIST OF FIGURES

Figure 1 The cyclic integrated reversible-bending fatigue tester developed by ORNL.............. 3

Figure 2 Determination of the rod’s bending curvature by use of deflections measured at 

three points................................................................................................................ 4

Figure 3 Key fuel and cladding features and fuel rod phenomena of (a) an unloaded fuel 

rod, and (b) a fuel rod under a bending load. ............................................................. 5

Figure 4 (a) Horizontal layout of ORNL U-frame setup; (b) rod specimen and three 

LVDTs for curvature measurement; and (c) front view of CIRFT installed in 

ORNL hot cell. ........................................................................................................... 8

Figure 5 Schematic drawings of U-frame setup for reversal bending when rigid arms are 

(a) closing, (b) neutral, and (c) opening. ∆ is the rigid rod end displacememnt at 

loading joint. .............................................................................................................. 9

Figure 6 Assembled grip blocks (left) and two views (right) of detached grip body.................10

Figure 7 Image showing the grip design of CIRFT with one end-block removed....................10

Figure 8 (a) Vise mold for assembling octagonal rigid sleeves into a rod, (b) rod 

(Demo1, 606B2) to be inserted, (c) one pin turned away, rigid sleeve loaded 

with epoxy inserted into the lower chamber. ............................................................12

Figure 9 LVDT disk-type probe bias: effect of LVDT probe with disk shape on curvature 

measurement where contact points are shifted outwards (on the compression 

side of the rod) and inwards (on the tension side of the rod)....................................14

Figure 10 Diagram of circle showing quantities in calculating half gauge length of 

bent rod. ..................................................................................................................15

Figure 11 Curvature measurements of polycarbonate rod based on M-based estimate, 

and disk type probe, 2.50 mm sensor spacing correction.........................................16

Figure 12 Variations of (a) peak strain and (b) valley strain as a function of driving 

frequeny. At 5 Hz, corrected strain level is about 18% higher than strain gauge 

estimate for both peak and valley strain plots. .........................................................17

Figure 13 Segment 608C4B for DL2 taken from rod G10, with the storage tube......................18

Figure 14 Sample temperature as a function of time for HR1...................................................20

Figure 15 Flowchart for cyclic testing of SNF rod.....................................................................21

Figure 16 Moment, curvature time series, and moment-curvature hysteresis loop for rod 

D1. Two measurements are displayed with relative displacement 0.8 and 1.2 

mm at the loading points of U-frame. .......................................................................23

Figure 17 Empty-run conducted without specimen installed, for which the load and 

displacement were plotted; associated with static test on D1...................................25

Figure 18 Curve of moment versus curvature.....................................................................26

Figure 19 Mating fracture surfaces (a) and (b); lateral sides near the neutral axis of the 

bending rod for S1 (606C3C) (c) and (d).................................................................. 27

Figure 20 (a) Moment-curvature curves measurements and (b) stress and strain curves in 

static tests................................................................................................................29

Figure 21 Characteristic points of moment-curvature curve. ....................................................32

Figure 22 Comparison of CIRFT global data with theoretical PNNL moment-curvature 

derived from PNNL cladding only stress-strain data.................................................34

Figure 23 Moment and curvature as a function of time and moment-curvature loops 

based on measurements when (a) N=1 and (b) N = 111,000 cycles for D1 

(607C4B); measurements were made with 0.8 and 1.2 mm relative 

displacements; Nf = 1.1 × 105 cycles under ±15.24 N∙m, 5 Hz. Fuel particles

collected <1.0 g. ......................................................................................................37
x 

Figure 24 (a) Moment-curvature relation and (b) moment-flexural rigidity relation at 

various numbers of cycles for D1 (607C4B); Nf = 1.1 × 105 cycles under ±15.24 

N∙m, 5 Hz. Fuel particles collected <1.0 g................................................................38

Figure 25 Variations of (a) curvature range, (b) moment range, (c) flexural rigidity as a 

function of number of cycles for D1 (607C4B); Nf = 1.1 × 105 cycles under 

±15.24 N∙m, 5 Hz; fuel particles collected <1.0 g. ................................................. 38

Figure 26 Variations of (a) curvature range, (b) applied moment range, (c) flexural rigidity, 

(d) maximum and minimum values of curvature, (e) maximum and minimum 

values of moment, and (f) flexural hysteresis as a function of number of cycles 

for D1 (607C4B); Nf = 1.1 × 10

5 cycles under ±15.24 N∙m, 5 Hz; fuel particles 

collected <1.0 g. ......................................................................................................39

Figure 27 Moment and curvature as a function of time at (a) 26 and (b) 1.10 × 105

 cycles; 

results based on online monitoring;the initial unsymmetrical curvature 

deformation mode of the tested rod was stabilized and replaced with a

symmetrical curvature deformation mode at higher cycles.......................................41

Figure 28 (a) and (b) mating fracture surfaces, (c) frontal, and (d) back sides for D1 

(607C4B), Nf = 1.1 × 105 cycles under ±15.24 N∙m, 5 Hz. Fuel particles 

collected <1.0 g. ......................................................................................................42

Figure 29 Equivalent strain amplitudes (∆ε/2) as a function of the number of cycles to 

Markers with arrows indicate that the tests stopped without failure..........................49

Figure 30 Equivalent strain amplitudes (∆ε/2) as a function of number of cycles to failure 

with curve fitting extended to include data points without failure;markers with 

arrows indicate that the tests stopped without failure...............................................50

Figure 31 (a) Maxima of absolute curvature extremes and curvature amplitudes as a 

function of number of cycles to failure, and (b) maxima of absolute strain 

extremes and strain amplitudes as a function of number of cycles to failure;

markers with arrows indicate that the tests were stopped without failure.................52

Figure 32 (a) Curvature amplitudes, (b) maxima of absolute curvature extremes, and 

(c) maxima of absolute strain extremes as a function of number of cycles to 

failure; Curve-fitting was extended to include the no-failure data points; markers

wth arrows indicate that the tests were stopped without failure..............................54

Figure 33 Images of untested segment 605D1D for (a) overall axial cross section and (b) 

enlarged area A.......................................................................................................56

Figure 34 (a) Axial cross section of S2/ 605D1E, (b) fracture surface near back side of 

rod and an enlarged area, and (c) fracture surface near front side of rod and an 

enlarged area (A = delamination of circumferential hydride layers, B = spalling 

and cracking of oxide layer).....................................................................................57

Figure 35 (a) High magnification micrograph showing radial hydrides of Sample HR-1 

(H ≈ 360–400 ppm); the HBU HBR specimen was pressurized to 140 MPa at 

400 °C with five thermal cycles; (b) high magnification micrograph showing 

radial hydrides of a sample of unirradiated, hyrodgen charged zircaloy-4 (H ≈

286 ± 20 ppm); the specimen was sectioned at the mid-plane of a sample that 

was 6 inches long. ...................................................................................................58

Figure 36 (a) Moment amplitude vs. number of failure cycles, (b) strain amplitude vs. 

number of failure cycles, as function of clad oxide thickness; HR4 only has 

thermal annealing treatment without pressurization applied.....................................60

Figure 37 Metallographic examination of fuel and cladding: results of HR1, which show 

debonding phenomenon and gap formation at pellet-clad interface and high 

density of radial hydride formation in HRT treated rod. ............................................61
xi

LIST OF TABLES

Table 1 Correction Factors for LVDT Measurement Data .....................................................13

Table 2 Specifications of specimens used in hot-cell testing.................................................19

Table 3 Summary of static tests............................................................................................30

Table 4 Characteristic points and quantities based on curvature-moment curves.................33

Table 5 Comparison of flexural rigidity results between CIRFT testing and PNNL data ........35

Table 6 Summary of dynamic test results.............................................................................43

Table 7 Failure modes of rods from dynamic tests................................................................44

Table 8 Curve-fitting parameters of curvature and strain as a function of cycles* .................54

Table 9 HBR SNF CIRFT data with the associated tested rod’s oxide thickness ..................59

xiii

EXECUTIVE SUMMARY

This report describes testing designed to determine the ability of high burnup (HBU) 

(>45 GWd/MTU) spent nuclear fuel (SNF) to maintain its integrity under conditions relevant to 

storage and transportation. The project included two phases. Phase I involved testing HBU fuel 

in the as-irradiated state, in which hydrides are expected to be predominantly in the 

circumferential-axial orientation. Phase II involved testing HBU fuel segments following a 

treatment designed to reorient the hydrides in the cladding to be predominantly in the radialaxial orientation.

In 2011, Oak Ridge National Laboratory (ORNL) proposed a bending fatigue system for testing

HBU SNF rods. This system is composed of a U-frame equipped with load cells and is used to

impose pure bending loads on SNF rod specimens and measure the in-situ curvature of the 

SNF rod specimen during bending. This is accomplished using a set-up of three linear variable 

differential transformers (LVDTs). Pure bending is a condition of stress in which a bending 

moment is applied to a beam without the simultaneous presence of axial, shear, or torsional 

forces. 

On August 19, 2013, the testing system was installed in ORNL’s Irradiated Fuels Examination 

Laboratory (IFEL) hot cell and was formally named the cyclic integrated reversible-bending 

fatigue tester (CIRFT). After tuning the CIRFT and performing benchmark testing in September 

2013, Phase I testing began on HBU SNF rod segments with Zircaloy-4 (Zry-4) cladding

irradiated in the H. B. Robinson (HBR) nuclear plant. ORNL completed four static tests under 

displacement control at the rate of 0.1mm/s to a maximum displacement of 12.0 mm. In early 

November 2013, the benchmark and static test results were critically reviewed at a meeting 

between representatives from the US Nuclear Regulatory Commission (NRC) and ORNL. 

Dynamic testing was then initiated, and 16 cyclic tests were completed in the IFEL. Load ranges

varied from ±5.08 to ±35.56 N∙m; there were 12 tests with rod fracture and 4 tests without rod 

fracture. One of the cyclic tests reached 1.3 × 107

 cycles with no rod fracture and was 

terminated for practical reasons. 

Phase II testing began in 2016 using HBU HBR SNF rods with Zry-4 cladding which had been 

subjected to a hydride reorientation treatment (HRT). ORNL completed CIRFT testing on four

specimens following HRT: one in static loading, and three in dynamic loading. The samples 

subjected to HRT showed reduced lifetimes and reduced flexural rigidity when compared to asirradiated samples. 

The following observations can be made regarding static testing: 

• The static CIRFT results indicate a significant increase in the flexural rigidity of a 

fueled SNF rod compared to the flexural rigidity in a calculated response for cladding 

only. This applies to both as-irradiated and HRT SNF rods.

• The sample that was subjected to HRT and tested under a static bending load 

showed reduced flexural rigidity compared to as-irradiated samples. Nevertheless, 

materials tested in the as-irradiated and HRT states both had higher flexural rigidity 

than materials considered in the calculated response for cladding only. 

• The HBU HBR rods in the as-irradiated state survived static unidirectional bending to 

a maximum curvature of 2.2–2.5 m−1

, or a maximum moment of 85–87 N∙m. The 
xiv

maximum static unidirectional bending values were bounded by the CIRFT device’s

displacement capacity. The maximum equivalent strain was 1.2–1.3%.

• The HBU HBR rods in the as-irradiated state exhibited a multiple-stage constitutive

response, with the two linear stages followed by a nonlinear stage. The flexural

rigidity at the initial stage was 63–78 N∙m2

, corresponding to an elastic modulus of

101–125 GPa. The flexural rigidity at the second stage was 55–61 N·m2

, and the

corresponding elastic modulus was 88–97 GPa.

• Most HBU HBR rods in the as-irradiated state under static unidirectional loading

fractured at a location coincident with the pellet-to-pellet interface, as validated by

the post-test examinations showing pellet end faces in most of the fracture surfaces.

Fragmentation of the pellets also occurred to a limited degree, along with cladding

failure.

The following observations can be made regarding dynamic testing:

• The fatigue life of HBU HBR rods in the as-irradiated state in the cyclic tests

depended on the level of loading. Under loading with moments of ±8.89 to ±35.56

N∙m—namely ±0.03 to ±0.38% strain at 5 Hz—the fatigue life ranged from 5.5 × 103

to 2.3 × 106

 cycles.

• The ε-N curve of the HBU HBR rods in the as-irradiated state can be described by a

power function of y = 3.839·x−0.298, where x is the number of cycles to failure, and y is

the strain amplitude (%).

• Maxima of the curvature during dynamic tests in the as-irradiated state ranged from

±0.08 to ±0.78 m−1 at 5 Hz. The κ-N curve of the HBU HBR rods can be described by

a power function of y = 6.864·x−0.283, where x is the number of cycles to failure, and y

is the maxima of cladding tensile curvature |κ|max (m-1). A fatigue limit is likely located

between 0.08 and 0.15 m−1 if it is defined at 107

 cycles (as is typical for material

fatigue endurance limits).

• The failure of HBU HBR rods under cyclic loading often occurred near pellet-to-pellet

interfaces.

• The samples subjected to HRT showed a slightly reduced lifetime compared to asirradiated samples in dynamic testing.
xv 

ACKNOWLEDGMENTS

This research was sponsored by the US Nuclear Regulatory Commission (NRC). The authors 

would like to thank Michelle Flanagan-Bales, NRC Program Manager, for providing guidance 

and support, and Gordon Bjorkman, Bob Einziger, and Patrick Raynaud, NRC staff members, 

for providing invaluable input and comments during program development and execution. 

The authors gratefully acknowledge Oak Ridge National Laboratory program managers Bruce 

Bevard and Rob Howard for their guidance, Ting Tan and Hao Jiang for providing finite element 

method simulation support, Yong Yan for providing hydride reorientation (HR) specimens, 

Chuck Baldwin for providing post-irradiation examination support, Josh Schmidlin for performing 

fuel rod cutting and dimension measurement, Brian Woody and Scot Thurman for providing hotcell operation support, Thomas Cox, Brian Sparks, and Randy Parten for providing drawing and 

machining support, and Rose Raney for providing editorial review of this report.

xvii

LIST OF ACRONYMS, ABBREVIATIONS, AND UNITS

ANL Argonne National Laboratory

BE bonding efficiency

CFR Code of Federal Regulations

CIRFT cyclic integrated reversible-bending fatigue tester

EPRI Electric Power Research Institute

FEA finite element analysis

FEM finite element method

GWd gigawatt-day

HBR H. B. Robinson

HBU high burnup

HR hydride reorientation

HRT hydride reorientation treatment

ID inner diameter

IFEL Irradiated Fuels Examination Laboratory

LVDT linear variable differential transformer

MET metallographic examination

MOX mixed oxide

MTU metric ton of uranium

NCT normal conditions of transport

N∙m newton-meter

NRC US Nuclear Regulatory Commission 

OD outer diameter

ORNL Oak Ridge National Laboratory

PIE post-irradiation examination 

PNNL Pacific Northwest National Laboratory

PWR pressurized water reactor

SEM scanning electron microscope

SNF spent nuclear fuel

SS stainless steel

Zry-4 Zircaloy-4 

d1, d2, d3 deflection of rod at LVDT1, LVDT2, LVDT3, mm

E1, E2, E3 elastic modulus, GPa

EI, EI1, EI2, EI3 flexural rigidity, N∙m2

F load, N
xviii

h, h1, h2 sensor spacing, m

I moment of inertia, m4

Ic moment of inertia for clad, m4

Ip moment of inertia for pellet, m4

L loading arm, m

κ, κA, κB, κC curvature, m-1 

max k || maximum cladding curvature, 

M, MA, MB, MC moment, N∙m

∆M moment range, N∙m 

∆κ curvature range, m-1

σ, σA, σB, σC engineering stress, MPa

ε, εA, εB, εC engineering strain, %

ymax maximum y of cross section of rod, m
1 

1 INTRODUCTION

Current practice is for nuclear fuel to be regularly irradiated to high burnup (HBU) values (>45 

GWd/MTU). Several phenomena associated with HBU fuel have been identified, such as 

hydride reorientation (HR) that could influence cladding integrity. Increased burnup, the use of 

advanced materials, and variability in material properties may also influence fuel performance 

during transportation.1,2

As part of the planned fuel cycle, spent nuclear fuel (SNF) will be transported at least once 

before permanent disposal. Under normal conditions of transport (NCT), packages will 

experience cyclical force due to vibration. This cyclical force will be transmitted to the contents 

of the packages and SNF assemblies, resulting in localized stresses in the fuel rods. The 

response of SNF to NCT oscillation has been investigated: the US Nuclear Regulatory 

Commission (NRC) tasked Oak Ridge National Laboratory (ORNL) to investigate the bending 

stiffness and fatigue life of HBU fuel during transportation. Bending stiffness (or flexural rigidity) 

is the critical rod property that dictates the rod’s vibration response. The testing was designed to 

evaluate the fuel rod as a composite system, including fuel pellets and any pellet/cladding

bonding effects. Although the original objective was to determine the response of HBU SNF to

vibration under NCT, the stresses applied to the HBU SNF rods throughout the program can be 

applied to all transportation and storage conditions. In addition to determining the results of 

fatigue testing, another objective was to determine the effect of the fuel pellets’ presence on the 

overall behavior of the fuel rod.

Testing was conducted on HBU SNF with Zircaloy-4 (Zry-4) cladding from a pressurized water 

reactor (PWR). Five static tests were conducted using fuel rod segments from five rods with 

very similar power histories and burnups. This was followed by nineteen cyclic fatigue tests in a 

range of load amplitudes. The average burnup of the tested rod segments ranged between 

63.8–66.8 GWd/MTU. The estimated oxide layer thickness on the waterside surface was 

approximately 40–110 μm, and corresponding hydrogen contents of hydrides within the 

cladding wall were estimated to be between 360–800 wppm. In the first phase of testing, all fuel 

rod segments were tested in an as-irradiated state in which there is no oxide layer removal, with 

most hydrides in the circumferential orientation. Because both circumferential and radial 

hydrides are oriented in the planes parallel to the principal normal tensile stress during bending 

loading, SNF rod fatigue strength was not expected to be sensitive to the hydride orientation.3

However, the potential effect of radial hydrides on the bending fatigue response of the HBU fuel 

rods was investigated in Phase II of the program. 

Since the first publication of this NUREG-7198 in May 2015, several factors that influence the 

recorded curvature measurement data were identified. In this first revision, updated results from

previous CIRFT tests are reported. In addition, after the first publication, a new phase of testing 

began using samples treated to induce radial hydrides. This revision includes results of CIRFT 

tests on segments subjected to hydride reorientation treatment (HRT), broadening the 

applicability of the research results. This first revision of NUREG-7198 is entirely replaces the 

initial May 2015 publication.

