Mechanical Fatigue Testing 

of High-Burnup Fuel for 

Transportation Applications

Office of Nuclear Regulatory Research

NUREG/CR-7198

ORNL/TM-2014/214
NRC Reference Material 

As of November 1999, you may electronically access 

NUREG-series publications and other NRC records at 

NRC’s Public Electronic Reading Room at 

http://www.nrc.gov/reading-rm.html. Publicly released 

records include, to name a few, NUREG-series 

publications; Federal Register notices; applicant, 

licensee, and vendor documents and correspondence; 

NRC correspondence and internal memoranda; bulletins 

and information notices; inspection and investigative 

reports; licensee event reports; and Commission papers 

and their attachments. 

NRC publications in the NUREG series, NRC 

regulations, and Title 10, “Energy,” in the Code of

Federal Regulations may also be purchased from one 

of these two sources. 

1. The Superintendent of Documents

 U.S. Government Printing Office

Mail Stop SSOP

Washington, DC 20402–0001

Internet: bookstore.gpo.gov

Telephone: 202-512-1800

Fax: 202-512-2250 

2. The National Technical Information Service 

Springfield, VA 22161–0002

www.ntis.gov

1–800–553–6847 or, locally, 703–605–6000 

A single copy of each NRC draft report for comment is 

available free, to the extent of supply, upon written 

request as follows:

Address: U.S. Nuclear Regulatory Commission

 Office of Administration 

Publications Branch

 Washington, DC 20555-0001

E-mail: DISTRIBUTION.RESOURCE@NRC.GOV

Facsimile: 301–415–2289 

Some publications in the NUREG series that are 

posted at NRC’s Web site address 

http://www.nrc.gov/reading-rm/doc-collections/nuregs 

are updated periodically and may differ from the last 

printed version. Although references to material found on 

a Web site bear the date the material was accessed, the 

material available on the date cited may subsequently be 

removed from the site.

Non-NRC Reference Material 

Documents available from public and special technical 

libraries include all open literature items, such as books, 

journal articles, transactions, Federal Register notices, 

Federal and State legislation, and congressional reports. 

Such documents as theses, dissertations, foreign reports 

and translations, and non-NRC conference proceedings 

may be purchased from their sponsoring organization. 

Copies of industry codes and standards used in a 

substantive manner in the NRC regulatory process are 

maintained at—

The NRC Technical Library

Two White Flint North

11545 Rockville Pike

Rockville, MD 20852–2738 

These standards are available in the library for reference 

use by the public. Codes and standards are usually 

copyrighted and may be purchased from the originating 

organization or, if they are American National Standards, 

from—

American National Standards Institute

11 West 42nd Street

New York, NY 10036–8002

www.ansi.org

212–642–4900 

AVAILABILITY OF REFERENCE MATERIALS

IN NRC PUBLICATIONS

Legally binding regulatory requirements are stated only 

in laws; NRC regulations; licenses, including technical 

specifications; or orders, not in NUREG-series 

publications. The views expressed in contractorprepared publications in this series are not necessarily 

those of the NRC.

The NUREG series comprises (1) technical and 

administrative reports and books prepared by the staff 

(NUREG–XXXX) or agency contractors (NUREG/CR–

XXXX), (2) proceedings of conferences (NUREG/CP–

XXXX), (3) reports resulting from international 

agreements (NUREG/IA–XXXX), (4) brochures 

(NUREG/BR–XXXX), and (5) compilations of legal 

decisions and orders of the Commission and Atomic and 

Safety Licensing Boards and of Directors’ decisions

under Section 2.206 of NRC’s regulations (NUREG–

0750).

DISCLAIMER: This report was prepared as an account 

of work sponsored by an agency of the U.S. 

Government. Neither the U.S. Government nor any 

agency thereof, nor any employee, makes any warranty, 

expressed or implied, or assumes any legal liability or 

responsibility for any third party’s use, or the results of 

such use, of any information, apparatus, product, or 

process disclosed in this publication, or represents that 

its use by such third party would not infringe privately 

owned rights.
Mechanical Fatigue Testing 

of High-Burnup Fuel for 

Transportation Applications

Manuscript Completed: August 2014

Date Published: May 2015

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 Codes N6686/N6789

Office of Nuclear Regulatory Research

NUREG/CR-7198

ORNL/TM-2014/214

iii 

ABSTRACT 

This report describes testing designed to determine the ability of high burnup (HBU) (>45 

GWd/MTU) spent fuel to maintain its integrity under normal conditions of transportation. An 

innovative system, Cyclic Integrated Reversible-bending Fatigue Tester (CIRFT), has been 

developed at Oak Ridge National Laboratory (ORNL) to test and evaluate the mechanical 

behavior of spent nuclear fuel (SNF) under conditions relevant to storage and transportation. 