3 

2 APPROACH

The cyclic integrated reversible-bending fatigue tester (CIRFT) shown in Figure 1 is the test 

hardware developed by ORNL.

4,5,6,7,8,9,10,11,12 Test protocols were designed to oscillate short 

segments of HBU SNF until fatigue failure occurred. The results of these tests were evaluated 

to gain a better understanding of the expected fatigue lifetime of SNF during NCT and the 

potential effects of phenomena such as HR. 

Figure 1 The cyclic intgrated reversible-bending fatigue tester developed by ORNL

Static tests were used to define the fuel rod system’s elastic and plastic responses. Dynamic 

testing defined the fatigue life of the fuel rod system under constant amplitude loading. The 

effects of cyclic loading were tested over a wide range of loading amplitudes to develop an 

understanding of HBU fuel fatigue life. Relations of fuel rod strain vs. the number of cycles to 

failure (ε-N), or curvature vs. the number of cycles to failure (κ-N), were developed to estimate 

the influence of imposed bending on the fatigue strength, where 

ε is the gauge strain, or strain amplitude, experienced by the SNF composite segment 

during the test, and equal to κ × ymax,

ymax is the maximum distance to the neutral axis of the test rod section and is set equal to the 

outer radius of the cladding,

N is defined as the number of recorded cycles to failure, and

κ is defined as the average curvature of the oscillating rod, assumed circular, measured 

during the test using three sensors (d1, d2, and d3), and evaluated as follows:

2/12

0

2

20 ])[(

− κ +−= ydx , (1)

)(2

2 )()( 32 21

0

ab

aba b

mm

ddmddmhmm

x −

+++−− =

,

22

1 21

0 0

hdd

x

m

y

a

 −



 



 + −−=

,
4 

where

dd 12

h

ma − =

,

dd 23

h

mb − =

, and

h is the sensor distance, 12 mm. 

Figure 2 Determination of the rod’s bending curvature by use of deflections measured at 

three points

Under the CIRFT load-controlled mode, the curvature response of a rod (Figure 2) is not 

necessarily symmetric in one cycle of loading due to potential nonsymmetry in the SNF rod, as 

well as the associated SNF system flaws and defects. Thus the curvature is described by a 

mean value of the maximum curvature, κmax, and minimum curvature, κmin, from the tension and 

compression cycle of CIRFT testing, κm: 

κm = 0.5 × (κmax + κmin), (2) 

and the maximum of absolute curvature extremes, |κ|max, is

max( ), κ max = max κκ min . (3) 

For a given specimen, the |κ|max given by Eq. (3) corresponds to the curvature that creates the 

maximum tensile stress in the cladding. Time series plots of applied bending moment and 

curvature were also generated. The uniform moment applied to the rod is derived simply from 

the following equation: 

M = F × L, (4) 

where F is the averaged value of applied bending loads registered from two load cells, and L is 

the fixed loading arm length, 101.60 mm. 

Figure 3 illustrates some of the terms used in the discussion and indicates the surfaces and 

locations involved in bonding, pellet-cladding interaction, and pellet-to-pellet pinning. The effects 


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7 

3 EXPERIMENTAL METHOD

The objective of this testing program was to evaluate the bending deformation and the 

associated fatigue mechanical properties for HBU fuel rods relevant to transportation and 

storage conditions. Currently, most mechanical testing devices require a particular design of the 

specimen such as a pre-notch or reduced gauge section to prevent failure at the grip location 

and to ensure meaningful data. For this testing program, the testing equipment was designed to 

obviate the need for the test segment to be altered by machining a gauge section. Such 

alteration could have introduced or removed localized features relevant to the bending 

deformation. Therefore, an innovative U-frame and grip design were developed, key features of 

which are discussed in this section. Cost considerations also had a significant influence on the 

development of the testing program. Hot-cell space and time are extremely limited and 

expensive, as is testing material. For this testing program, the intent was to use a small test 

segment, design a compact test device, and develop a test protocol for efficiency. Important 

features of the test segments and test protocol are discussed later in this report. Finally, the 

curvature response of the test segment was to be measured with as little interpretation as 

possible. It was not practical to weld strain gauges onto the segments or use lasers to measure 

deflection on HBU material in a hot cell. Therefore, three adjacent linear variable differential 

transformers (LVDTs) were used to measure lateral deflections of the test specimen so the 

curvature of the specimen could be calculated without interpretation or assumptions about the 

deformation of various components of the testing system. The important aspects of the three 

LVDTs and the calculation of specimen curvature are discussed in this section. 

3.1 The U-Frame Setup

The integrated CIRFT U-frame is shown in Figure 4. The U-frame includes two rigid arms, 

connecting plates, and universal testing machine links. The rod specimen is coupled to the rigid 

arms through two specially designed grips. The U-frame setup is oriented in a horizontal plane 

and is driven by Bose dual linear electromagnetic motors. With help from the coupling, linear 

motions applied at the loading points of the rigid arms are converted into bending moments 

exerted on the rod. The dual linear motor (model LM2) test bench has a maximum load capacity 

of ±3,000 N for each motor and a total maximum stroke of ±25.6 mm. 


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11

control and the curvature resulting from the applied load is measured directly on the sample, the 

properties of the compliant layer do not need to be accounted for when determining the loadcurvature response. 

Epoxy casting and durability has been demonstrated over time in the hot cell environment with a 

series of tests of samples exposed to high radiation in the hot-cell. A vise mold was designed 

and built to cast the epoxy, and procedures were developed to cast the epoxy and mount the 

rigid sleeves into the rod specimen.5,6 Several modifications were made to adapt the vise mold 

to hot cell testing.10 Images in Figure 8(a) show different views of the vise mold, and those in 

Figure 8 (b) and (c) illustrate the process of epoxy casting in the hot cell. 


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13

3.3 LVDT Measurements and Corrections

Theoretically, the bending radius and maximum strain of a rod can be estimated based on the 

traveling displacement at the loading points of the rigid arm. However, the displacement 

measured includes the contribution of the compliant layers and depends on the materials used 

in them and the level of loading. To take these factors into consideration, direct measurement of 

the specimen displacement using LVDTs at three adjacent points along the rod method was 

proposed14 and implemented to evaluate the curvature of a bending rod in this study,

5,8

 as 

established in Section 2. 

A number of factors influence the translation of LVDT measurements into curvature 

measurement data. Since the first publication of NUREG-7198 in May 2015, some new factors 

were identified. The primary factors accounted for in the CIRFT data processing procedure are 

(1) load cell and LVDT measurement data smoothing, (2) LVDT probe contact point/sensor 

spacing corrections, and (3) LVDT system dynamic vibration correction factor. Table 1

summarizes the primary correction factors and their relative impact on the curvature 

measurement data. Additional discussion is provided below on correcting for the probe width 

and dynamic vibration effects.

Table 1 Correction Factors for LVDT Measurement Data

Origin of Correction Δh

Impact 

on 

results

Correction 

made to 

data in 

Rev 0

Correction 

made to 

data in 

Rev 1

Probe width effect as 

evaluated by tests with strain 

gauges and half-gauge 

length methodology

2.5

50% (+)

37% (-) 

No Yes

Data smoothing No No No Yes

LVDT probe dynamic 

vibration adjustment 

procedure

No 18% No Yes

3.3.1 Correction for Disk-Type LVDT Head Bias

The arrangement of the three LVDTs and their installation in the CIRFT can be seen in 

Figures 2 and 7. The LVDTs measure the gauge strain as defined in Figure 3. A series of static 

tests was performed to confirm the measurement method by using multiple surrogate rods 

composed of various materials. The curvature obtained by the LVDT measurements has been 

shown to correlate well with the calculated flexural rigidity of the specimen based on the 

estimated moment of inertia and Young’s modulus for various materials. However, the offset in 

the contact position of the disk-shaped LVDT probes with the rod (Figure 9) introduced a bias in 

the curvature measurements. 


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16

so the associated gauge strain is:

ε = κ × ymax (11) 

where ymax is the radius of the clad’s outer surface. The three LVDT-based curvature curves 

obtained with a disk probe are presented in Figure 11 with and without the correction applied. 

Figure 11 Curvature measurements of polycarbonate rod based on M-based estimate

and disk type probe, 2.50 mm sensor spacing correction

Implementation of the correction procedure in the polycarbonate rod example suggested that a 

∆h of 2.50 mm is needed. The corrected curvature showed a good correspondence with the 

M-based curvature according to Eq. (10). In the monotonic test, ∆h = 2.50 mm obtained from 

the calibration can be used for spacing correction in data analysis. In the reversed cyclic 

bending, the ∆h is applied directly using Eq. (9) because the peak and valley pair is available in 

the data block. This is preferred since, as discussed further in Section 4.2, with the same 

loading in both directions, the flexure/ curvature of an irradiated fuel rod can be asymmetric due 

to local bonding and pellet-cladding interface effects. The experimental details and results 

obtained in the out-of-cell study using surrogate rods can be found in Appendix G. All data 

reported here have been adjusted for LVDT head bias.

3.3.2 Curvature Adjustment for LVDT Stem Dynamic Stability

To verify the curvatures calculated from the three LVDT readings under dynamic loading, an 

out-of-cell benchmark exercise was conducted using a strain gauge fixed on the sample. A 

strain gauge (EA-06-250BF-350, Appendix G) was mounted on calibration rod SS30402 and 

tested under load control at varying frequencies. Correction was applied to the probe sensor 

spacing as shown in Figure 12. 
17

Figure 12 Variations of (a) peak strain and (b) valley strain as a function of driving 

frequeny. At 5 Hz, corrected strain level is about 18% higher than strain gauge 

estimate for both peak and valley strain plots

Peak and valley strains based on the corrected exhibit have demonstrated a trend similar to the 

response of strain gauge as a function of driving frequency as shown in Figure 12. The level of 

the three-LVDT-based strains is usually higher than those of the strain gauge in both peak and 

valley directions, which may be induced by the dynamic response of LVDT. The overestimate is 

~18% at 5Hz. Therefore, all curvatures and gauge strains reported here are reduced by 18% in 

addition to the LVDT bias correction to take into account the variance induced by LVDTs stem 

dynamic instability. 

3.4 Test Segment Selection and Preparation

The Zry-4 15 × 15 cladding (1.41 wt. % Sn, 0.22 Fe, 0.12 Cr, and 0.14 O) tested in this study 

was irradiated for seven cycles in the HBR Unit 2 PWR. The rod-average fuel burnup was 

67 GWd/MTU. The test segments for Phases I and II were taken from rods at E04, E14, F07, 

G10 and R05 locations in assembly S-15H. The irradiation history of these rods and their 

associated details can be found in the literature.1 The specimen cutting plan and the specimen 

locations in the father rods can be found in Appendix A. According to the cutting plan presented 

in Appendix A, 6-inch segments were cut. Figure 13 shows one segment (608C4B) taken from 

the father rod G10, along with the tube used for specimen storage. The rod diameter of the fuel 

segments was measured using a dedicated setup in the hot cell.17 The measurement details are 

provided in Appendix B. Readings were taken at each half inch from one end of the rod and at 

three angles (0, 120, and 240°) on cross sections at each reading distance. The mean value 

and standard deviation of the diameter readings were found to be 10.763 mm (0.4237 in.) and 

(a) (b)
18

0.003 mm (0.0012 in.), respectively. It can be seen that the calculated mean value is close to 

the nominal diameter of the fuel cladding at HBR.

1

Figure 13 Segment 608C4B for DL2 taken from rod G10, with the storage tube

Twenty CIRFT tests were completed in Phase I on the HBR specimens in the as-irradiated 

state, including four static tests (S1–S4) and sixteen dynamic tests (D0–D15). Four CIRFT tests 

were completed in Phase II on the HBR specimens following HRT. These consisted of one 

static test (HR2) and three dynamic tests (HR1, HR3, and HR4). A summary of specimen 

information for hot cell testing is provided in Table 2. 
19

Table 2 Specifications of specimens used in hot-cell testing

Specimen 

label 

ORNL 

inventory label

(“Seg. ID”)

Span Burnup1 

(GWd/MTU)

Estimated 

oxide1 (μm)

Estimated 

hydrogen of span*

(wppm)

Outer 

diameter 

(mm)

Static Tests

S1 606C3C 4 66.8 70–100 550–750 10.748

S2 605D1E 2–3 66.5 40–70 360–550 10.746

S3 609C5 4 66.5 70–100 550–750 10.736

S4 609C6 4 66.5 70–100 550–750 10.740

HR2 607D4A 2 63.8 40-44 360-400 10.720

Dynamic Tests

D0 605D1F 2–3 66.5 40–70 360–550 10.746

D1 607C4B 4 63.8 70–100 550–750 10.834

D2 608C4B 4 63.8 70–100 550–750 10.781

D3 605C10A 4 66.5 70–100 550–750 10.774

D4 605D1C 2–3 66.5 40–70 360–550 10.750

D5 605D1B 2–3 66.5 40–70 360–550 10.771

D6 609C4 4 66.5 70–100 550–750 10.748

D7 609C3 4 66.5 70–100 550–750 10.777

D8 606C3E 4 66.8 70–100 550–750 10.739

D9 609C7 4 66.5 70–100 550–750 10.738

D10 606C3A 4 66.8 70–100 550–750 10.759

D11 607C4A 4 63.8 70–100 550–750 10.832

D12 608C4A 4 63.8 70–100 550–750 10.821

D13 606B3E 5 66.5 100–110 750–800 10.765

D14 606B3D 5 66.5 100–110 750–800 10.765

D15 606B3C 5 66.5 100–110 750–800 10.768

HR1 607D4C 2 63.8 40–44 360–400 10.760

HR3 608D4A 2 63.8 40–44 360–400 10.720

HR4 608D4C 2 63.8 40–44 360–400 Not measured

*From References 18,19

3.4.1 Hydride Reorientation Treatment (HRT) Procedure

Phase II involved testing HBU SNF segments following a treatment designed to reorient the 

hydrides in the cladding to be predominantly in the radial-axial orientation. The HBU SNF 

samples for the HRT tests were prepared in the Irradiated Fuel Examination Laboratory (IFEL). 

After samples were sectioned into pieces 6 inches long, the outer surface oxide layer, fuel, and 

inner surface oxide were removed from both ends to a depth of approximately 0.5 inch. End 

caps were welded at both ends, and the HBR fuel specimens were then pressurized with argon 

gas. Thermal cycling was performed to maximize the radial hydrides (see Appendix H for 

additional information on the extent of radial hydrides induced in a range of HRT conditions). 

The specimens were placed in a holder in the furnace for heating. The specimen was 

pressurized to induce a maximum hoop stress of 140 MPa at a target temperature of 400 °C. 

The test specimens were held at 400 °C and 140 MPa for 3 hours, cooled at 1 °C/min to 170 °C, 

and then heated at 1 °C/min to the hold temperature of 400 °C (see Figure 14). This thermal 


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22

amplitudes were chosen based on the static testing results and the desire to provide 

a wider range of fatigue life data.

2. Perform the dynamic cyclic test under constant load control using a sine wave input

in reverse bending mode.

3. Monitor the SNF fatigue evolution at defined intervals: perform static measurements

of the rod deformation at the end of each target cycle with a frequency of 0.05 Hz

and reduced loading amplitude relative to the dynamic loading amplitude under

displacement control. Using the reduced loading amplitude for the periodic

measurements ensures there is no fatigue damage accumulated during the

measurement procedure.

4. Stop the dynamic test when fracture is detected or when the preselected number of

cycles is reached.

5. Weigh any fuel fragments that fall out of the fracture.

3.6 Data Processing 

Online monitoring data are converted into the applied moment and curvature based on the load 

channel (load1 and load2) information, the loading arm length (101.60 mm), and LVDT data 

(LVDT1, 2, and 3). This information is used to generate the time series plots of moment and 

curvature and the moment-curvature hysteresis loops. These loops are those illustrated in

Figure 16 for specimen D1. The data are processed using the approach defined in Section 2 for 

the applied deflection, moment, and strain. 


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24

compressive stress regions shown in Figure 16. At the compressive cladding stress region, the 

fuel reinforcement remains intact in the form of the fuel pellet pinning effect. This results in a 

shift of neutral axis in an SNF system under reversal bending. 

This phenomenon was also observed from the CIRFT test online monitoring data, in which the 

tensile cladding stress stage shows a higher curvature reading than the compressive stress 

stage, as shown in Figure 26(d). The EI values also change accordingly in each moment 

reversal cycle, as shown in the moment-curvature plot of Figure 16 above, where the EI (the 

slope of moment-curvature) in the clad compression cycle is higher than that of the clad tension 

cycle. Furthermore, a detailed three-dimensional finite element analysis with a 6.25 N∙m uniform 

moment also predicted that the localized strain in the cladding at the pellet-to pellet cladding

interface region is about three to four times (depending on the interface cohesive bond 

parameters)21 that of the global strain for the tensile cladding stress region. The associated 

localized stress is about 2.6 times the average global tensile cladding stress. Another

complication is that the neutral axis of the SNF rod will no longer reside in the geometric center 

of the SNF system, and the EI value will shift alternatively around the geometry center under 

cyclic loading reversals.

21,22 Therefore, the conventional approach, as stated in Eqs. (9) and 

(10) from a global M-κ consideration, is no longer valid for describing the cladding failure 

mechanism associated with localized flexural rigidity degradation at the rod’s pellet-to-pellet 

interface. Conducting detailed localized stress-strain evaluation at the pellet-to-pellet interface 

using global measurements is beyond the current scope of the project. Therefore, only the 

mechanical response derived from direct measurements (the LVDTs) will be used in this study 

and discussion. 

Generally, the moment range (∆M), curvature range (∆κ), and flexural rigidity EI are used to 

characterize the mechanical properties of the fuel rod. These are defined as: 

∆M = Mmax − Mmin, (14)

∆κ = κmax − κmin, and

EI = ∆M / ∆κ, 

where the subscripts max and min represent the maximum and minimum waveforms in each 

loading cycle (Figure 16).

Under a load controlling mode, the curvature response of a rod is not necessarily symmetric in 

one cycle of loading (Section 4.1). The offset of the M-κ loop on the κ axis with respect to the 

origin can be described by a mean value of curvatures, κm: 

κm = 0.5 × (κmax + κmin), (15) 

and the maximum of absolute curvature extremes, |κ|max, is

max( ), κ max = max κκ min . (16) 

For a given specimen, the |κ|max given by Eq. (16) corresponds to the curvature that creates the 

maximum tensile stress in the cladding. The resistance force of the CIRFT system may be 

significant, depending on the amplitude of the rigid arm movement. The CIRFT system static 

resistance force was measured at different displacement levels without a specimen loaded. The 

net applied load at each displacement level was estimated by subtracting the measured 

resistance from the applied static test load shown in Figure 17. 
25

Figure 17 Empty-run conducted without specimen installed, for which the load and 

displacement were plotted; associated with static test on D1 

The CIRFT system resistance was tested and evaluated by using empty runs in which the 

specimen was not loaded on the testing machine. These empty runs were conducted in the 

specified period, or whenever the calibration is needed.

0 50 100 150 200

-2

0

2

4

6

8

10

12

14

Time (sec)

Disp (mm)

test001.dat

disp1

disp2

0 50 100 150 200

-5

0

5

10

15

20

25

Time (sec)

Relative Disp (mm)

test001.dat

0 50 100 150 200

-20

-10

0

10

20

30

40

50

60

70

80

Time (sec)

Load (N)

test001.dat

load1

load2

0 50 100 150 200

-20

-10

0

10

20

30

40

50

60

70

Time (sec)

Average Load (N)

test001.dat
26

4 RESULTS

4.1 Static Tests

Three specimens were tested under static loading in the hot cell for test system benchmarking. 

Five static tests were subsequently conducted: four on HBU segments as irradiated, and one on 

an HBU segment following HRT. This chapter focuses on these five static tests, while Appendix 

C covers the static benchmarking tests. 

S1 (606C3C, 66.8 GWd/MTU burnup, 70–100 μm oxide layer) survived four static loading/ 

unloading cycles of unidirectional bending without failure. The moment-curvature curve for the 

initial cycle is presented in Figure 18(a). (The curves of subsequent cycles overlay one another

to a great extent and for clarity are not plotted.) The maximum moment achieved in this test was 

85.5 N·m, corresponding to a maximum curvature of 2.2 m-1. The slope changes of the linear 

portion of moment-curvature curves could be seen near moment levels 12 and 59 N∙m. Due to 

pellet end dish-in design, the pellet-to-pellet interface’s bonding strength would be fairly weak, 

and the pellet-to-pellet debonding would likely occur in the early phase of the bend loading. The 

slope-changed points indicate that the flexural rigidity changed in the SNF rod system within the 

linear elastic range. 

Figure 18 Curve of moment versus curvature 

Based on a static loading cycle in which the maximum relative displacement was 24 mm; rates 

at the loading points of the U-frame were set at 0.1 and 0.2 mm/s for loading and unloading, 

respectively.

Since the machine’s capacity was reached and no failure occurred to the specimen after four 

unidirectional loadings, a post-static dynamic test was conducted with reversible bending under 

±25.40 to 30.48 N∙m at 5 Hz. As mentioned above, the purpose of this procedure was to test the 

specimen to failure for further postmortem examination. The rod failed after approximately

1.4 × 104

 cycles. The failure occurred near motor 1 (right side) with a fuel release of 0.9 g. 

Optical images of the fracture surfaces are shown in Figure 19. The pellet end-face can be 

identified from the haze zone characterized by a brown-blue color indicated by the triangular 

marker in Figure 19(a). This indicates that the rod failure may have initiated at the pellet-topellet interface. At the same time, a serrated shear fracture surface of the cladding was 


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One HBU segment following HRT, HR2, was initially tested under unidirectional bending six 

times, and then it was tested under the reverse 4 more times, and no failure was observed

before the machine stoke capacity was reached. HR2 was further tested under 16.26Nm and 

failed at 9.47 × 103

 cycles. The details of the HRT procedure are provided in Appendix H.

In the static tests described above, loading/unloading cycles did not result in apparent or 

catastrophic failure of specimens because of the limited stroke of the CIRFT machine. 

(Measurement details of all static tests can be found in Appendix D.) However, failure did occur 

in the initial loading cycle of two benchmarking tests on HBU SNF fueled specimens, Scal1 and 

Scal2. The peak moment of Scal1 was 78.2 N·m, and the peak moment of Scal2 was 80.7 N·m. 

Details of these benchmarking tests can be found in Appendix C. Therefore, it is reasonable to 

assume that the maximum bending moment for static tests S1–S4 were not significantly greater 

than 80 N·m.

A summary of the moment-curvature curves is shown in Figure 20 for the five static tests. A 

summary of static test results is provided in Table 3. Several important observations can be 

made with respect to the responses of the test specimens during the first loading cycle. 

• The results from the four Phase I static test results (as-irradiated segments) all show

relatively similar elastic behavior.

• The initial moment-curvature response of Phase I static tests is characterized by a

flexural rigidity of 63–79 N∙m2

.

• There is a change in the flexural rigidity in the Phase I static tests at the moment of

12–21 N∙m. This change is quantified in Section 4.1.1.

• After this change, the moment-curvature response of the Phase I static tests is

characterized by a decreased flexural rigidity of 55–61 N∙m2.

• Both Phase I and Phase II static test segments experienced 1.2–1.4% strain without

failure.

• The unloading flexural rigidity of the Phase I static tests is close to the loading

flexural rigidity in the second stage after the moment reaches 42 to 53 N∙m.

• The single Phase II static test, HR2, showed a similar loading behavior, but with

lower flexural rigidity compared to that of as-irradiated specimens under similar

curvature. Furthermore, the unloading of the HR2 sample appeared to be a curve

with a varying slope. A significant load drop was observed shortly after unloading.

• Overall, all tested rods failed within the gauge sections. The majority of the rod

fractures involved the pellet-to-pellet interfaces. S3 was exceptional in that the failure

location seemed away from the pellet-to-pellet interface, as no end face of the pellet

can be identified from the fracture surface of the fuel.
29

Figure 20 (a) Moment-curvature curves measurements and (b) stress and strain curves in 

static tests 

(a) 

(b) 


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4.1.1 Characteristics of Moment-Curvature Curve

The moment-curvature responses of the four HBU SNF as-irradiated specimens were similar. 

They are characterized by two distinct linear responses, EI1 and EI2, followed by a nonlinear 

response during the loading and a linear response upon unloading, EI3. It was observed that 

when reloaded, the moment-curvature response of the rod followed the same slope of the linear 

unloading curve and proceeded with nonlinear response after passing the previous maximum 

load. Such loading/unloading/reloading responses can be found in many mechanical systems 

where irreversible changes occur during loading.

8An effort was made to characterize the 

moment-curvature response based on characteristic points to facilitate understanding of the test 

results. The values for EI1, EI2, and EI3 were obtained, corresponding to the slopes of the first 

and second linear regions and of the unloading region, by using curve fitting with the first order 

polynomial (Figure 21). 

Figure 21 Characteristic points of moment-curvature curve.

The characteristic curvatures and moments at the slopes’ changed points A and B were then 

identified. In addition, the moment at point C corresponding to a 0.37 m−1 irreversible curvature, 

or 0.2% equivalent plastic strain, was found by using a line with the same slope as that of 

unloading, EI3, and horizontal axis intercept 0.37 m−1

. The quantities corresponding to points A, 

B, and C are designated by κA, κB, κC, and MA, MB, and MC. The results are summarized in 

Table 4. For as-irradiated rods, the points are generally EI1 > EI2 > EI3, with a marginal 

difference between the latter two. However, for HR2, the linear stages are not defined as well as 

those in as-irradiated specimens and EI1 > EI3 > EI2.
33 

Table 4 Characteristic points and quantities based on curvature-moment curves

Spec label Segment ID EI1 EI2 EI3 κA κB κC κD MA MB MC MD N∙m2 N∙m2 N∙m2 m−1 m−1 m−1 m−1 N∙m N∙m N∙m N∙mS1 606C3C 78.655 57.33 51.027 0.202 0.968 2.009 2.166 16.695 60.599 83.595 85.413S2 605D1E 73.016 60.848 52.699 0.32 1.009 2.001 2.154 20.18 62.133 85.914 87.294S3 609C5 71.517 59.369 47.101 0.311 0.933 2.149 2.308 22.338 59.288 83.728 85.235S4 609C6 63.117 54.849 41.704 0.503 0.862 2.329 2.507 28.54 48.244 81.656 85.02As-irradiated Avg. 71.576 58.099 48.133 0.334 0.943 2.122 2.284 21.938 57.566 83.723 85.741As-irradiated Std. Dev. 6.422 2.603 4.886 0.125 0.062 0.154 0.164 4.977 6.322 1.741 1.048HR2 607D4A 62.769 41.517 43.333 0.487 1.007 1.585 2.158 30.301 51.884 66.809 79.606
34

The HR2 sample had a lower flexural rigidity than the as-irradiated test segments. While HR2 

might be expected to survive larger deformation before failure than the as-irradiated segments, 

limitations of the CIRFT test device indicate that this could not be confirmed. Nevertheless, the 

flexural rigidity is consistently reduced under the similar curvature, and the maximum moment 

level of HR2 was 80 Nm instead of 86 Nm for the as-irradiated condition. 

4.1.2 Comparison of Static Results with PNNL Cladding Data

To investigate the contribution of fuel pellets in the fuel rod structure, analysis was performed to 

compare the measured static results to predicted values considering cladding alone. The 

cladding properties used for this analysis were obtained from a database maintained by Pacific 

Northwest National Laboratory (PNNL).23 The database allows the user to specify the cladding 

type, temperature, fluence and cold work of the cladding of interest. For this study, the following 

values were specified:

• Cladding – Zry-4

• Temperature – 75 °F

• Fluence – 12 × 1025 n/m2

• Cold work – 0.5

• Cladding ID - 0.364 inches

• Cladding OD – 0.423 inches

• Cladding Thickness – 0.0295 inches

• Calculated cladding moment of inertia (Ic) – 0.0114 in.4

To estimate the contribution of the fuel pellets to the composite bending response, the 

comparison based on CIRFT M-κ data was completed as described below. 

Figure 22 Comparison of CIRFT global data with theoretical PNNL moment-curvature 

derived from PNNL cladding only stress-strain data
35

For comparison, the PNNL data (σ−ε curve) were converted to a moment-curvature curve

(shown in Figure 22 as “PNNL”) using the following equations, with the consideration of stress 

distribution appropriate for a thin-walled tube under uniform bending,

M = σ·I/ymax, (17) 

and 

κ = ε/ymax, (18) 

where I = Ic, and other quantities have the same meanings as in Eqs. (1)–(12). The hypothetical 

cladding only moment-curvature response was plotted, together with the moment-curvature 

response of the high burnup fuel rod system, as shown in Figure 22, where data for the SNF rod 

show much higher flexural rigidity compared to that of PNNL data with cladding alone. A 

comparison of CIRFT testing results with cladding-only rigidity based on PNNL data is given in

Table 5. 

Table 5 Comparison of flexural rigidity results between CIRFT testing and PNNL data

EI1 (N⋅M2) EI2 (N⋅M2) EI3 (N⋅M2)

As-irradiated 71.576 58.099 48.133

HR2 62.769 41.517 43.333

PNNL data 26.933

The data comparison shows a distinct difference from that of the HBU fuel rod systems. This 

difference was observed in the slope of the PNNL data. The slopes of both the as-irradiated and 

HR-treated HBU fuel rod systems are greater than those shown in the PNNL data. By focusing 

on the initial slope between 0–15 N∙m, it was seen that the slope (flexural rigidity) of HBU fuel

rods was approximately twice that of the PNNL data for cladding alone. 

The increase of SNF rod system stiffness is attributed to the fuel pellet’s mechanical properties 

and moment of inertia. However, the measured flexural rigidity of a fuel rod system is much less 

than that estimated from a direct summation of EcIc + EpIp , assuming a perfect mechanical bond 

between the fuel and the cladding, as well as perfect bonding between the pellet-to-pellet 

interfaces and a single pellet. However, bonding at fuel-cladding and fuel pellet-to-pellet 

interfaces is not perfect, and it evolves during the test. The initial fuel-clad gap size also 

influences the system’s response. Furthermore, due to segmental pellet structure, numerous 

stress riser sites within cladding at pellet-to-pellet interfaces in an SNF rod are created. In 

general, the stress riser sites will reduce the lifetime of the structure as compared to the very 

same system without stress riser sites or discontinuous materials interfaces. The stress riser 

effect due to an HBU rod segment structure was further validated from the dynamic testing, 

where the CIRFT test specimens are all failed at pellet-to-pellet interfaces. Moreover, the 

intensity of the stress riser or the pellet-clad interaction is strongly dependent on the loading 

intensity. Thus, at a low loading level, pellet-to-pellet interface stress risers are expected to be 

small, so the cladding alone likely dictates the SNF rod’s composite flexural response.

All these factors affect the degree to which the presence of fuel results in an increase in the

rod’s flexural rigidity relative to cladding alone. This research effort did not attempt to account 

for and quantify each of these factors, so the preliminary conclusion based on CIRFT SNF static 

bending testing is that flexural rigidity was approximately twice what would be obtained if 

cladding properties alone were used to predict behavior. 
36

4.2 Dynamic Testing

Nineteen dynamic tests were conducted on the HBU HBR fuel in the hot cell, including tests on 

16 as-irradiated (one for benchmark testing and fifteen according to dynamic testing plan) and 3 

HRT fuel rod specimens. Input amplitudes varied from ±5.08 to ±35.56 N∙m; 12 tests were 

completed with specimen failure and 4 without failure. The fatigue life ranged from 5.5 × 103 to 

2.3 × 106

 cycles. The tests without failure were conducted under ±5.08 to ±8.89 N∙m, with the 

accumulated cycles beyond 6.4 × 106

. Details of the dynamic tests can be found in Appendix E; 

only typical results are described in this section. 

The test on D1 (607C4B, 63.8 GWd/MTU burnup, 70–100 μm oxide layer) was conducted under 

±15.24 N∙m, 5 Hz. A fatigue life of 1.1 × 105 cycles was obtained, and less than 1 g of fuel 

fragments was collected at the failure site. 

For each cyclic fatigue test, periodic quasi-static measurements of rod deformation were 

conducted at a reduced load at the target intervals described in Section 3.5.2. Time series of 

moment and curvature and moment-curvature loops obtained at the first cycle and after 

approximately 11,000 cycles are shown in Figure 23. 
37

Figure 23 Moment and curvature as a function of time and moment-curvature loops 

based on measurements when (a) N=1 and (b) N = 111,000 cycles for D1 

(607C4B); measurements were made with 0.8 and 1.2 mm relative 

displacements; Nf = 1.1 × 105 cycles under ±15.24 N∙m, 5 Hz. Fuel particles 

collected <1.0 g

Comparing the moment-curvature plots in Figures 23(a) and (b), the loops became slender with 

accumulated cycles. The relations of moment-range versus curvature-range and flexural rigidity 

are illustrated in Figure 24. 

(a) (b)


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online measurement of the cyclic test. This will ensure that the measurement process did not 

affect the dynamic cycle data. 

In general, the flexural rigidity tends to increase with decreasing curvature. This is likely due to a 

better interface bond at a relatively low load resulting in less stiffness reduction. Overall, a 

stable rod response was exhibited before the final failure. The curvature history shown in 

Figure 26(d) clearly indicated a non-symmetric deformation under reversal loading where the 

cladding tension site has much higher deformation (about 1.6 times) compared to that at the 

cladding compression site. This phenomenon could be the consequence of debonding at the 

pellet-to-pellet interface and the stress riser occurring at the pellet-to-pellet interface region at 

the tension side of the cladding, in addition to sensor probe sensitivity.

In general, the flexural hysteresis history shown in Figure 26(f) remains quite uniform throughout 

the reversal bending test, except with slightly reduction at higher cycles. Similar flexural 

hysteresis reduction at higher cycle was also observed from the periodic measurement data 

shown in Figure 23; at 1.1 × 105 cycle the moment-curvature loading and unloading curves had 

smaller loop compared to that of 26th cycle. Under relatively low loading amplitude of dynamic 

testing, the clad or fuel pellet would be mainly under linear elastic behavior. The root cause of 

hysteresis energy dissipation under cyclic loading could be the system’s nonlinear response 

associated with segment pellets induced by the stress riser at interface regions, as well as 

uncertainly resulting from sensor probe sensitivity.

Moment and curvature time history based on online monitoring data collected during dynamic 

tests at 26 and 1.1 × 105 cycles are shown in Figure 27. The asymmetrical curvature 

deformation mode of the tested rod at 26th cycle and 1.1 × 105 cycle was also observed.
41

Figure 27 Moment and curvature as a function of time at (a) 26 and (b) 1.10 × 105 cycles; 

results based on online monitoring; the initial unsymmetrical curvature 

deformation mode of the tested rod was stabilized and replaced with a 

symmetrical curvature deformation mode at higher cycles

The end faces of the two neighboring pellets were essentially clean. Both stressed cylinder 

surfaces of the rod were found covered with equally spaced circumferential cracks throughout 

the gauge section. Spalling only occurred on the local area near the fracture. The degree of 

damage shown on both stressed cylinder surfaces of specimen D1 is different from the damage

in test specimens undergoing only unidirectional bending. The large deformation arising from 

the unidirectional bending can produce a greater extent of damage or spalling on one side of the 

test specimen, with much less spalling on the other side of the specimen (e.g., S3 [609C5] in 

Table 3). 

The failure in test D1 was observed in the gauge section near motor 2 (left side of the U-frame 

setup). The failure occurred at the pellet-to-pellet interface, as illustrated in Figure 28. 

(a) (b)


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Table 6 Summary of dynamic test results

Spec 

label

Segment

ID

Load 

amp.

Moment 

amp.