The CIRFT system is composed of a U-frame equipped with load cells for imposing the pure 

bending loads on the SNF rod test specimen and measuring the in-situ curvature of the fuel rod 

during bending using a set up with three linear variable differential transformers (LVDTs). 

The HBU H.B. Robinson (HBR) SNF rods with Zircaloy-4 (Zry-4) cladding and intact fuel were 

studied under both static and dynamic (cyclic) bending. The HBR HBU rods survived static 

unidirectional bending to a maximum curvature of 3.1 to 3.5 m-1, or a maximum moment of 86 to 

87 N·m. The maximum longitudinal cladding offset strain before failure or before reaching 

CIRFT displacement capacity was 1.7 to 1.9 %, and the maximum bending stress was 733 to 

748 MPa. The 0.2% offset yield stress of the HBU irradiated clad was estimated to be 687 to 

727 MPa based on the equivalent stress-strain curves. 

In general, the static CIRFT test results indicate a significant increase in flexural rigidity 

compared to that of the defueled HBU rod specimen. Nevertheless, the segment composite 

structure of a HBU rod also introduces numerous stress concentration sites into a HBU rod 

system, ultimately resulting in HBU specimen fractured at the pellet-pellet interface regions 

under dynamic CIRFT fatigue testing. 

The majority of the HBR dynamic test samples failed at or near a pellet-pellet interface. The 

fatigue life of HBU HBR rods during dynamic testing mainly depended on the level of loading 

and fuel pellet-clad interaction and the conditions of the rod. Under loading with moments of 

±8.89 to ±35.56 N·m, resulting in an equivalent strain  of ±0.07 to ±0.49% 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 HBR fuel cladding system, it was significant to find that the -N data for 

the HBU HBR was characterized by a curve that would be expected of standard uniform 

materials. The -N curve of the HBU HBR rods can be described by a power function of y = 

3.5693·x0.252, 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, then 

the HBU HBR exhibited an endurance limit. The endurance limit for HBR HBU is likely located 

at approximately 0.1% strain. 

The maxima of the imposed curvature  during dynamic testing ranged from ±0.16 to ±1.19 m1

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

=8.1941·x0.252, where x is the number of cycles to failure, and y is the maxima of clad tensile 

curvature |max (m-1). An endurance limit is likely located at approximately 0.2 m1

 when it is 

defined at 107

 cycles. 

v 

FOREWORD 

The U.S. Nuclear Regulatory Commission (NRC) requires the evaluation of nuclear fuel 

performance under normal and accident conditions, as specified in 10 CFR Parts 71 and 72, as 

part of spent nuclear fuel (SNF) storage and transportation cask certification. The certification of 

SNF storage and transportation systems requires a demonstration of safety and security. As 

part of demonstrating safety, the applicant must make an assessment about the condition of the 

SNF. 

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

permanent disposal. During this transport phase, transportation casks will experience some 

level of oscillation associated with normal conditions of transport (NCT). This oscillation will be 

transmitted to the contents of the casks and the SNF assemblies, resulting in localized stresses. 

Nuclear fuel is now regularly irradiated to high-burnup (HBU) values (>45 GWd/MTU). Several 

phenomena associated with HBU fuel have been identified, such as hydride reorientation that 

could impact the fatigue resistance of SNF under normal conditions of storage and 

transportation. 

For these reasons, the fatigue resistance of HBU fuel should be investigated. The NRC has 

tasked Oak Ridge National Laboratory to investigate the flexural rigidity and fatigue life of HBU 

fuel. The testing was designed to evaluate the fuel rod as a system, including the presence of 

intact fuel inside the cladding and any fuel/cladding bonding effects. The properties measured in 

this testing program can be used in the evaluation of SNF integrity under NCT, when combined 

with details of an SNF cask design and expected transportation loading conditions. 