Number of 

cycles Failure* κa |κ|max σa εa |ε|max

N N∙m m-1 m−1 MPa % %

D0 605D1F 250 24.068 2.50E+04 Yes 0.439 0.444 206.109 0.236 0.239

D1 607C4B 150 14.107 1.10E+05 Yes 0.215 0.24 117.26 0.117 0.13

D2 608C4B 50 4.207 6.40E+06 No 0.046 0.067 35.496 0.025 0.036

D3 605C10A 100 9.17 1.00E+06 Yes 0.125 0.171 77.938 0.067 0.092

D4 605D1C 75 6.726 1.10E+07 No 0.089 0.12 57.596 0.048 0.065

D5 605D1B 90 8.201 2.30E+06 Yes 0.114 0.123 69.706 0.061 0.066

D6 609C4 125 11.624 2.50E+05 Yes 0.205 0.218 99.546 0.11 0.117

D7 609C3 200 18.923 6.50E+04 Yes 0.351 0.37 160.835 0.189 0.199

D8 606C3E 87.5 7.743 1.28E+07 No 0.107 0.118 66.309 0.057 0.063

D9 609C7 350 33.667 7.10E+03 Yes 0.576 0.624 288.308 0.31 0.335

D10 606C3A 125 11.552 1.80E+05 Yes 0.174 0.213 98.185 0.094 0.115

D11 607C4A 300 29.021 5.50E+03 Yes 0.469 0.564 241.223 0.254 0.306

D12 608C4A 110 9.986 3.86E+05 Yes 0.144 0.171 83.617 0.078 0.092

D13 606B3E 135 12.551 1.29E+05 Yes 0.151 0.199 106.677 0.081 0.107

D14 606B3D 87.5 7.842 2.74E+05 Yes 0.112 0.135 66.652 0.06 0.073

D15 606B3C 75 6.639 2.24E+07 No 0.087 0.125 56.426 0.047 0.067

HR1 607D4C 150 15.152 4.19E+04 Yes 0.424 0.433 128.788 0.228 0.233

HR3 608D4A 100 8.982 2.44E+05 Yes 0.219 0.233 76.342 0.118 0.125

HR4 608D4C 160 14.759 5.47E+04 Yes 0.323 0.344 125.449 0.174 0.185

Optical microscopy was conducted on the fracture surfaces of selected test specimens. Among 

the seven specimens examined, at least five failures were determined to be dominated by the 

pellet-to-pellet interfaces (PPIs). Table 7 shows the failure modes of rods from the dynamic

tests. Most of the SNF rod failures initiated and propagated through the pellet-to-pellet clad 

interface region in a typical tensile failure mode except for the HR1 tested sample. The HR1 

fractured sample showed that the crack initiated at a pellet-to-pellet interface, but the crack 

propagation was in a mixed-mode (tensile + shear) failure mechanism. The postmortem 

examination of HR1 fractured sample also indicated that a large area of the pellet-cladding bond 

was disrupted during testing or never formed during operation (see the lateral view provided in 

Table 7 of item HR1 607D4C).
Table 7 Failure modes of rods from dynamic tests

Spec label

Applied 

moment 

amplitude

Average 

curvature 

(κa)

Number of 

cycles to failure Cross section Lateral view

N∙m m−1

D0, 605D1F

24.068 0.439 2.50 × 104

D1, 607C4B

14.107 0.215 1.10 × 105

44
Table 7 Failure modes of rods from dynamic tests (continued)

Spec label

Applied 

moment 

amplitude

Average 

curvature 

(κa) Number of cycles to failure Cross section Lateral view

N∙m m−1

D2, 608C4B 4.207 0.046 6.40 × 106 without 

failure; follow-up 35.56 N∙m, 5 Hz with 1.8 × 103 cycles to 

failure

D3, 605C10A

9.170 0.125 1.00 × 106

45
Table 7 Failure modes of rods from dynamic tests (continued)

Spec label

Moment 

amplitude ∆κ Number of cycles to failure Cross section Lateral view N∙m m−1

D5, 605D1B 8.201 0.114 2.30 × 106D6, 609C4 11.624 0.205 2.50 × 105D7, 609C3 18.923 0.351 6.50 × 104

46
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Table 7 Failure modes of rods from dynamic tests (continued)

Spec label

Moment 

amplitude

∆

κ Number of cycles 

to failure Cross section Lateral view

N∙m

m−1 

HR4, 608D4C

14.759 0.323 5.47 × 104

48


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50

Figure 30 Equivalent strain amplitudes (∆ε/2) as a function of number of cycles to failure

with curve fitting extended to include data points without failure; markers with 

arrows indicate that tests stopped without failure

Several important observations can be made with respect to the dynamic test results. 

• In both as-irradiated and HRT specimens that failed during testing, the fatigue life

decreases with increasing strain amplitude, as expected.

• An endurance limit for as-irradiated HBU rods is likely between 0.03–0.05% strain, if

it is defined at 107

 cycles. Below this limit, there does not appear to be a number of

cycles that will cause failure.

• There was a large variation of hydrogen content in the cladding for both as-irradiated

and HRT test specimens used in the dynamic tests. However, the contribution of

hydrogen content to the number of cycles to failure appears to be secondary to the

effect of imposed loading amplitudes.

• When reviewing the fatigue life as a function of the calculated gauge strain

amplitude, HRT rods showed extended life relative to the as-irradiated rods, but

when reviewing the fatigue life as a function of the applied moment amplitude, the

HRT rods showed reduced life relative to as-irradiated rods. Further discussion on

this subject is provided in Section 4.5, “HRT CIRFT Test Data Evaluation.”

4.2.1 κ-N Curve

The gauge strain amplitude was used to describe the loading condition imposed on the rods. 

The use of equivalent strain provides an expedient means to estimate the global response of 

the tested rod. However, the more elaborated localized deformation mechanism involved with 

pellet-to-pellet interfaces cannot be readily represented by this simplified approach. To provide 

a more accurate representation of a fatigue lifetime estimate (in addition to the curvature range), 

a new parameter of maximum curvature was defined. Due to dishing on the pellet end and the 

existence of an oxide contact interface, HBU fuel is expected to have a weak interface bond at 
51

the pellet-to-pellet interface. This theory is supported by CIRFT test results, in which the flexural 

rigidity dropped off quickly in the initial cycles of testing, and a nonsymmetric curvature 

response was revealed in the reverse loading cycle. 

Asymmetrical deformation response is an important signature of the HBU SNF due to pellet-topellet cladding interaction. The curvature at the cladding tensile stress site has the highest 

deformation during fatigue cycles, so it was used to develop the maximum curvature plots 

shown in Figure 31 to serve as another important index for the HBU SNF fatigue life. 

Observations based on curvature amplitude (κa, half of curvature range ∆κ /2) are similar to 

those of strain amplitude since conversion simply involved a multiplication factor as shown in 

Eq. (11).

• The fatigue life decreases with increasing curvature amplitude with a defined κ-N 

curve. The curve fitting generated a power function y = 7.033·x−0.297 with R2 =0.847, 

where x is the number of cycles to failure, and y is the curvature amplitude (m-1). 

• An endurance limit may be located near 0.05–0.10 m−1 if it is defined at 107

 cycles. 

Dynamic tests were conducted under load control, and the curvature response of the rods was 

not necessarily symmetric, as shown above in Figure 26(d).

An investigation of the maxima of absolute values of curvature extremes (|κ|max, defined in 

Section 2) was pursued as an alternative quantity for describing the loading imposed on fuel. 

This is relevant because the maxima actually reflect the maximum tensile loading levels to 

which the specimen is subjected during the dynamic testing. The results are shown in Figure 31. 
52

(a)

(b)

Figure 31 (a) Maxima of absolute curvature extremes and curvature amplitudes as a 

function of number of cycles to failure, and (b) maxima of absolute strain 

extremes and strain amplitudes as a function of number of cycles to failure; 

markers with arrows indicate that the tests were stopped without failure

The κ-N curve in terms of maxima of cladding curvature |κ|max has been demonstrated to be 

similar to that of the curvature amplitude κa; the exponents of the curve-fit power function are 

similar. The scatter plots based on the two approaches are similar, with some changes in the 

relative positions of points between 1.1 and 3.9 × 105 cycles. Overall, the maxima of absolute 

curvature are larger than the curvature amplitude, depending on the magnitude of κm. 
53

The following observations can be made with regard to absolute curvature maxima:

• For as-irradiated specimens that failed during the tests, the fatigue life decreases

with increasing curvature maxima according to a well-defined κ-N curve. The data

points can be fit by using the power function y = 6.864·x−0.283 with a R2 value as high

as 0.893, where x is the number of cycles to failure, and y is the curvature maxima

(m-1).

• A fatigue limit may be located near 0.07 to 0.12 m−1 if it is defined at 107

 cycles.

• Due to the unsymmetrical deformation nature of HBU SNF under dynamic cycling,

the bending neutral axis of the deformed rod is expected to shift away from the

cladding tensile stress site. A first order approximation of simplified strain conversion

from the curvature was used to derive the strain maxima.

The associated strain maxima and strain amplitude vs. the number of cycles to failure are given

in Figure 31(b). Overall, the maxima of absolute strain are larger than the strain amplitude.

Additional curve fitting to the fatigue lifetime as a function of curvature amplitudes, curvature 

maxima, and strain maxima was completed using piece-wise power fits; the graphical results 

are presented in Figure 32, and the curve-fitting parameters are given in Table 8. Note that R2

values associated with the high-cycle power fit are generally quite low. This is likely because 

many of the tests at low amplitudes were discontinued without reaching sample fracture. 
54

Figure 32 (a) Curvature amplitudes, (b) maxima of absolute curvature extremes, and 

(c) maxima of absolute strain extremes as a function of number of cycles to 

failure; Curve-fitting was extended to include the no-failure data points; 

markers with arrows indicate that the tests were stopped without failure. 

Table 8 Curve-fitting parameters of curvature and strain as a function of cycles*

A a R2 A a R2

N <= 1E+06 N > 1E+06

Curvature amplitude 9.434 -0.325 0.850 0.131 -0.027 0.004

Strain amplitude 5.130 -0.325 0.851 0.069 -0.026 0.004

N <= 8E+05 N > 8E+05

Curvature maxima 11.257 -0.330 0.906 0.522 -0.096 0.136

Strain maxima 6.094 -0.330 0.909 0.283 -0.097 0.138

*Based on y = A·xa, where x represents cycles and y represents amplitude or maxima.

(a) (b)

(c)
55

4.3 Post-Irradiation Examination

The objective of the post-irradiation examination (PIE) was to gather meaningful information to 

identify underlying failure mechanisms. The primary focuses were: 

• to generate information on cladding features such as hydrides and incipient cracks

near pellet-to-pellet interfaces and near the mid-pellet,

• to reveal information about the fuel/cladding bond, and

• in order to preserve SNF system damage profile, the PIE was performed with fuel

pellet intact.

4.3.1 As-Irradiated Samples PIE

PIE was conducted on five as-irradiated fuel segments taken from the same father rod 

(E02-605) but with different mechanical loading histories:

• Untested – 605D1D

• Tested and failed under static loading condition at a pellet-to-pellet interface: S2/

605D1E

• Tested but non-failed under dynamic loading condition: D4/ 605D1C

• Tested and failed under dynamic loading at a pellet-to-pellet interface: D3/ 605C10A

• Tested and failed under dynamic loading but not at pellet-to-pellet interface: D5/

605D1B

PIEs were generally conducted on axial sections, but one PIE was carried out on a transverse 

section. For the tested and failed specimen, the preparation of axial sections involved the 

following steps: 

(1) Cut a one-inch segment on the fracture end. 

(2) Mount the segment with the bending plane in the horizontal direction, and cut the top 

half away. 

(3) Back-pot the remaining half to keep all the fuel in place. 

(4) Grind and polish the section to the required finish for metallography. 

Examinations were then performed on an optical microscope.21 The procedure for preparing 

sections for all the segments/specimens was similar to the procedure outlined above. 

PIE on the transverse section was conducted only on D3. The section was prepared using a 

piece of fuel rod approximately a 0.25-inch long cut away from the fracture surface of the 

specimen. The segment was defueled using hot nitric acid. The defueled cladding was removed 

from the hot cell and cleaned with water and alcohol in an ultrasonic cleaner. The inside and 

outside surfaces of the cladding were then coated with a thin layer of epoxy. Then the section 

was ground and polished to the required finish for examination using scanning electron 

microscopy (SEM). Observations on the transverse section can be found in Appendix F, and 

those for the axial sections will be briefly discussed below. 


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57

Figure 34 (a) Axial cross section of S2/ 605D1E, (b) fracture surface near back side of rod 

and an enlarged area, and (c) fracture surface near front side of rod and an 

enlarged area (A = delamination of circumferential hydride layers, B = spalling 

and cracking of oxide layer) 

Information on the PIEs of the other fuel segments can be found in Appendix F, including data 

on five axial sections and one transverse section.

The following observations were made based on the PIE. 

A

B

(a)

A

(b)

(c)
58

• Pellet-to-pellet interface and pellet-cladding interfaces (PCIs) near the edges or

peripheral areas of pellets appeared to be in good contact. No significant gap or

filling was observed in the interfaces.

• The pellets exhibited fracturing that was normal for as-irradiated fuel. These fractures

were observed in both untested and tested fuel segments.

• Tested specimen failures occurred primarily at pellet-to-pellet interface.

• The density of the hydride layers increases significantly toward the outside of

cladding wall. The cladding outer wall was covered with an oxide layer that was

75 μm thick.

4.3.2 HRT Samples PIE

The segments subjected to HRT were subjected to CIRFT testing at room temperature to 

evaluate the materials’ performance after HRT. After the tests were completed in the CIRFT, the

specimens were examined to verify that radial hydrides had been generated. Figure 35 shows 

the hydride morphology of HR1 sample following HRT. 

Figure 35 (a) High magnification micrograph showing radial hydrides of Sample HR-1 

(H ≈ 360–400 ppm); the HBU HBR specimen was pressurized to 140 MPa at 

400 °C with five thermal cycles; (b) high magnification micrograph showing 

radial hydrides of a sample of unirradiated, hyrodgen charged zircaloy-4 (H ≈

286 ± 20 ppm); the specimen was sectioned at the mid-plane of a sample that 

was 6 inches long

The hydrogen concentration of this sample was about 360–400 wppm. The specimen used for 

metallography was sectioned at the midplane of a sample 6 inches long. Although the in-cell 

HRT test conditions were the same as the out-of-cell test HR-HBR#2 (detailed in Appendix H),

and the hydrogen concentration of these two samples were compatible, the morphology of HBU 

sample HR1 was different from hydrided sample HR-HBR#2. For the unirradiated sample, the 

circumferential hydrides were uniformly distributed across the wall, and the maximum length of 

its radial hydrides was about 50–60 μm. For HBU samples, the radial hydrides were mainly 

distributed near the inner surface, and the circumferential hydrides were mostly seen near the 

outer surface. The length of the radial hydride of the HBU sample shown in Figure 35(a) was 

also much longer than the unirradiated sample shown in Figure 35(b). 

(a) (b)
59

4.4 Cladding Oxide Effect Study

CIRFT test data and the associated oxide thickness are shown in Table 9. 

Table 9 HBR SNF CIRFT data with the associated tested rod’s oxide thickness

Spec 

label ID

Moment 

amplitude

Number 

of cycles

Flexural 

rigidity κa sa εa

Oxide 

thickness1

Unit N∙m # N-m2 m-1 MPa % μm

As-irradiated CIRFT data

D8 7.74 >1.28E+07 72.3 0.107 66.31 0.057 70–100

D14 7.84 2.74E+05 70.0 0.112 66.65 0.06 100–110

D5 8.20 2.30E+06 71.9 0.114 69.71 0.061 40–70

D3 9.17 1.00E+06 73.4 0.125 77.94 0.067 70–100

D12 9.99 3.86E+05 69.4 0.144 83.62 0.078 70–100

D10 11.55 1.80E+05 66.4 0.174 98.19 0.094 70–100

D6 11.62 2.50E+05 56.7 0.205 99.55 0.11 70–100

D13 12.55 1.29E+05 83.1 0.151 106.68 0.081 100–110

D1 14.11 1.10E+05 65.6 0.215 117.26 0.117 70–100

D7 18.92 6.50E+04 53.9 0.351 160.84 0.189 70–100

D0 24.07 2.50E+04 54.8 0.439 206.11 0.236 40–70

D11 29.02 5.50E+03 61.9 0.469 241.22 0.254 70–100

D9 33.67 7.10E+03 58.5 0.576 288.31 0.31 70–100

HRT CIRFT data

HR1 15.15 4.19E+04 35.7 0.424 128.79 0.228 40-44

HR3 8.98 2.44E+05 41.0 0.219 76.34 0.118 40-44

HR4 14.76 5.47E+04 45.7 0.323 125.45 0.174 40-44

Under similar loading amplitudes, the CIRFT data of as-irradiated specimens suggest reduced 

fatigue life for higher oxide thicknesses For instance, the D8 sample with 70–100 μm oxide 

thickness and the D14 sample with 100–110 μm oxide thickness had similar loading amplitudes; 

the D8 sample had much longer fatigue life (>12M cycles) compared to the D14 sample (2.74E5 

cycles). The CIRFT data are shown with distinct labels for four oxide thickness ranges in 

Figure 36. 

The plots of as-irradiated baseline data for moment amplitude and strain amplitude input from 

HBR baseline data show a self-consistent trend: as-irradiated CIRFT data with higher oxide 

thickness showed less fatigue life, while for HRT data, a different trend was observed from 

moment amplitude and strain amplitude input plots. In the moment amplitude plot, HRT data 

showed reduced fatigue life below the mean trend curve of 70–100 μm oxide thickness, while as 

in strain amplitude, plot HRT data showed increased fatigue life above the mean trend curve. 

The HRT samples had the lowest oxide thickness, as shown in Figure 36. Thus, HRT data 

would have been expected to trend above the mean fatigue strength trend curve of 70–100 μm 

oxide thickness per the as-irradiated results, which suggests that reorientation may dominate 

any oxide effects. 
60

Figure 36 (a) Moment amplitude vs. number of failure cycles, (b) strain amplitude vs. 

number of failure cycles, as function of clad oxide thickness; HR4 only has 

thermal annealing treatment without pressurization applied

4.5 HRT CIRFT Test Data Evaluation

The detailed HRT procedure used for developing HRT CIRFT test samples is provided in 

Appendix H. Reoriented hydrides can change the SNF rod system response. In addition, the 

procedure used to reorient the hydrides may result in other physical changes to the specimen, 

including changing pellet-cladding bonding efficiency or developing gaps at pellet-clad interface 

regions due to the thermal treatment (400 °C) and clad tubing pressurization (140 MPa hoop 

HR4

y = 3.7093x-0.292

R2 = 0.9291

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

1.00E+03 1.00E+04 1.00E+05 1.00E+06 1.00E+07

Strain Amplitude (%)

Nf (Cycles)

40-70μm

70-100μm

100-110μm

40-44μm (HR)

Power (70-100μm)

HR4

y = 301.7x-0.26

R2 = 0.961

0

5

10

15

20

25

30

35

40

1.00E+03 1.00E+04 1.00E+05 1.00E+06 1.00E+07

Moment Amplitude (N-m)

Nf (cycles)

40-70μm

70-100μm

100-110μm

40-44μm (HR)

Power (70-100μm)

(a)

(b)


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63

5 CONCLUSIONS

High burnup HBR fuel rods were tested and evaluated under both static and dynamic loading 

conditions using the CIRFT equipment developed at ORNL. The CIRFT system is composed of 

a U-frame equipped with load cells for imposing pure bending loads on the spent fuel rod test 

specimen and measuring the in-situ curvature of the fuel rod during bending using a 3-LVDT

set-up. In general, the static CIRFT test results indicate a significant increase in flexural rigidity 

compared to that of the defueled HBU rod specimen, even considering the numerous stress 

riser sites into an HBU rod system.

The static tests were conducted under displacement control at a rate of 0.1mm/s to a maximum 

displacement of 12.0 mm at each loading point. Dynamic tests were conducted under load 

control of ±5.08 to ±35.56 N∙m, 5 Hz. PIEs were performed on selected specimens. Both asirradiated fuel rods and HR-treated fuel rods were tested. 

• The HBU as-irradiated HBR rods survived static unidirectional bending to a 

maximum curvature of 2.2–2.5 m−1

, or a maximum moment of 85–87 N∙m (Table 3). 

The maximum equivalent strain was 1.2–1.3%. 

• The HBU as-irradiated HBR rods exhibited a multiple-stage constitutive response 

with two linear stages followed by a nonlinear stage. The flexural rigidity at the initial 

stage was 63–79 N∙m2

, corresponding to an elastic modulus of 101–125 GPa (Table 

4). 

• The deformation of the HBU HBR subjected to HRT showed a similar loading 

behavior, but with lower flexural rigidity compared to that of as-irradiated specimens

under similar curvature. 

• A large majority of the specimen failures of the HBU HBR rods for both as-irradiated 

or HRT samples under static unidirectional loading occurred or initiated near the 

pellet-to-pellet interface, as validated by the post-test or PIEs. 

• Comparing the HBU HBR static bending response to a theoretical calculated 

response of cladding alone, the flexural rigidity was approximately twice what would 

be obtained if cladding properties alone were used to predict behavior.

• The fatigue life of HBU as-irradiated HBR rods in the cyclic test depended on the 

level of loading. Under loading with moments of ±8.89 to ±35.56 N∙m and strain of

±0.03 to ±0.38% at 5 Hz (Table 6), the fatigue life ranged from 5.5 × 103 to 

2.3 × 106

 cycles. 

• The ε-N curve of the HBU as-irradiated HBR rods can be described by a power 

function of y = 3.839·x−0.298, where x is the number of cycles to failure, and y is the 

strain amplitude (%).

• Based on strain amplitude, a fatigue limit is likely located between 0.03–0.05% if it is 

defined at 107

 cycles. 
64

• Maxima of the curvature in dynamic tests ranged from ±0.07 to ±0.62 m−1 at 5 Hz

(Table 6). The κ-N curve of the HBU as-irradiated HBR rods can be described by a

power function of y = 6.864·x−0.283

, where x is the number of cycles to failure, and y is

the maxima of cladding tensile curvature |κ|max (m-1). A fatigue limit based on maxima

of curvature is likely located between 0.07–0.12 m−1 if it is defined at 107

 cycles.

• The failure of HBU HBR rods under cyclic reverse loading was primarily due to

pellet-cladding interaction and generally resided at the pellet-to-pellet interface, as

validated by the PIE.

• Review of the fatigue life as a function of the calculated gauge strain amplitude

indicates that the HRT rods showed extended life relative to the as-irradiated rods.

On the other hand, review of the fatigue life as a function of the applied moment

amplitude indicates that the HRT rods showed reduced life relative to the asirradiated rods. However, reduction is largely within the scatter of the as-irradiated

rods.
65

6 REFERENCES

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Burnup PWR Fuel from H. B. Robinson Plant for Use in the NRC Experimental Program at 

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2001.

2 B. Hanson, H. Alsaed, C. Stockman, D. Enos, R. Meyer, and K. Sorenson, “Gap Analysis 

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3 D. Tang, A. Rigato, R. Einziger, “Flaw Effects and Flaw Reorientation on Spent Fuel Rod 

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4 J.-A. J. Wang, H. Wang, Y. Yan, R. Howard, and B. Bevard. “High Burn-up Spent Fuel 

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Ridge, TN, January 2011.

5 J.-A. J. Wang, H. Wang, T. Tan. H. Jiang, T. Cox, and Y. Yan. “Progress Letter Report on 

U Frame Test Setup and Bending Fatigue Test for Vibration Integrity Study (Out-of-Cell 

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Laboratory, Oak Ridge, TN, January 2012.

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Materials Fatigue Aging Performance,” ORNL Invention Disclosure 201102593, DOE S 

124, 149, April 8, 2011. Patent in review, 13/396,413, February 14, 2012. 

8 H. Wang, J.-A. J. Wang, T. Tan, H. Jiang, T. S. Cox, R. L. Howard, B. B. Bevard, and M. 

E. Flanagan, “Development of U-Frame Bending System for Studying the Vibration 

Integrity of Spent Nuclear Fuel,” Journal of Nuclear Materials 440, 201–213, 2013.

9 J.-A. J. Wang, B. B. Bevard, R. L. Howard, and M. E. Flanagan, “SNF Test System for 

Bending Stiffness and Vibration Integrity,” International High-Level Radioactive Waste 

Management Conference, Albuquerque, NM, April 28–May 2, 2013.

10 J.-A. J. Wang, H. Wang, T. Cox, and C. Baldwin, “Progress Letter Report on Bending 

Fatigue Test System Development for Spent Nuclear Fuel Vibration Integrity Study (Outof-Cell Fatigue Testing Development, Task 2.4),” ORNL/TM-2013/225, Oak Ridge 

National Laboratory, Oak Ridge, TN, July 2013.
66

Packaging and Transportation of Radioactive Materials, PATRAM 2013, San Francisco, 

CA, August 18, 2013. 

12 J.-A. J. Wang and H. Wang, “Progress Letter Report on Reversible Bending Fatigue 

Testing of Zry-4 Surrogate Rod (Out-of-Cell Fatigue Testing Development—Task 2.4),” 

ORNL/TM-2013/297. Oak Ridge National Laboratory, Oak Ridge, TN, August 2013.

13 Jiang Hao, Jy-An Wang, and Hong Wang, “The Impact of Interface Bonding Efficiency on 

High-Burnup Spent Nuclear Fuel Vibration Integrity,” Journal of Nuclear Engineering and 

Design, 309, 40–52, 2016.

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Review Meeting, Oak Ridge National Laboratory, August 8, 2011.

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Fuel Integrity Study in Transportation Environments, ORNL/TM-2014/63, April 2014. 

16 H. Wang and J.-A. J. Wang, “Bending testing and characterization of surrogate nuclear 

fuel rods made of Zircaloy-4 cladding and aluminum oxide pellets,” Journal of Nuclear 

Materials, 479, 470–482, 2016.

17 C. Baldwin, Oak Ridge National Laboratory, Mar. 6, 2013.

18 M. C. Billone, T. Burtseva, and Y. Y. Liu, “Baseline Properties and DBTT of High-Burnup 

PWR Cladding Alloys,” Proceedings of the 17th International Symposium on the Packing 

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20 M.C. Billonea, T.A. Burtsevaa, R.E. Einziger, “Ductile-to-brittle transition temperature for 

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000604, September 2014. 

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Surrogate Rod Testing to Analyze High Burnup Used Nuclear Fuel Mechanical 

Properties,” ORNL/TM-2014/257, DOE FCRD-UFD-2014-000603, August 2014.

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Northwest National Laboratory, Richland, Washington, July 2008.

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Bending Fatigue Test System for Investigating Vibration Integrity of Spent Nuclear Fuel 

During Transportation,” Proceedings of the 17th International Symposium on the 
APPENDIX A ROD SEGMENTS AND CUTTING PLAN 

Table A.1 Specification of rod segments to be used in hot cell testing

Spec Label ORNL Label

NRC 

Original Label

ORNL

Inventory Label (“Seg. ID”) Span Burnup* (rod average, GWd/MTU) Estimated Oxide* (μm) Estimated Hydrogen of Span**(wppm) Estimated Local Hydrogen (wppm)

S1 S1 S3 606C3C

4 66.8 70

–100 550

–750 650

S2 S2 S2 605D1E

2

–

3 66.5 40

–70 360

–550 400

S3 DCAL DCAL 609C5

4 66.5 70

–100 550

–750 550

S4 SCAL SCAL 609C6

4 66.5 70

–100 550

–750 550

D1 DL1 DL1 607C4B

4 63.8 70

–100 550

–750 700

D2 DL2 DL2 608C4B

4 63.8 70

–100 550

–750 700

D3 DL3 DL3 605C10A

4 66.5 70

–100 550

–750 550

D4 DM1 DM1 605D1C

2

–

3 66.5 40

–70 360

–550 500

D5 DM2 DM2 605D1B

2

–

3 66.5 40

–70 360

–550 550

D6 DH1 DH1 609C4

4 66.5 70

–100 550

–750 700

D7 DH2 DH2 609C3

4 66.5 70

–100 550

–750 750

D8 DM3 DM3 606C3E

4 66.8 70

–100 550

–750 550

D9 DH3 DH3 609C7

4 66.5 70

–100 550

–750 550

D10 S5 S5 606C3A

4 66.8 70

–100 550

–750 750

D11 R1 R1 607C4A

4 63.8 70

–100 550

–750 700

D12 R2 R2 608C4A

4 63.8 70

–100 550

–750 700

D13 R3 R3 606B3E

5 66.5 100

–110 750

–800 750

D14 R4 R4 606B3D

5 66.5 100

–110 750

–800 750

D15 R5 R5 606B3C

5 66.5 100

–110 750

–800 750

D16/ Dcal S3 S1 605D1F

2

–

3 66.5 40

–70 360

–550 360

Demo1 DEMO1 DEMO1 606B2

5 66.5 100

–110 750

–800 750

Scal1 DEMO2 DEMO2 606B3B

5 66.5 100

–110 750

–800 750

Scal2 S4 S4 606C3B

4 66.8 70

–100 550

–750 700

HR1 HR1 HR1 607D4C

2 63.8 40-44 360-400 400

HR2 HR2 HR2 607D4A

2 63.8 40-44 360-400 400

HR3 HR3 HR3 608D4A

2 63.8 40-44 360-400 400

HR4 HR4 HR4 608D4C

2 63.8 40-44 360-400 400

* E. J. Ruzauskas, and K. N. Fardell, “Design, Operation, and Performance Data for High Burnup PWR Fuel from H. B. Robinson Plant for Use in the NRC Experimental Program at

Argonne National Laboratory,” Electric Power Research Institute, EPRI Report 1001558, 2001.** M.C. Billone, T. Burtseva, and Y.Y. Liu, “Baseline Properties and DBTT of High-Burnup PWR Cladding Alloys,” Proceedings of the 17th International Symposium on the Packaging 

and Transportation of Radioactive Materials, PATRAM 2013, San Francisco, CA, August 18–23, 2013; C. Baldwin, Oak Ridge National Laboratory, Mar. 6, 2013.

A-1

APPENDIX A ROD SEGMENTS AND CUTTING PLAN 
(a)

S2 S1 DL3 DM1 DM2

Used for 

drill 

practice

A2
(b)

Demo 

1 

Demo 

2 R5 R4 R3 S5 S4 S3 DM3

A3
(c)

DL1 R1 HR1 HR2

A4
(d)

DL2 R2 HR3 HR4

A5
(e)

Figure A.1 Cutting plan for (a) E02-605, (b) E14-606, (c) F07-607, (d) G10-608, and (e) R05-609

DH1 DH2 DH3 Scal Dcal

A6
APPENDIX B METROLOGY OF ROD SEGMENTS 

Table B.1 Diameter measurements on fuel rod segment R5 (606B3C) (unit: inch)*

Distance 

from End 

(Inches)

0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 5.5

A - (000) 0.42450 0.42436 0.42450 0.42440 0.42402 0.42390 0.42388 0.42400 0.42430 0.42392 0.42368B - (120) 0.42428 0.42416 0.42400 0.42416 0.42392 0.42396 0.42370 0.42404 0.42402 0.42374 0.42382C - (240) 0.42432 0.42430 0.42386 0.42412 0.42374 0.42390 0.42446 0.42506 0.42474 0.42384 0.42382Maximum 0.42450 0.42436 0.42450 0.42440 0.42402 0.42396 0.42446 0.42506 0.42474 0.42392 0.42382Minimum 0.42428 0.42416 0.42386 0.42412 0.42374 0.42390 0.42370 0.42400 0.42402 0.42374 0.42368Mean 0.42437 0.42427 0.42412 0.42423 0.42389 0.42392 0.42401 0.42437 0.42435 0.42383 0.42377Std. Dev. 0.00012 0.00010 0.00034 0.00015 0.00014 0.00003 0.00040 0.00060 0.00036 0.00009 0.00008OverallMaximum 0.42506Minimum 0.42368Mean 0.42410Std. Dev. 0.00032

*Based on C. Baldwin, ORNL, Mar. 6, 2013.

B-1

APPENDIX B METROLOGY OF ROD SEGMENTS 
Table B.2 Diameter measurements on HBR fuel rod segments (unit: inch)*

Spec Label ORNL Label NRC Original Label Segment ID 0 Degree 90 Degree Average Std. Dev.

-0.5 0.0 +0.5 -0.5 0.0 +0.5

S1 S1 S3 606C3C 0.42324 0.42304 0.42306 0.42320 0.42306 0.42332 0.42315 0.00012S2 S2 S2 605D1E 0.42322 0.42298 0.42298 0.42310 0.42310 0.42310 0.42308 0.00009S3 DCAL DCAL 609C5 0.42222 0.42248 0.42272 0.42298 0.42288 0.42288 0.42269 0.00029S4 SCAL SCAL 609C6 0.42320 0.42308 0.42288 0.42228 0.42252 0.42298 0.42282 0.00035

D1 DL1 DL1 607C4B 0.42634 0.42650 0.42664 0.42690 0.42640 0.42640 0.42653 0.00021D2 DL2 DL2 608C4B 0.42442 0.42422 0.42410 0.42482 0.42462 0.42454 0.42445 0.00026D3 DL3 DL3 605C10A 0.42398 0.42402 0.42422 0.42444 0.42418 0.42420 0.42417 0.00016D4 DM1 DM1 605D1C 0.42314 0.42336 0.42352 0.42302 0.42310 0.42316 0.42322 0.00019D5 DM2 DM2 605D1B 0.42376 0.42412 0.42396 0.42426 0.42424 0.42410 0.42407 0.00019D6 DH1 DH1 609C4 0.42338 0.42336 0.42286 0.42288 0.42310 0.42322 0.42313 0.00023D7 DH2 DH2 609C3 0.42464 0.42404 0.42394 0.42430 0.42446 0.42426 0.42427 0.00026D8 DM3 DM3 606C3E 0.42250 0.42280 0.42268 0.42284 0.42292 0.42292 0.42278 0.00016D9 DH3 DH3 609C7 0.42324 0.42296 0.42304 0.42236 0.42256 0.42230 0.42274 0.00039

D10 S5 S5 606C3A 0.42368 0.42378 0.42374 0.42344 0.42326 0.42348 0.42356 0.00020D11 R1 R1 607C4A 0.42642 0.42690 0.42676 0.42614 0.42616 0.42644 0.42647 0.00031D12 R2 R2 608C4A 0.42630 0.42588 0.42618 0.42622 0.42574 0.42592 0.42604 0.00022D13 R3 R3 606B3E 0.42402 0.42396 0.42408 0.42334 0.42356 0.42400 0.42383 0.00030D14 R4 R4 606B3D 0.42416 0.42404 0.42378 0.42368 0.42368 0.42358 0.42382 0.00023D15 R5 R5 606B3C 0.42430 0.42390 0.42388 0.42392 0.42396 0.42370 0.42394 0.00020

D16/ Dcal S3 S1 605D1F 0.42326 0.42306 0.42298 0.42324 0.42294 0.42306 0.42309 0.00013

Demo1 DEMO 1 DEMO 1 606B2 0.42342 0.42356 0.42364 0.42346 0.42364 0.42372 0.42357 0.00012

Scal1 DEMO 2 DEMO 2 606B3B 0.42406 0.42388 0.42424 0.42402 0.42394 0.42406 0.42403 0.00012Scal2 S4 S4 606C3B 0.42376 0.42354 0.42340 0.42390 0.42376 0.42374 0.42368 0.00018

HR1 HR1 HR1 607D4C 0.42360HR2 HR2 HR2 607D4A 0.42200HR3 HR3 HR3 608D4A 0.42200HR4 HR4 HR4 608D4C

Average 0.4237

Std. Dev. 0.0012

*Based on C. Baldwin, ORNL, Mar. 6, 201.

B2
APPENDIX C BENCHMARKING OF TEST SYSTEM

C-1

C.1 TUNING AND STATIC TEST ON DEMO1/ 606B2 (+ DYNAMIC)

Tuning was completed September 23, 2013 on test specimen Demo1 (606B2) by using TuneIQ 

on both motors (Table C.1). 

Table C.1 Tuning parameters obtained for test specimen Demo1

Axial 1 Axial 2

TuneIQ1 1.5860 TuneIQ1 1.8869

TuneIQ2 0.0192 TuneIQ2 0.0193

TuneIQ3 -9.6663 TuneIQ3 -11.0162

Demo1 (606B2, 66.5 GWd/MTU burnup, 100–110 μm oxide layer, 750 ppm H content) was 

initially tested to the relative displacement of 2 mm at a rate of 0.1 mm/s at each loading point 

on the U-frame. Such a displacement level was low to serve as a pilot test to detect the 

response of the rod. The rod was then tested for 11 cycles with an increased relative 

displacement of 20 mm, corresponding to a moment of 77.7 Nm (load correction applied 

regarding the resistive force from testing system; the same applies in the following unless 

specified otherwise), and the rod appeared to be stable. Four additional cycles were carried out 

with further increased displacement of 24 mm (moment 85 N∙m), and the rod was found to still 

work fine. The responses of the rod during the loading processes specified above are presented

in Figure C.1. 