The research effort was not intended to quantitatively evaluate SNF failure modes during NCT 

with respect to regulatory requirements. The results of this study are particular to the specific 

fuel tested. The results are not intended to be generically applicable and may not be inclusive of 

all cladding types and fuel burnup levels. 

vii 

CONTENTS 

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

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

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

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

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

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

LIST OF ACRONYMS AND UNITS ........................................................................................... xvii

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

2. EXPERIMENTAL METHOD .................................................................................................. 3

2.1 THE U-FRAME SETUP ................................................................................................ 3

2.2 GRIP DESIGN .............................................................................................................. 6

2.3 LVDT MEASUREMENTS AND CURVATURE CALCULATION ................................... 9

2.4 TEST SEGMENTS ..................................................................................................... 12

2.5 TEST PROTOCOL ..................................................................................................... 14

2.5.1 Static Testing .................................................................................................. 14

2.5.2 Dynamic Testing ............................................................................................. 14

2.6 DATA PROCESSING ................................................................................................. 15

2.6.1 HBU SNF Characteristics ............................................................................... 16

3. STATIC TESTING ............................................................................................................... 21

3.1 RESULTS ................................................................................................................... 21

3.2 DISCUSSION ............................................................................................................. 27

3.2.1 Characteristics of Moment-Curvature Curve ................................................... 27

3.2.2 Comparison of Static Results with PNNL Cladding Data ................................ 30

3.2.3 Effect of LVDT Probe Contact on Large Curvature Measurement .................. 32

4. DYNAMIC TESTING ........................................................................................................... 35

4.1 RESULTS ................................................................................................................... 35

4.2 DISCUSSION ............................................................................................................. 47

4.2.1 -N Curve ........................................................................................................ 47

4.2.2 Post-Irradiation Examination (PIE) .................................................................. 51

5. CONCLUSIONS .................................................................................................................. 55

6. REFERENCES .................................................................................................................... 57

APPENDIX A.................................................................................................................................................    
viii 

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
ix 

LIST OF FIGURES 

Figure Page 

Figure 1. (a) Horizontal layout of ORNL U-frame setup; (b) rod specimen under test 

and three LVDTs for curvature measurement (operator is facing the three 

LVDTs); and (c) front view of CIRFT installed in ORNL hot cell. .............................. 5

Figure 2. Schematic drawings of U-frame setup for reversal bending when rigid 

arms are (a) closing, (b) neutral, and (c) opening. ................................................... 5

Figure 3. Assembled grip (left) and two views (right) of detached grip body. .......................... 6

Figure 4. Image showing the grip design of CIRFT with one end-block removed. .................. 7

Figure 5. (a) Vise mold for assembling octagonal rigid sleeves into a rod; (b) rod 

(Demo1, 606B2) to be inserted; (c) one pin was turned away and a rigid 

sleeve loaded with epoxy inserted into the lower chamber. ..................................... 8

Figure 6. Determination of the bending curvature of the rod by use of deflections 

measured at three points. ......................................................................................... 9

Figure 7. (a) Diagrams showing the effect of the LVDT probe with disk shape on 

curvature measurement, where the contact points are shifted. 

Comparisons of curvature measurement between disk and chisel head 

probe for LVDTs located at clad compression and tension site are shown 

in (b) and (c), respectively. (c). ............................................................................... 11

Figure 8. Segment 608C4B for DL2 taken from rod G10, with the storage tube. .................. 12

Figure 9. Flowchart for cyclic testing of spent fuel rod. ......................................................... 15

Figure 10. 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. ........................................................ 16

Figure 11. “Empty-run” conducted on HBR S1, for which the load and displacement 

were plotted. ........................................................................................................... 19

Figure 12. (a) Curve of moment versus curvature, and (b) equivalent strain–stress 

curve for S1 (606C3C). .......................................................................................... 22

Figure 13. Mating fracture surfaces (a) and (b); lateral sides near the neutral axis of 

the bending rod for S1 (606C3C) (c) and (d). The rod was subjected to 

four repeat static cycles to 24 mm maximum relative displacement 

followed by dynamic cyclic loading ±25.40 to 30.48 N·m at 5 Hz; 1.4×104

failure cycles, 0.9 gram fuel particles collected. ..................................................... 23

Figure 14. Moment-curvature curves measurements in static tests (Corresponding 

stress and strain displayed on right and top axes, respectively). ........................... 24

Figure 15. Characteristic points of moment-curvature curve. .................................................. 28

Figure 16. Comparison of CIRFT global data with PNNL moment-curvature curve 

converted from PNNL cladding stress-strain data. ................................................. 31

Figure 17. Moment-curvature curves based on stainless steel alumina pellet (SSAP) 

rod testing when rod is bending with respect to three LVDTs (a) 

concavely and (b) convexly. ................................................................................... 32

Figure 18. Moment-curvature curves as a function of sensor spacing adjustment 

when rod is bending with respect to three LVDTs (a) concavely and (b) 

convexly. The unit of dh is mm. .............................................................................. 33