A follow-up cycle test was conducted by using ±300 N or 30.48 N∙m 5 Hz. The rod eventually 

fractured in the gauge section near motor 1 around 3.96 × 103 cycles. The monitored curvature, 

moment, and deduced flexural rigidity as a function of number of cycles are provided in Figure 

C.2. The fracture surfaces (Figure C.3) of the rod indicate that the failure likely involved the 

fracture of a fuel pellet near the interface as can be seen from the attached fragments.

It has been observed that at a low displacement level, the rod exhibited a linear response. A 

flexural rigidity of 40 N∙m2 was obtained from the slope of the curvature-moment curve. A 

nonlinear response was exhibited under increased displacement. Three deformation stages can 

be delineated with two turning points near 20 and 55 N∙m, at which point the slope of the curve 

changed. Below the first turning point, the response appeared repeatable under loading. Beyond 

the second turning point, the response of the rod became nonlinear. Unloading and loading 

curves were mostly overlaid if the loading level was less than the previous maximum level; 

otherwise, the loading curve followed the response of the rod in a monotonic loading process.

The following be seen based on testing of the first fuel rod: 

• The testing system worked stably for both static and dynamic loadings when the

tuning parameters generated above were used.

• A three-stage deformation process was evident when the deformation was

substantially large.

• No failure was seen from the use of Demo1 under static loading, even though the

displacement imposed almost reached the system capacity (25 mm). Therefore, a

cyclic loading process is required to observe the failure mode of the rod.

APPENDIX C BENCHMARKING OF TEST SYSTEM 
C-2

• The testing setup worked as designed, confining the final failure within the gauge

section of the rod.

(a) (b)

Figure C.1 (a) Curve of moment versus curvature with maximum relative displacement of 

20 mm, and (b) moment versus curvature in the subsequent repeats of static 

tests with relative displacement of 20 or 24 mm for Demo1 (606B2).

The rate at the U-frame loading point was set at 0.1 and 0.2 mm/s for loading and 

unloading, respectively.
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The specimen had been subjected to repeated static loading cycles before the dynamic cycle 

test.

C.2 STATIC TEST ON SCAL1/ DEMO2/ 606B3B

The test on Scal1/ Demo2 (606B3B, 66.5 GWd/MTU burnup, 100–110 μm oxide layer, 750 ppm 

H content) was conducted at 0.1 mm/s at each loading point of the U-frame. The rod fractured 

when the relative displacement reached 22.920 mm, corresponding to a fracture moment of 

78.2 N∙m. Again, two turning points can be defined clearly near 20 and 50 N∙m prior to the 

failure, as shown in Figure C.4. Rod failure occurred at the middle of the gauge section. The 

fracture surface of the cladding (Figure C.5) appeared to be very serrated. A large tooth-shaped 

fragment was attached on the compressive side of the rod, but it broke off during specimen 

handling. The final fracture of the rod was seemingly related to the fracture of a fuel pellet near 

a pellet to pellet interface. 

Observations:

• The moment at fracture for Scal1 is lower than the maximum moment experienced

by Demo1, as discussed above. Therefore, a substantial mechanical property

difference is seen between the two rods, although they have the same thickness of

oxide layer and the same level of H content.

• Lateral surfaces of the rod were covered with a large number of equally spaced

circumferential cracks throughout the gauge section on both the front (facing

operator) and the back sides of the rod. Whether the cracks were linked around the

whole periphery of the rod is uncertain.

• Extensive subsurface spalling can be seen over the back or tensile side of the rod.

Circumferential cracks and spalling also occurred on Demo1 and several other rods,

as discussed below.

Figure C.4 Moment-curvature curve for Scal1/ Demo2 (606B3B) with a rate of 0.1 mm/sec 

at the U-frame loading points/ the rod broke at peak moment = 83 N∙m 
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APPENDIX D STATIC TEST RESULTS 

D-1

APPENDIX D STATIC TEST RESULTS 

D.1 S1/ 606C3C (+ DYNAMIC)

S1 (606C3C, 66.8 GWd/MTU burnup, 70–100 μm oxide layer, 650 ppm H content) was tested 

under four repeated static loading cycles before being taken into a dynamic cyclic test. The 

moment-curvature curve and equivalent stress-strain curve for the initial cycle are presented in 

Figure D.1 (the curves of subsequent cycles overlaid to a great extent and are not plotted for 

clarity). Again, the turning points were revealed near 12 and 59 N∙m. The rod fractured in a 

follow-up cyclic test under ±25.40 to 30.48 N∙m 5 Hz near 1.4×104

 cycles. 

Fracture occurred near motor 1 (right side) with a fuel release of 0.9 grams. Fractography using 

an optical microscope as shown in Figure D.2 revealed that the rod failure was related to 

fractures of a pellet adjacent to a pellet-to-pellet interface. The exit of the pellet end-face can be 

identified from the haze zone characterized by the brownish blue color. At the same time, a 

serrated fracture surface of cladding was observed with a slender attachment (about one roddiameter length). 

The lateral surfaces shown in Figures D.2(c) and (d) near the rod’s neutral axis of bending were 

mostly free of circumferential cracks. On the other hand, one of the stressed lateral surfaces 

was covered with a large number of circumferential cracks and related subsurface spalling, as 

seen in Figure D.2(d). It is uncertain whether the spalling was related to tensile stress in static 

loading, as the specimen ID was 90o off to the stressed sides.

(a) (b)

Figure D.1 (a) Curve of moment versus curvature, and (b) equivalent strain–stress curve 

for S1 (606C3C), based on static loading cycle in which the maximum relative 

displacement was 24 mm and the rate at the U-frame loading points was set at 

0.1 and 0.2 mm/s in loading and unloading, respectively
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D-3 

Fracture occurred on a pellet-to-pellet interface, Figure D.4 clearly shows the smooth brownish 

end face of the adjacent pellet. The attachment shown in Figure D.4(a) is convex-shaped and 

covers almost 1⁄3 of the end face. Several small attachments appear around the peripheral area. 

The cladding fracture appeared very serrated with a sizable tooth-shaped attachment on the 

compressive side of the rod. The lateral stressed surface of the rod was covered by sparsely 

distributed circumferential spalling. The rod’s minor damaged lateral surface of bending agreed 

with the results of Dcal/ S3, as they were prepared from same segment (605D) and with the 

same level of preexisting oxides and hydrides.

The fuel release was about 0.6 grams.

(a) (b)

Figure D.3 (a) Curve of moment versus curvature and (b) equivalent strain-stress curve 

for S2 (605D1E), based on static loading cycle in which maximum relative 

displacement was 24 mm and the rate at the U-frame loading points was set at 

0.1 and 0.2 mm/s in loading and unloading, respectively
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D-5 

Rod failure was seemingly related to the fracture of a pellet (Figure D.6). No involvement of the 

pellet-to-pellet interface can be seen. The fracture surface of the pellet was characterized by 

attached fragments that were about a quarter the diameter of a typical pellet in size. 

A serrated fracture surface was seen again over the cladding. Furthermore, the fracture surface 

was tilted to a certain degree to the longitudinal axis of the rod. This is in contrast to rods in 

which the failure actively involves the pellet-to-pellet interface and the fracture surfaces are 

more or less normal to the longitudinal axis of the rod. Circumferential cracks and spalling can 

be observed on lateral stressed sides, especially on the tensile side in static loading. The spalls 

had an extended dimension in the circumferential direction, some of which were overlaid in the 

longitudinal axis of the rod. However, the extent of overlaying is not as large as seen in Demo1 

and Scal1/ Demo2. This can be attributed to less development of oxides and hydrides.

(a) (b)

Figure D.5 (a) Curve of moment versus curvature and (b) equivalent strain-stress curve 

for S3/ Dcal (609C5), based on a static loading cycle in which maximum 

relative displacement was 24 mm and the rate at the U-frame loading points 

was set at 0.1 and 0.2 mm/s in loading and unloading, respectively
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D-7 

Fracture occurred near motor 2 (left side) and was shown to be related to pellet fractures near a 

pellet-to-pellet interface as illustrated in Figure D.8. The exit of the end face of the involved 

pellet can be seen clearly from the bluish haze zone. A serrated fracture surface of cladding can 

also be seen. 

The specimen ID was found to be 90o off the stressed surfaces of the rod. A conspicuous sharp 

attachment on one side suggests where the intensified compression once resided, inducing the 

shearing fracture. The opposite side is covered by a certain amount of spalling, as discussed 

above. The less stressed lateral side near the rod’s neutral axis of bending appeared to be clear 

of circumferential cracks.

Failure of the rod involved a significant amount of fragmentation in the fractured pellet. The 

released fuel weighed 1.3 grams.

(a) (b)

Figure D.7 (a) Curve of moment versus curvature and (b) equivalent strain-stress curve 

for S4/ Scal (609C6), based on static loading cycle in which maximum relative 

displacement was 24 mm and the rate at the U-frame loading points was set at 

0.1 and 0.2 mm/s in loading and unloading, respectively
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D-9 

The responses of the specimen under static loading are represented in Fig. D.9 in terms of 

moment-curvature and stress-strain curve, and those under dynamic loading are presented in 

Fig.D.10. 

(a) (b)

Figure D.9 (a) Curve of moment versus curvature and (b) equivalent strain-stress curve 

for HR2 (607D4A); tests were conducted with static loading to 12 mm or +/- 

12 mm at each U-frame loading point
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APPENDIX E DYNAMIC TEST RESULTS

E-1

E.1 D1/ DL1/ 607C4B (±15.24 N∙m 5 Hz)

The test on D1/ DL1 (607C4B, 63.8 GWd/MTU burnup, 70–100 μm oxide layer, 700 ppm H 

content) was conducted under ±15.24 N∙m 5 Hz. A lifetime of 1.1 × 105 cycles was obtained with 

less than 1 gram of fuel particles collected. The cycle-to-failure of DL1 is apparently longer than 

that obtained for Dcal. The latter was subjected to a higher amplitude load of 25.40 N∙m. 

Measurements were conducted under displacement control at two relative displacement levels: 

0.8 and 1.2 mm. The curvature-moment loops obtained for N = 1 and 111,000 cycles are given 

in Figure E.1. The loops actually became quite slender with the accumulated cycles, whil the 

amplitude of the loops decreased to a certain degree. The ranges of moment in relation to 

curvature and flexural rigidity are illustrated in Figure E.2. It can be seen that most of the rigidity 

degradations occurred in the first 1,000 cycles. Variations of these quantities as a function of the 

number of cycles are shown in Figure E.3. While the curvature stayed at a consist level as a 

result of displacement control, the moment decreased. The rigidity of measurements at two 

displacements appeared to converge before rod failure while exhibiting a slightly declining trend. 

The curvature, moment, and flexural rigidity based on online monitoring data are presented in

Figure E.4. It can be seen that the 0.65 m-1 curvature obtained in the cyclic test is higher than 

those used in measurements as part of the experiment design. Accordingly, the flexural rigidity 

based on online monitoring was about 50 N∙m2

, which is a little lower than that acquired in 

measurements. The result agrees with the declining flexural rigidity with increasing curvature as 

shown in Figure E.3. Overall, the rod exhibited a relatively stable response prior to its final 

breakage. 

Failure was observed in the gauge section near motor 2 (left side). The failure occurred at the 

pellet-to-pellet interface as illustrated in Figure E.5. The end faces of two neighboring pellets 

were essentially clean, while the fracture surfaces of the cladding were serrated. Some 

striations had developed in the circumferential direction, but they were not well differentiated at 

this magnification. 

Both lateral stressed sides of the rod were covered with equally spaced circumferential cracks 

throughout the gauge section. Spalling only occurred at the local area near the fracture. A 

similar degree of damage on the lateral stressed sides exhibited by D1/ DL1 is different from 

that seen from unidirectional bending. The large deformation arising from unidirectional bending 

can produce a greater extent of damage or spalling on one side and less on the other side, as 

seen in S3/ Dcal (Figure D.6). The extensive circumferential cracks revealed in D1/ DL1 did not 

occur in Dcal/ S3, even though the load amplitude in the latter was higher. Therefore, the 

parallel equally spaced cracks were likely related to a preexisting condition, as D1/ DL1 actually 

had a thicker layer of oxidation and a higher H content. 

APPENDIX E DYNAMIC TEST RESULTS 
E-2

(a) (b)

Figure E.1 Moment-curvature loops based on measurements when (a) N = 1 and 

(b) N = 111,000 cycles for D1/ DL1 (607C4B); measurements were made with 

0.8 and 1.2 mm relative displacements; Nf = 1.1 105 cycles under ±15.24 N∙m 

5 Hz; fuel particles collected < 1.0 gram

(a) (b)

Figure E.2 (a) Moment-curvature relation and (b) moment-flexural rigidity relation at 

various numbers of cycles for D1/DL1 (607C4B); Nf = 1.1 × 105 cycles under 

±15.24 N∙m 5 Hz; fuel particles collected < 1.0 gram
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APPENDIX F POST-IRRADIATION EXAMINATION (PIE)

F-1

F.1 Preparation of specimens

Axial section

PIE on the axial section was conducted for one untested rod segment, on a fractured specimen 

from static test S2, and on three tested specimens from dynamic tests (one unfailed—DM1, and 

two fractured—DL3 and DM2).

For the tested and fractured specimen, the preparation of axial cross sections mainly involved 

the following steps: (1) cut a one-inch segment on the fracture end, (2) mount the segment with 

the bending plane in horizontal direction and cut the top half away, (3) back-pot the remaining 

half to keep all the fuel in place, and (4) grind and polish the section to the required finishing of 

metallography. Examinations were then carried out by using an optical microscope.18 

The section preparation for other segments/specimens was similar to the above procedure. The 

untested rod segment was taken from a section of the same father fuel rod with same estimated 

amount of hydrogen content as those in other PIE specimens. The section for a tested and 

unfailed specimen was taken from the gauge section of the specimen. 

Transverse section

PIE on transverse section was conducted only on D3. 

The section was prepared by using an approximately 0.25-inch length of cladding cut away from 

the remaining half of specimen D3 (specimen holder A). The segment was defueled by using 

hot nitric acid. The defueled cladding was removed from the hot cell and cleaned with water and 

alcohol in an ultrasonic cleaner. The inside and outside surfaces of the clad were then coated 

with a thin layer of epoxy. The section was then ground and polished to the required finish for 

examination using scanning electron microscopy (SEM).

F.2 Untested segment 605D1D

The untested section was prepared one-inch fuel segment from segment 605D1D. Adjacent 

segments indicated that the estimated hydrogen content is around 400–500 ppm. An optical 

image of the section is given in Figure F.1a. 

The three dished fuel pellets can be easily identified. Two primary fractures were developed 

roughly along and normal to the axial direction. A detail study (Figure F.1b) showed that the 

pellet-to-pellet interface near the edge of dish and the pellet-to-clad interface were fairly close, 

without any visible gap. Secondary fractures were developed within pellets near the peripheral 

area close to the pellet-to-clad interfaces. The cavity is also observed around some triple 

boundaries of pellets and clad. 

The parallel axial hydride layers can also be seen clearly over the whole thickness. The axial 

hydrides are widely spaced in the large middle part of wall. The density of hydrides becomes 

much higher toward the outside of clad and in the clad-pellet interface area. The axial hydride 

layers are sizable, and in the middle of clad wall, the layers can be over 1000 μm.

The outside surface of clad is covered with a continuous thin oxide layer that is measured at 

about 75μm.

APPENDIX F POST-IRRADIATION EXAMINATION (PIE) 
)2

D

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F-3

(a)

(b)

(c)

Figure F.2 (a) Axial cross section of S2/605D1E, (b) fracture surface near back side of rod 

and an enlarged area, and (c) fracture surface near front side of rod and an 

enlarged area; label A shows delamination, and label B illustrates cracking of 

the oxide layer

F.4 Tested and un-failed in dynamic test D4/DM1/605D1C

D4/DM1 was based on segment 605D1C, with estimated hydrogen and oxide layer thicknesses 

of 500 ppm and 40–70 μm. D4 survived 1.1 × 107 cycles of dynamic test under ±7.62 N·m, 5Hz) 

without failure. 

A

A

B
F-4

The axial section of a one-inch segment is shown in Figure F.3. The primary fractures in the 

same range of pellet size can be seen clearly. The fracture size and fracture pattern were 

similar to those observed for untested fuel segments. The figure also illustrates closely spaced 

fractures that interacted and resulted in slender fragments aligned in the axial direction. One 

fragment was pulled out in the grinding process.

The pellet-to-clad interface appeared tight. The axial hydrides are similar to those observed in 

the untested fuel segment, as they were neighboring segments in the same father fuel rod. In 

addition, the surface oxide layer was intact, even though the specimen experienced a high-cycle 

reversible bending fatigue test.

(a)

(b)

(c)

Figure F.3 (a) Axial cross section of DM1/605D1C, (b) back side of rod, and (c) front side 

of rod
F-5

F.5 Fractured in dynamic test D3/ DL3/ 605C10A

Axial section

D3/DL3 was based on segment 605C10A with and estimate hydrogen contact of 550 ppm and 

an oxide layer of 70–100μm. As mentioned earlier, D3 was tested under ±10.16 N·m 5Hz with a 

lifetime 106

 cycles. 

The optical image of axial cross section is presented in Figure F.4 . The facture surface was 

located at the left side at a pellet-to-pellet interface. It can be seen that the development of 

irradiation-induced primary fractures was similar to that of the specimens examined above. 

The part of clad fracture surface near the front side was not preserved. The fractured surface 

near back side of specimen is highly serrated. The cavity can be seen near the fracture that 

may have resulted from the hydride cracking since they are coincident with the axial hydride 

layers. Spalling of outside oxide layer was again seen near the fracture surface.
F-6

(a)

(b)

(c)

Figure F.4 (a) Axial cross section of D3/605C10A, (b) fracture surface near front side of 

rod, and (c) fracture surface near back side of rod and an enlarged area; label

A indicates spalling of the oxide layer, and label B shows the cavity aligning 

with the axial hydride layer

Transverse section

A set of SEM images is presented in Figures F.5 (a–d), which show the different areas of the 

transverse cross section. The images corresponding to the 12 and 6 o’clock positions represent 

A

B
F-7

the maximum alternating stress areas of the rods subjected to potential crack initiation. The 

cladding was apparently flattened at the maximum stress areas after the fracture, and the failure 

origin was not preserved. However, the areas near the 3 and 9 o’clock positions of the fractured 

surface revealed a clear laminate structure, as seen in Figure F.5(e). This may be related to the 

preexisting circumferential hydride structure in the HBU Zry-4 cladding.