Figure 19. 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). ............................................................................................................... 37
x 

Figure 20. (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 21. 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 22. 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 = 1.1×105

 cycles under ±15.24 

N·m, 5 Hz. Fuel particles collected <1.0 g. ............................................................. 39

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

at (a) 26 and (b) 1.10×105

 cycles; results are based on online monitoring. ........... 40

Figure 24. (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. ..................................................................................................... 41

Figure 25. Equivalent strain amplitudes (/2) as a function of number of cycles. .................. 46

Figure 26. Equivalent strain amplitudes (/2) as a function of number of cycles with 

curve fitting extended to include the data points without failure. Solid 

markers represent tests with specimens failure; open markers indicate the 

tests stopped without failure. .................................................................................. 47

Figure 27. (a) Maxima of absolute curvature extremes and curvature amplitudes as a 

function of number of cycles, (b) Maxima of absolute strain extremes and 

strain amplitudes as a function of number of cycles. ( ........................................... 49

Figure 28. (a) Curvature amplitudes, (b) maxima of absolute curvature extremes, 

and (c) maxima of absolute strain extreme as a function of number of 

cycles with curve-fitting extended to include the no-failure data points. ................. 50

Figure 29. Optical images of untested segment 605D for (a) overall axial cross 

section and (b) enlarged area A. ............................................................................ 52

Figure 30. (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, B – cracking of oxide layer. ............. 53
xi 

LIST OF TABLES 

Table Page 

Table 1. Physical and mechanical properties of related materials ............................................ 7

Table 2. Specifications of specimens used in hot-cell testing ................................................. 13

Table 3. Summary of static tests ............................................................................................. 25

Table 4. Characteristic points and quantities based on curvature-moment curves ................. 29

Table 5. Characteristic points and quantities based on equivalent stress-strain curves ......... 29

Table 6. Comparison of flexural rigidity results in CIRFT tests and PNNL data ...................... 32

Table 7. Summary of dynamic test results .............................................................................. 42

Table 8. Failure modes of rods from dynamic tests ................................................................ 43

xiii 

EXECUTIVE SUMMARY 

This report describes testing designed to determine the ability of high burnup (HBU) (>45 

GWd/MTU) spent fuel to maintain its integrity under normal conditions of transportation. 

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

HBU SNF rods. The proposed system is composed of a U-frame equipped with load cells for 

imposing the pure bending loads on the SNF rod test specimen and measuring the in-situ

curvature of the fuel rod during bending using a set-up of three linear variable differential 

transformers (LVDTs). 

By August 2013, ORNL had completed all the modifications needed for the testing system and 

the associated vise mold. On August 19, 2013, the completed testing system was installed in 

the Building 3525 hot cell and formally named the Cyclic Integrated Reversible-bending Fatigue 

Tester (CIRFT). After tuning the test system and benchmark testing in September 2013, ORNL 

completed four static tests under displacement control at the rate of 0.1mm/s to a maximum 

displacement of 12.0 mm. The benchmark and static testing results were critically reviewed at 

the NRC-ORNL review meeting in early November 2013. The cyclic testing was then initiated. 

At this writing, 16 cyclic tests have been conducted in the hot cell. Load ranges varied from 

±5.08 to ±35.56 N·m; there were 12 tests with failure and 4 tests without failure. One of the 

cyclic tests reached 1.3 × 107

 cycles and was stopped. 

In general, the static CIRFT results indicate a significant increase in flexural rigidity compared to 

the results for the defuel HBU rod specimen. Nevertheless, the segment composite structure of 

an HBU rod also introduces numerous stress concentration sites into an HBU rod system, 

ultimately resulting in HBU specimen fractures at the pellet-pellet interface regions under 

dynamic CIRFT fatigue testing. 

The following observations were noted: 

 The HBU HBR rods survived static unidirectional bending to a maximum curvature of 3.1 to 

3.5 m1

, or a maximum moment of 86 to 87 N·m. The maximum equivalent strain was 1.7 to 

1.9 %, corresponding to equivalent stress of 733 to 748 MPa. 

 The HBU HBR rods exhibited a multiple-stage constitutive response, with the two linear 

stages followed by a nonlinear stage. The flexural rigidity at the initial stage was 52 to 63 

N·m2

, corresponding to an elastic modulus of 83 to 101 GPa. The flexural rigidity at the 

second stage was 37 to 41 N·m2

, and the corresponding elastic modulus was 59 to 66 GPa. 

 The equivalent 0.2 % yield strength of the HBU irradiated cladding was estimated to be 687 

to 727 MPa based on the equivalent stress-strain curves. 

 Most of the failures of HBU HBR rods under static unidirectional loading were coincident 

with the pellet-to-pellet interface, as validated by the outcrop of pellet end faces in most of 

the fracture surfaces. Fragmentation of the pellets also occurred to a limited degree, along 

with cladding failure. 

 The fatigue life of HBU HBR rods in the cyclic test depended on the level of loading. Under 

loading with moments of ±8.89 to ±35.56 N·m—namely ±0.07 to ±0.49% strain at 5 Hz—the 

fatigue life ranged from 5.5 × 103

 to 2.3 × 106 cycles. 
xiv 

 The -N curve of the HBU HBR rods can be described by a power function of y = 

3.5693·x0.252, where x is the number of cycles to failure and y is the strain amplitude (%). 

 Maxima of the imposed curvature during dynamic tests ranged from ±0.16 to ±1.19 m1

 at 5 

Hz. The -N curve of the HBU HBR rods can be described by a power function of y = 

8.1941·x0.252, where x is the number of cycles to failure and y is the maxima of cladding 

tensile curvature |max (m-1). A fatigue limit is likely located between 0.226 and 0.245 m1

 if it 

is defined at 107

 cycles. 

 The failure of HBU HBR rods under cyclic reverse loading often occurred near pellet-topellet interfaces, as validated by the outcrop of clean pellet end faces. 
xv 

ACKNOWLEDGMENTS 

This research was sponsored by the U.S. 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; Chuck Baldwin for providing post-irradiation examination support; 

Brian Woody and Scot Thurman for providing hot-cell operation support; and Thomas Cox, 

Brian Sparks, and Randy Parten for providing drawing and machining support. 

xvii 

LIST OF ACRONYMS AND UNITS 

CFR Code of Federal Regulations 

CIRFT Cyclic Integrated Reversible-bending Fatigue Tester 

EPRI Electric Power Research Institute 

FEM finite element method 

GWd gigawatt-day 

HBR H. B. Robinson 

HBU high-burnup 

LVDT linear variable differential transformer 

MTU metric ton of uranium 

NCT normal conditions of transport 

N·m newton-meter 

NRC U.S. Nuclear Regulatory Commission 

ORNL Oak Ridge National Laboratory 

PCI pellet-cladding interface 

PIE post-irradiation examination 

PNNL Pacific Northwest National Laboratory 

PPI pellet-pellet interface 

PWR pressurized water reactor 

SEM scanning electron microscope 

SNF spent nuclear fuel 

SS stainless steel 

SSAP stainless steel alumina pellet 

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 

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 

, A, B, C curvature, m-1 

max k || maximum clad tensile curvature, 

M, MA, MB, MC moment, N·m 
xviii 

∆M moment range, N·m 

 curvature range, m-1

A BC engineering stress, MPa 

 A, B, C engineering strain, % 

ymax maximum y of cross section of rod, m 
1 

1. INTRODUCTION 

Spent nuclear fuel (SNF) is expected to be transported at least once before permanent disposal. 

During this transportation phase, transport casks will experience a certain level of oscillation due 

to normal conditions of transport (NCT). This oscillation will be transmitted to the contents of the 

casks and SNF assemblies, resulting in localized stresses in the fuel rods. SNF is now regularly 

irradiated to high-burnup (HBU) values (>45 GWd/MTU). Several phenomena associated with 

HBU fuel have been identified, such as hydride reorientation 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 

For these reasons, the response of SNF to NCT is being investigated. The U.S. Nuclear 

Regulatory Commission (NRC) has tasked Oak Ridge National Laboratory (ORNL) to 

investigate the bending stiffness and fatigue life of HBU fuel during transportation. The testing 

was designed to evaluate the fuel rod as a system, including the presence of fuel pellets and 

any fuel/cladding bonding effects. The objective is to report the results of vibration fatigue 

testing of HBU SNF rods and the effect of the presence of the fuel pellets 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). Using fuel rod segments from 5 separate rods, 4 static tests were conducted, 

followed by 16 cyclic fatigue tests in a range of load amplitudes. The burnup of the tested rod 

segments was ranged between 63.8 and 66.8 GWd/MTU. The estimated oxide layer thickness 

on the waterside surface was approximately 40 to 110 m, and corresponding hydrogen 

contents of hydrides within the cladding wall was estimated between 360 and 800 wppm. All of 

the fuel rod segments were tested in an as-received status in which there is no oxide layer 

removal, and most hydrides are 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, it is expected that SNF rod fatigue strength may not be sensitive to the 

hydride orientation if there is no significant clad internal pressure. Nevertheless, the potential 

effect of radial hydride reorientation on the bending fatigue response of the HBU fuel rods is 

being investigated and will be reported in future. 