(a) (b)

(c) (d)

(e)

Figure F.5 SEM images of clad cross section based on the defueled D3: 

(a) 12 o’clock (reference mark), (b) 3 o’clock, (c) 6 o’clock, (d) 9 o’clock, and 

(e) enlarged area (labeled A) 

A
F-8

F.6 Fractured in dynamic test D5/ DM2/ 605D1B

D5/DM2 was based on segment 605D1B, with and estimated hydrogen contact of 550 ppm and 

an oxide layer of 40–70μm. As mentioned above, D5 was tested under ±9.14 N·m 5Hz with a 

lifetime of 2.3 × 106

 cycles. 

The optical image of the axial cross section is shown in Figure F.6 . The facture surface was 

located at the left side within a pellet. The primary fractures, which were the same size scale as 

a pellet, were similar to those revealed in the specimens examined above. The primary fractures 

were preexisting and may result from the radiation effect because they were similar in size and 

pattern to those observed in untested fuel segments. One primary fracture system normal to the 

axial direction of fuel corresponded to the fracture surface of fuel and thus served as a main 

contributor to the failure.

The part of the clad fractured surface near the front side was flattened due to the impacts of 

fracture surfaces before the machine reached the predetermined termination condition. The part 

of fractured surface near the back side of specimen was preserved and exhibited a highly

serrated fractured surface. The delamination and connected cavities can be seen near the 

fracture, which may be attributed to the hydride cracking since they are coincident with the axial 

hydride layers. Spalling of outside the oxide layer was seen near the fractured surface.
F-9

(a)

(b)

(c)

Figure F.6 (a) Axial cross section of D5/605D1B, (b) fractured surface near the front side 

of the rod, and (c) fractured surface near the back side of the rod and an 

enlarged area; label A indicates the spalling of the oxide layer; label B 

indicates the delamination, and label C shows the connected cavities aligning 

with the axial hydride layers

A

C

B

APPENDIX G VERIFICATION OF CURVATURE MEASUREMENT

G-1

G.1 TECHNICAL APPROACH

Several methods were used to verify and justify the three-LVDT curvature measurement. The 

work was conducted in an out-of-cell study by using surrogate fuel rods; the observations and 

suggestions are presented in this appendix. While the curvature measurement based on diskhead LVDTs is used as a baseline, the following methods are examined:

1) Use of the chisel head in the curvature measurement. It is believed that the main

contribution of curvature measurement deviating from the ideal measurement

condition is caused by the disk head. The contact point may be shifted during the

measurement, which may change the probe or sensor spacing (h) used in the

equation of curvature. Use of the chisel head is expected to reduce such effect.

2) Use of sensor spacing adjustment (∆h) in the curvature measurement with

disk-head. The adjustment or correction was based on the assumption of the equal

gauge length in both the peak and the valley of one data block (two cycles of input).

The half gauge length Lg/2 was calculated according to the equation as follows:

Lg = d2 /22/ κ

where d2 is the middle LVDT reading in m, and κ is the calculated curvature in m-1. 

3) Use of strain based on curvature measurement. The outer fiber strain of the

bending rod can be obtained according to the measured curvature (either disk head

or chisel head):

max = κε y ,

where ymax is the maximum distance to the neutral axis of the test rod and is 

measured by the radius of the rod.

4) Use of strain gauge. The gauge-based strain is obtained according the following

equation,

10 /( .0 25 ) 6 = VBR ASg ε ,

where ε is micro strain or με, VBR is the bridge voltage, A is the amplification, and Sg

is the gauge factor.

APPENDIX G VERIFICATION OF CURVATURE MEASUREMENT 
G-2

G.2 EXPERIMENTAL TECHNIQUE

Two surrogate fuel rods were used. 

1) The first rod was made of SS304 tube-only and tested under small amplitude input.

The tests were conducted using sine waveforms under displacement control at

0.05Hz and load control at 5Hz. Tested items include disk-head LVDT, chisel-head

LVDT, and strain gauge.

2) The second was made of polycarbonate rod and was tested in a large amplitude

input under displacement. A ramping waveform and a 0.05 Hz sine wave were

tested. Disk head, chisel -head, and strain gauge were investigated.

The tests using ramping wave followed the procedure established in the static test, and those 

using 0.05 and 5Hz sine waves followed the dynamic test procedure.

A P3 strain indicator and recorder and a strain signal conditioning amplifier 2310B were used. 

Two types of strain gauges (Micro-Measurements, Raleigh, NC) were identified for this study, 

and the gauge factors are given in Table G.1. The setups of surrogate rods with the strain 

gauge mounted are shown in Figures G.1 and G.2. 

Table G.1 Strain gauges used in out-cell verification tests

Surrogate rod Strain gauge designation Gauge factor

SS304 surrogate rod EA-06-250BF-350 2.055 +/- 0.5%

Polycarbonate rod CEA-06-125UW-350 2.095+/-0.5%

5) Calculation of strain based on given flexural rigidity. The curvature can be estimated

by

ε = (M / EI) ymax .

where M is the moment amplitude input and EI is given rigidity of the rod. 

EI= 42.8 Nm2

, ymax = 4.78 mm for a SS304 rod.
G-3

Figure G.2 Setup of polycarbonate surrogate rod with strain gauge mounted in CIRFT 

testing

Figure G.1 Setup of SS304 surrogate rod with strain gauge mounted in CIRFT testing
G-4

G.3 RESULTS ON SS304 SURROGATE ROD

The proposed curvature correction was tested against several established methods, including 

curvature measurement, using a chisel type probe and a strain gauge. 

G.3.1 Chisel type probe as calibration method

Testing on surrogate rod SS30402 was conducted at 0.05 Hz and 5 Hz to simulate the static 

and dynamic loading (Table G.2). 

Table G.2 Testing condition of surrogate rod SS30402 with disk and chisel types

No. pk1 (mm, N) pk2 (mm, N) Frequency Cycles Sum LVDT type

44 -0.4 0.4 0.05 3 Flat disk, D6

45 -0.6 0.6 0.05 3 Flat disk, D6

46 -0.8 0.8 0.05 3 Flat disk, D6

47 -75 75 5 500 500 Flat disk, D6

48 -0.4 0.4 0.05 3 Flat disk, D6

49 -0.6 0.6 0.05 3 Flat disk, D6

50 -0.8 0.8 0.05 3 Flat disk, D6

51 -100 100 5 500 1000 Flat disk, D6

52 -0.4 0.4 0.05 3 Flat disk, D6

53 -0.6 0.6 0.05 3 Flat disk, D6

54 -0.8 0.8 0.05 3 Flat disk, D6

55 -125 125 5 500 1500 Flat disk, D6

56 -0.4 0.4 0.05 3 Chisel, D6

57 -0.6 0.6 0.05 3 Chisel, D6

58 -0.8 0.8 0.05 3 0 Chisel, D6

59 -75 75 5 500 500 Chisel, D6

60 -0.4 0.4 0.05 3 Chisel, D6

61 -0.6 0.6 0.05 3 Chisel, D6

62 -0.8 0.8 0.05 3 Chisel, D6

63 -100 100 5 500 1000 Chisel, D6

64 -0.4 0.4 0.05 3 Chisel, D6

65 -0.6 0.6 0.05 3 Chisel, D6

66 -0.8 0.8 0.05 3 Chisel, D6

67 -125 125 5 500 1500 Chisel, D6

The curvature values based on the disk-type probe and the corrected values are presented in 

the Figures G.3–G.6, along with those based on flexural rigidity (EI) and the chisel type probe. 

The major observations can be summarized as follows,

1) Measurement with the chisel type probe showed uncertainty in curvature; no defined

trend can be seen in the measured curvatures as compared to those using the disktype probe.
G-5

2) The tests with the chisel- and disk-type probes were not conducted concurrently. The

probe contact condition was changed in the session when the probe was changed,

and the three LVDTs were remounted. The sensor spacing was sensitive to the

change of contact condition.

3) The corrected curvature was shown to be close to the EI-based curvature.

4) Spacing correction (∆h) seemed to exhibit a trend to increase with the level of

displacement input (disp1 on the side of motor 1). However, a large variance of

correction was seen near the small displacement.

5) The level of correction was also far less than those based on the calibration of the

chisel-type probe (2.40 or 2.90 mm).

(a) 0.4 mm
G-6

(b) 0.6 mm

(c) 0.8 mm

Figure G.3 Curvature curves based on disk-type probe, corrected, EI-based, and chiseltype probe, for 0.05 Hz displacement at loading point of U-frame arm: (a) 0.4 

mm, (b) 0.6 mm, and (c) 0.8 mm. EI=42.8 Nm2
G-7

(a) Peak

(b) Valley
G-8

(c) Spacing correction

Figure G.4 (a) Peak and (b) valley responses of curvature waveform, and (c) spacing 

corrections under various disp1 amplitudes at 0.05Hz

(a) 75N
G-9

(b) 100N

(c) 125N

Figure G.5 Curvature curves based on disk-type probe, corrected, EI-based, and chiseltype probe for 5 Hz load at loading point of U-frame arm: (a) 75 N, (b) 100 N, 

and (c) 125 N. EI=42.8 Nm2
G-10

(a) Peak

(b) Valley
G-11

(c) Spacing correction

Figure G.6 (a) Peak and (b) valley responses of curvature waveform, and (c) spacing 

corrections under various load1 amplitudes at 5Hz

G.3.2 Strain gauge as calibration method

The strain gauge was installed on the surface of the bending rod opposite to the three LVDTs, 

as shown in Figure 1. The test conditions using the 0.05 Hz sine wave with varying amplitudes 

are presented in Table 3. The strains were acquired in items 71–73 using a P3 strain indicator 

and recorder (Vishay Micro-Measurement, Raleigh, NC) and in items 81–84 using a strain 

signal conditioning amplifier 2310B (Vishay Micro-Measurement, Raleigh, NC).

Table G.3 Test condition with strain gauge installed

No. pk1 (mm, N) pk2 (mm, N) Frequency Cycles LVDT type

71 -0.4 0.4 0.05 3 Flat disk, D5

72 -0.6 0.6 0.05 3 Flat disk, D5

73 -0.8 0.8 0.05 3 Flat disk, D5

81 -0.4 0.4 0.05 3 Flat disk, D5

82 -0.6 0.6 0.05 3 Flat disk, D5

83 -0.8 0.8 0.05 3 Flat disk, D5

84 -100 100 5 500 Flat disk, D5

The curvatures measured using three LVDTs were converted into the maximum strain at the 

outer fiber of the bending rod. The results of test items 71– 73 are illustrated in Figures G.7 and 

G.8. 
G-12

1) Overall, the strain gauge has a synchronized response with curvature-based strain,

although the curvature-based strain at 0.6 mm lagged in the first half of period.

2) Peak/ valley response of curvature-based strain with the correction applied matched

the strain gauge measurement well.

3) Again, spacing correction (∆h) demonstrated an increasing trend with the

displacements, but the correction at small displacement input showed a lower level

than in the previous test session (Figure G.4).

0.4 mm
G-13

0.6 mm

0.8 mm

Figure G.7 Strain curves based on original disk-type curvature, corrected curvature, and 

strain gauge for 0.05 Hz displacement at the loading point of the U-frame arm: 

(a) 0.4 mm, (b) 0.6 mm, and (c) 0.8 mm
G-14

(a) Peak

(b) Valley
G-15

(c) Spacing correction

Figure G.8 (a) Peak and (b) valley responses of strain waveform, and (c) spacing 

corrections under various disp1 amplitudes at 0.05 Hz

The results of test items 81– 84 are given in Figure G.9 for 0.05 Hz and G.10 for 5 Hz. The 

following observations can be made:

1) Strain gauge showed a lower level of strain that those of curvature-based strain.

2) The spacing correction ∆h and its variation with displacements were smaller

compared with previous tests (Figures G.4 and G.8). The degree of correction on κbased strain was smaller as well.

3) The corrected curvature-based strains were seen close to the EI-based strains.
G-16

(a) Displacement amplitude 0.4 mm

(b) Displacement amplitude 0.6 mm
G-17

(c) Displacement amplitude 0.8 mm

Figure G.9 Strain results based on original curvature, corrected curvature, stain gauge, 

and flexural rigidity EI for 0.05 Hz tests under three displacement amplitudes

(a) Data block 5
G-18

(b) Data block 6

Figure G.10 Strain results based on original curvature, corrected curvature, stain gauge, 

and flexural rigidity EI for 5 Hz tests under amplitude 100N

G.3.3 Summary

The use of SS304 surrogate rod in out-cell study provided the opportunity to examine the small 

curvature measurement currently used in CIRFT tests. The following summarizes the major 

observations:

1) The corrected curvature was shown to be close to the EI-based curvature; the

corrected-curvature-based strain was also demonstrated to agree with the strain

gauge measurement.

2) The spacing correction (∆h) was smaller than those based on a large curvature

measurement. The degree of correction on curvature-based strain was smaller, as

well.

3) The spacing correction exhibited a defined trend to increase with the level of

displacement input, but it also showed a large variance at the small displacements.

4) The level of spacing correction in specified displacement input varied among tests

sessions. The change of contact condition between the probe and the surface of rod

is believed to contribute to the variation.
G-19

G.4 RESULTS ON POLYCARBONATE SURROGATE ROD

The test condition on a polycarbonate rod PC01 consists of two load waveforms: ramp and sine.

G.4.1 Ramp waveform test

The displacement at each loading point of the U-frame ramped up to a designated level and 

then down to zero; one loading cycle was used. The bending in the ramp was unidirectional, 

with a single pule pointed in either a positive or negative direction. 

Two types of LVDT probes were tested: flat disk type and chisel type. Test condition can be 

seen in Table G.4. 

Table G.4 Test condition using ramp process

No. pk1 (mm ) pk2 (mm) rate(mm/s) Cycles LVDT

1 0 6 0.2 1 flat disk, D5

2 0 -6 0.2 1 flat disk, D5

7 0 6 0.2 1 chisel, D5

8 0 -6 0.2 1 chisel, D5

The data sets for a specified probe were assembled to obtain a whole waveform of reversed 

bending. The three LVDT-based curvatures obtained with both disk type and chisel type probes 

are presented, along with the corrected curvature in Figure G.11, in which EI = 1.8 Nm2 was 

used for the EI-based estimate. 

1) The correction procedure suggested that a sensor spacing correction of 2.50 mm is

needed. With the suggested procedure, only one correction is needed instead of two

corrections in the positive (2.90 mm) and negative (2.40 mm) directions as used

before.

2) The corrected curvature showed a good correspondence with the both-chisel type

measurement and the EI-based estimate.
G-20

Figure G.11 Curvature measurements of polycarbonate surrogate rod based on disk- and 

chisel-type probe, 2.50 mm sensor spacing correction, and EI-based estimate

G.4.2 Sine waveform test

Reversed bending was conducted directly by using a 0.05 Hz sine wave, and 12 amplitudes 

were tested with one LVDT probe, the flat disk type. Strain gauge measurement was enabled, 

and strains were acquired using a P3 strain indicator and recorder. The test conditions are 

detailed in Table G.5. 

Table G.5 Test condition using sine wave

No. pk1 (mm) pk2 (mm) F(Hz) Cycles LVDT

35 0.5 -0.5 0.05 3 flat disk, D5

36 1 -1 0.05 3 flat disk, D5

37 1.5 -1.5 0.05 3 flat disk, D5

38 2 -2 0.05 3 flat disk, D5

39 2.5 -2.5 0.05 3 flat disk, D5

40 0.3 -0.3 0.05 3 flat disk, D5

41 0.3 -0.3 0.05 3 flat disk, D5

42 0.3 -0.3 0.05 3 flat disk, D5

43 0.5 -0.5 0.05 3 flat disk, D5

44 0.5 -0.5 0.05 3 flat disk, D5

45 0.5 -0.5 0.05 3 flat disk, D5

46 1 -1 0.05 3 flat disk, D5
G-21

47 1 -1 0.05 3 flat disk, D5

48 1.5 -1.5 0.05 3 flat disk, D5

49 1.5 -1.5 0.05 3 flat disk, D5

50 2 -2 0.05 3 flat disk, D5

51 2 -2 0.05 3 flat disk, D5

52 2.5 -2.5 0.05 3 flat disk, D5

53 2.5 -2.5 0.05 3 flat disk, D5

54 3 -3 0.05 3 flat disk, D5

55 3 -3 0.05 3 flat disk, D5

56 3 -3 0.05 3 flat disk, D5

57 3.5 -3.5 0.05 3 flat disk, D5

58 3.5 -3.5 0.05 3 flat disk, D5

59 3.5 -3.5 0.05 3 flat disk, D5

60 4 -4 0.05 3 flat disk, D5

61 4 -4 0.05 3 flat disk, D5

62 4 -4 0.05 3 flat disk, D5

63 4.5 -4.5 0.05 3 flat disk, D5

64 4.5 -4.5 0.05 3 flat disk, D5

65 4.5 -4.5 0.05 3 flat disk, D5

66 5 -5 0.05 3 flat disk, D5

67 5 -5 0.05 3 flat disk, D5

68 5 -5 0.05 3 flat disk, D5

69 5.5 -5.5 0.05 3 flat disk, D5

70 5.5 -5.5 0.05 3 flat disk, D5

71 5.5 -5.5 0.05 3 flat disk, D5

72 6 -6 0.05 3 flat disk, D5

73 6 -6 0.05 3 flat disk, D5

74 6 -6 0.05 3 flat disk, D5

75 6 -6 0.05 3 flat disk, D5

The typical results from tests with 0.05 Hz sine wave are given in Figure G.12. The peak and 

valley values of the strain waveforms obtained for various cases are shown in Figure G.13, 

including strains based on original curvature (disk-type probe), corrected curvature, and strain 

gauge. The variation of spacing correction as a function of displacement input is also presented. 

1) The strains from the strain gauge usually agreed well with those based on the

corrected curvatures, but they deviated from the corrected curvature-based strain at

high levels.

2) The spacing correction (∆h) again exhibited a defined trend to increase with the

displacement input, and also a larger variance at the small displacements.

3) The spacing corrections were again lower than those obtained in ramping.
G-22

4) The corrected ultimate curvatures obtained in ramp and sine waveforms were the

same (around 2 m-1 at 6.00 mm disp1). There may have been a significant change of

the contact condition between probe and surface of rod, as the tests were carried out

in different periods of the project.

(a)

(b)

Figure G.12 Strain wave forms obtained from original curvature, corrected curvature, 

and strain gauge: (a) displacement amplitude 1.00 mm, and (b) 4.00 mm
G-23

(a) Peak

(b) Valley
G-24

(c) Spacing correction

Figure G.13 (a) Peak and (b) valley responses of strain waveform, and (c) spacing 

corrections under various disp1 amplitudes at 0.05 Hz

G.4.3 Summary

Based on the tests of polycarbonate rod, main observations with respect to disk-type curvature 

measurement and correction can be summarized as follows: 

1) The corrected curvature showed a good correspondence with chisel-type

measurement and EI-based estimate. The agreement of corrected-curvature–based

strain was also found with strain gauge measurement, particularly near small-toimmediate levels of displacement input.

2) The spacing correction (∆h) again exhibited a defined trend to increase with the

displacement input, and also a larger variance at the small displacements.

3) The correction procedure suggested a spacing correction of 2.50 mm (at 6.00 mm

disp1) other than dual corrections reported in ramping (2.40 mm in negative

curvature and 2.90 mm in positive curvature). However, the corrected ultimate

curvatures obtained in ramp and sine waveforms were the same (around 2 m-1 at

6.00 mm disp1).

4) The change of the contact condition between the probe and the surface of the rod in

different test sessions was believed to be responsible for the variation of spacing

correction observed.
G-25

G.5 RECOMMENDATION

The corrected curvature based on disk-type probes agrees to a varying extent with the EI-based 

estimate in both small and large curvature measurements and with chisel-type probe 

measurements in large curvature cases. The corrected-curvature–based strain was also 

illustrated to correspond to the strain gauge measurement, particularly at the small-to-immediate 

level of displacement input.

It is suggested that the proposed correction procedure be applied to data analysis. In the 

monotonic test, ∆h = 2.50 mm obtained from the calibration can be used for spacing correction 

in data analysis. In the reversed cyclic bending, the ∆h can be obtained directly using the data 

block because the peak and valley pair is available.

APPENDIX H HYDRIDE REORIENTATION SAMPLE PREPARATION 

H-1

H.1 INTRODUCTION

The objective of Phase II testing was to collect experimental data on HBR Zircaloy-4 HBU fuels 

following simulated drying-transfer operation conditions. This appendix describes a procedure 

used to reorient hydrides from the circumferential orientation to the radial orientation. Normal 

operation of nuclear fuel in a reactor results in the formation of a waterside corrosion layer 

(oxide) and the introduction of hydrogen into the zirconium cladding via the reaction 2H2O + Zr 

→ ZrO2 + 4H. With increasing corrosion, the hydrogen concentration in the cladding will exceed 

its terminal solid solubility, and brittle zirconium hydrides (Zr + 2H → ZrH2) may precipitate as 

the cladding cools. 

Under pool storage conditions after fuel is discharged from reactors, the precipitated hydride 

platelets are oriented in a circumferential direction. Circumferential hydrides—in combination 

with hardening caused by irradiation effects—will decrease cladding ductility in response to axial 

and hoop loads; however, even with circumferential hydrides, the cladding still retains some 

ductility. At elevated temperatures during drying-transfer operations, some of the hydrogen may 

go into solution (up to 200 wppm at 400 °C). The pressure-induced cladding tensile hoop stress 

during drying-transfer operations is high relative to in-reactor and pool storage conditions. 

During cooling under tensile hoop stress, some of the dissolved hydrogen will precipitate in the 

radial direction across the cladding wall when the hoop stress is sufficient. Further cooling 

during storage may result in radial-hydride–induced embrittlement. This was confirmed by 

recent research [1–5] that featured hydrided Zircaloy-4 cladding and high burnup zirconium alloy 

samples that were exposed to a tensile hoop stress from the internally pressurized cladding at 

≈ 400 °C—conditions similar to drying operations of the SNF. 

APPENDIX H HYDRIDE REORIENTATION SAMPLE PREPARATION 
H-2

H.2 MATERIALS, EQUIPMENT, AND TEST METHODS

H.2.1 Description of Cladding Material and High-Burnup Fuel Segments

The out-of-cell experiments were conducted with unirradiated, 15 × 15 pressurized water reactor 

(PWR) Zircaloy-4 cladding material provided by AREVA; the material has dimensions, oxygen 

content, and mechanical properties similar to those of the HBR cladding. After the cladding was 

received by Oak Ridge National Laboratory (ORNL), measurements were performed to 

determine the outer diameter (OD) and wall thickness. Table H.1 summarizes the dimensions 

and chemical compositions of the unirradiated AREVA-provided clad compared with the nominal 

commercial Zircaloy-4 cladding alloys. 

Table H.1 Dimensions and chemistry of H. B. Robinson (HBR) type Zircaloy-4 

used in the Oak Ridge National Laboratory test program (--- = below the 

detection limit)

Parameter 15 × 15 Zircaloy-4 

HBR-type claddinga

Nominal composition of 

commercial Zircaloy-4b,c

cladding alloys

Outer diameter, mm 10.76 10.72

Wall thickness, mm 0.76 0.617

Tin, wt % --- 1.29 ± 0.1

Niobium, wt % --- --- 

Oxygen, wt % ≈0.13 0.120

Iron, wt % --- --- 

Chromium, wt % --- --- 

Nickel, wppm --- --- 

Sulfur, wppm --- --- 

Carbon, wppm --- --- 

Hafnium, wppm --- --- 

Silicon, wppm --- --- 

Nitrogen, wppm --- --- 

Hydrogen, wppm ≈20 5 

aORNL data based on AREVA lots received in 2014.

bASTM B811. Source: M. C. Billone, T. Burtseva, and Y. Yan. “Overview of Spent Nuclear 

Fuel Program: Test Plan and High Burnup Cladding at ANL,” NRC Program Review 

Meeting, Argonne National Laboratory, Argonne, IL, July 7, 2010.

cORNL/TM-9591/V1&R1, “Physical and Decay Characteristics of Commercial LWR Spent 

Fuel,” January 1986.
H-3

Irradiated materials used in the ORNL test program are HBU HBR PWR fuel rods received by 

ORNL in 2008. They were from a 15 × 15 assembly of the HBR plant Unit 2 [6] that was 

operated for seven cycles and reached a rod-average burnup of 67 GWd/MTU (73 GWd/MTU 

peak pellet). The fuel enrichment is 2.90%. The nominal fuel pellet dimensions are 9.06 mm 

diameter × 9.93 mm height, and the active fuel height is 3.66 m. The cladding is coldworked/stress-relieved Zircaloy-4 measuring 10.77 mm OD and 9.25 mm inner diameter (ID), 

with a nominal tin content of 1.42%. The rods were pressurized with helium to 2.0 MPa during 

initial fabrication. A detailed description of the as-fabricated cladding, the irradiation history, and 

the nondestructive testing results (eddy current, profilometry, fission-gas release, etc.) is 

provided in a 2001 Electric Power Research Institute report [6].

Irradiated HBR fuel used in the ORNL test program is summarized in Table H.2. Detailed

characterization was performed to determine the fuel morphology, fuel-cladding bond, corrosion 

layer, and hydride morphology. Figure H.1 shows a low magnification image of the fuel 

morphology, which reveals the typical reactor start-up and shut-down cracks. 

Table H.2 Characteristics of HBU HR fuel rods for tORNL hydride reorientation (HR) testa

Parameter H. B. Robinson

Reactor Pressurized water reactor

Enrichment, wt % 2.90

Rod average burnup, GWd/MTU 63–67

Discharge date 1995

Fast fluence, 1025 n/m2 14

Cladding 15 × 15 Zircaloy-4 

Initial wall thickness, mm 0.76

OD oxide, μm ≤ 100

Hydrogen pickup, wppm ≤ 800

Fueled Yes
H-4

Figure H.1 HBU HBR fuel morphology for Sample 607D4A.

H.2.2 Equipment for Sample Preparation

Hydride reorientation (HR) tests were conducted with Zircaloy-4 cladding samples 152 mm in 

length (see Table H.1 for cladding geometry) which were hydrogen charged for a target 

hydrogen concentration of 300 wppm. After hydrogen charging, the hydrided samples were 

subjected to HR to simulate drying-transfer operation conditions. This which consists of a high 

pressure piping system and test chamber within a programmable crucible furnace. The hydrided 

sample was sealed (end-cap welded) at one end and connected to the high pressure piping 

system within the test chamber at the other end. The piping system was used to internally 

pressurize the sample to cause a hoop stress ranging up to 150 MPa.

The hydriding apparatus consists of a closed stainless steel tube that contains zirconium alloy 

specimens and hydrogen gas, as shown in Figure H.2. The tube is heated to facilitate hydrogen 

absorption by the metal. By controlling the initial hydrogen gas pressure in the vessel and the 

temperature profile, hydrogen concentration can be processed selectively. A typical pressure 

and temperature profile is shown in Figure H.3.
H-5

Figure H.2 Hydrogen charging apparatus

Figure H.3 Typical pressure and temperature profile of the hydrogen charging process

An Astro Arc Polysoude welder (see Figure H.4) was used for automatic orbital welding of the 

cladding. A circular movement of the welding torch/electrode around fuel cladding, controlled by 

the welder power source, creates an orbital weld. This welder can be operated remotely with 

manipulators in a hot cell. 
H-6

Figure H.4 An Astro Arc Polysoude tubing welder

H.2.3 Hydride Reorientation System

The reorientation system consists of a high pressure piping system and test chamber within a

programmable crucible furnace (see Figure H.5). The system was installed into the hot cell in 

FY16. Calibrations and benchmark tests were performed with unirradiated materials. For each 

experiment, a specimen 6 inches in length was welded with two end plugs, with one end sealed 

and the other end connected to the high-pressure piping system within the test chamber. The 

test chamber was purged with pure argon gas to remove air and moisture contents, and then 

the sample was internally pressurized to cause a hoop stress ranging from 0–150 MPa [5]. 

Pressurized samples were subjected to a preprogrammed temperature profile for up to five 

cycles, with each cycle being heated to 400 °C, held for 2 hours, and slow-cooled/heated (1

°C/min) to 170 °C; the final cycle cooled from 170 °C to room temperature (RT). Hoop stress 

was calculated using values r = 4.62 mm for inner radius and 

t = 0.76 mm for the wall thickness of the HBR Zircaloy-4 cladding.
H-7

(a)

(b)

Figure H.5 The HR system (a) before and (b) after being installed in the hot cell

H.3 MATERIALS HYDRIDING AND CHARACTERIZATION

The Zircaloy-4 cladding samples were hydrogen charged by the gaseous method as described 

in Section H.2. Gas charging was conducted under stress-free conditions. For this work, the 
H-8

Zircaloy-4 was hydrided to a target concentration of 300–400 wppm, which is similar to the 

hydrogen content of HBR HBU cladding to be tested in a hot cell.

Metallographic examinations were performed on hydrided Zircaloy-4 samples. As shown in 

Figure H.6, the hydride density increases as the hydrogen concentration in the sample 

increases. Hydrogen measurements indicated that the hydrogen content of the specimens 

shown in Figure H.6 range from 70–320 wppm.

Figure H.6 Micrographs showing typical, uniform circumferential hydride distributions

in hydrogen charged ry-4 
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H-10

Figure H.8 High magnification micrograph showing radial hydrides of Sample HR-HBR#1 

with hydrogen content H ≈ 274 wppm; the specimen was sectioned at the 

midplane of a sample 6 inches in length
Figure H.9 Radial hydride distribution shown in eight areas in different circumferential directions of Sample HR-HBR#1 the 

specimen was sectioned at the midplane of a sample 6 inches in length

H11
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H-13

Figure H.11 High magnification micrograph showing radial hydrides of Sample HRHBR#2 with hydrogen content H ≈ 286 ppm. The specimen was sectioned at 

the midplane of a sample 6 inches in length
Figure H.12 Radial hydride distribution shown in eight areas in different circumferential directions of Sample HR-HBR#2;

the specimen was sectioned at the midplane of a sample 6 inches in length

H14
H-15

H.4.3 HR-HBR#3, 100 MPa at 400°C, five cycles 

Test HR-HBR#3 was conducted on a hydrided 15 × 15 Zircaloy-4 specimen at the maximum 

hoop stress of 100 MPa and the hold temperature T = 400 °C. The specimen was assembled 

into a holder within the furnace for heating to 400 °C, held at 400 °C for 2 hours, cooled at 

1°C/min to 170 °C, and then heated at 1 °C/min to 400 °C again for five cycles. The sample was 

furnace cooled from 170 °C to RT for the last cycle. Figure H.13 shows the temperature and 

pressure history for HR-HBR#3. The plotted temperature is a thermocouple reading near the 

specimen surface. 

Figure H.14 shows the hydride morphology of Sample HR-HBR#3 with hydrogen content 

H ≈ 310 wppm. The specimen was sectioned at the midplane of a sample 6 inches in length. 

The radial hydrides can be observed clearly in Figure H.14. The maximum length of the radial 

hydride is < 30 μm. The radial hydride ratio of HR-HBR#3 also is lower than HR-HBR#2. Figure 

H.15 shows radial hydride distribution in eight areas in different circumferential directions of

Sample HR-HBR#3. 

Figure H.13 Pressure and sample temperature as a function of time for Test HR-HBR#3
H-16

Figure H.14 High magnification micrograph showing radial hydrides of Sample HRHBR#3 with hydrogen content H ≈ 310 ppm; the specimen was sectioned at the 

midplane of a sample 6 inches in length
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H-19

Figure H.17 High magnification micrograph showing radial hydrides of Sample HRHBR#4 with hydrogen content H ≈ 312 ppm; the specimen was sectioned at 

the midplane of a sample 6 inches in length
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H-21

H.5 DISCUSSION

In this work, the hydrided tubing samples have been pressurized internally to produce a hoop 

stress up to 150 MPa to form radial hydrides using the HR system (see Section H.4). Tests 

HR-HBR#1 and HR-HBR#2 indicate that multithermal cycling will increase the ratio of radial 

hydrides under the same maximum temperature and maximum internal pressure. In addition, 

the results of Tests HR-HBR#2 through HR-HBR#4 reveal that the ratio of radial hydrides 

decreases as the internal pressure decreases when the temperature and pressure profiles 

remain the same. 

The goal of Phase II testing was to investigate the impact of radial hydrides on HBU fuel 

cladding ductility. To gain insight into bounding effects of radial hydrides, it was desirable to 

produce a reasonably high percentage of radial hydrides within the cladding by the HR tests. 

The out-of-cell tests were important to support planning of the in-cell HR tests, as they 

established a test procedure and provided a guideline for in-cell test conditions. Based on the 

four out-of-cell tests, it was decided to perform in-cell HR treatment (HRT) test under the same 

test conditions of HR-HBR#2 to maximize the radial hydride condition.

H.5.1 Hydride Reorientation Treatment for Phase II HBU Tests

Four CIRFT evaluations were completed on the HBR specimens following HRT. Five cycles at 

140 MPa hoop stress and 400 °C were conducted for segments HR1–HR3. Due to a welded 

endcap leak, only thermal heat treatment was applied to the HR4 sample.

The HBU fuel samples for Phase II tests were prepared in the irradiation fuel examination 

laboratory. After the samples were sectioned into pieces 6 inches in length, the surface oxide 

layer and fuel material were removed from both ends to a depth of approximately 0.5 inches. 

Endcaps were welded at both ends, and the fuel specimens were then pressurized with argon 

gas at a maximum hoop stress level of 140 MPa and a target hold temperature (T = 400 °C). 

Thermal cycling was performed to increase the radial hydrides. All in-cell tests had the same 

temperature profiles. Figure H.19 shows the temperature history of the first in-cell HR test, HR1, 

with an HBU HBR sample. The same procedure was used for HR2 and HR3. Due to a welded 

endcap leak, only thermal heat treatment was applied to the HR4 sample.
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H-23

(a)

(b)

Figure H.20 (a) High magnification micrograph showing radial hydrides of HBU HBR 

Sample HR-1 (H ≈ 360–400 ppm) following HRT; the HBU HBR specimen was 

pressurized to 140 MPa at 400 °C with five thermal cycles; (b) high 

magnification micrograph showing radial hydrides of the unirradiated 

Sample HR-HBR#2 (H ≈ 286 ppm) following the same HRT procedure 
H-24

H.6 REFERENCES

1. M. C. Billone, T. Burtseva, and Y. Yan. “Overview of Spent Nuclear Fuel Program: Test Plan

and High Burnup Cladding at ANL,” NRC Program Review Meeting, Argonne National

Laboratory, Argonne, IL, July 7, 2010.

2. R. Daum, H. Tsai, Y. Liu, and M. Billone, “High-Burnup Cladding Mechanical Performance

during Cask Storage and Post-Storage Handling and Transportation,” Nuclear Safety

Research Conference, Washington, DC, October 25–27, 2004.

3. Y. Yan et al., “Observation and Mechanism of Hydride in Zircaloy-4 and Local Radial

Hydride Induced by High Pressure at High Temperatures,” 2013 International High-Level

Radioactive Waste Management, Albuquerque, NM, April 28–May 2, 2013.

4. Y. Yan, L. K. Plummer, H. Ray, T. Cool, and H. Z. Bilheux, “Ductility Evaluation of AsHydrided and Hydride Reoriented Zircaloy-4 Cladding under Simulated Dry-Storage

Condition,” 2013 MRS Fall Meeting Proceeding, Boston, MA, December 1–6, 2013.

5. M. C. Billone, T. A. Burtseva, and Y. Yan, Ductile-to-Brittle Transition Temperature for HighBurnup Zircaloy-4 and ZIRLOTM Cladding Alloys Exposed to Simulated Drying-Storage

Conditions, Argonne National Laboratory letter report to the Nuclear Regulatory

Commission, ANL-13/13, September 28, 2013. Available online as ML12181A238 at

http://www.nrc.gov/reading-rm/adams.html.

6. E. J. Ruzauska and K. N. Fardell, “Design, Operation, and Performance Data for High

Burnup PWR Fuel from the H. B. Robinson Plant for Use in the NRC Experimental Program

at Argonne National Laboratory,” Electric Power Research Institute report 1001558, May

2001.






NUREG/CR-7198, Rev. 1 Mechanical Fatigue Testing of High-Burnup Fuel for Transportation Application October 2017