The static tests defined the elastic and plastic responses of the fuel rod system, and the 

dynamic tests defined the fatigue life of the fuel rod system under constant loads. The effects of 

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

understanding of the fatigue life of HBU fuel. Relations of equivalent strain versus the number of 

cycles to failure (-N) or curvature versus the number of cycles to failure (-N) were developed 

to estimate the influence of the imposed bending on the fatigue strength. The impacts of 

important fuel rod phenomena such as fuel-to-cladding bonding efficiency and the pellet-topellet pinning interaction were recognized. The effects of the stress concentration at pellet-pellet 

interfaces and SNF system interfacial bonding status, including perfect bonding and de-bonding, 

on the global flexural rigidity and the local plastic deformation were studied by using finite 

element method (FEM), and the implication of numerical results on the interpretation of bending 

testing was discussed. 

 

3 

2. EXPERIMENTAL METHOD 

The objective of this testing program is to develop testing protocols to evaluate the bending 

deformation responses and the associated fatigue mechanical properties for HBU fuel rods 

relevant to transportation conditions. Thus, the HBU fuel rods would be tested in a condition and 

geometry as close as possible to normal conditions of transport (NCT). Currently, most of the 

mechanical testing devices require a specified specimen design, such as a pre-notch or 

reduced gage section, to prevent failure at the grip location and ensure meaningful data. For 

this testing program, it was recommended that the test segment not be altered by machining a 

gage section. Therefore, an innovative U-frame and grip design were developed. Key features 

of the U-frame and grip design will be discussed in the sections that follow. Cost considerations 

also had a significant influence on the development of the testing program. Hot-cell space and 

time are extremely limited and costly, as is testing material. For this testing program, we wanted 

to use a small test segment, design a compact test device, and develop a test protocol for 

efficient testing. Important features of the test segments and test protocol will be discussed later 

in this report. Finally, the curvature response of the test segment needed 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 that the curvature of the specimen could be calculated without interpretation or 

assumptions regarding 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 Section 2.3. 

2.1 THE U-FRAME SETUP 

Testing of the SNF rod specimen was accomplished using a cyclic integrated reversible-bending 

fatigue tester (CIRFT) recently developed by ORNL.3,4,5,6,7,8,9,10,11 The integrated U-frame is 

shown in Figure 1. 

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 Uframe setup is oriented in a horizontal plane and is driven by electromagnetic-force–based Bose 

dual linear 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 and a maximum stroke of 

±25.6 mm. 

Bending is imposed through a U-frame with dual driving points and a 101.60 mm loading arm. 

Under a pair of forces or displacements that face outward, the rigid arms are opened, and 

bending moments force the rod to deflect outward (namely, away from operator, as shown in 

Figure 2). Under a pair of forces facing each other, the rigid arms are closed, forcing the rod to 

deflect inward. With the special grips that will be described in Section 2.2, a pure reversible 

bending condition can be obtained. The CIRFT can deliver dynamic loading to a rod specimen 

in the load-control mode at 5 to 10 Hz. The current configuration enables the system to test a 

rod 9.70 to 11.74 mm in diameter, 152.40 mm (6 in.) in length, and 50.80 mm (2 in.) in gage 

section. Three LVDTs measure rod deflections at three adjacent points within the gage section 

to determine rod curvature, which is then correlated to the applied moment to characterize the 
4 

mechanical properties of the bending rod. Online monitoring can capture mechanical property 

changes to reveal fatigue behavior during testing. 

 

(a) 

(b) 

Rigid arms 

Connecting plates 

Universal testing machine links 

Three LVDTs for curvature 

measurement 

Rod 

specimen 
5 

(c) 

Figure 1. (a) Horizontal layout of ORNL U-frame setup; (b) rod specimen under test and 

three LVDTs for curvature measurement (operator is facing the three LVDTs); and 

(c) front view of CIRFT installed in ORNL hot cell. 

Figure 2. Schematic drawings of U-frame setup for reversal bending when rigid arms are 

(a) closing, (b) neutral, and (c) opening. 

SNF rod 

End-blocks 

LVDT clamp 





P 

(a) Rigid arms are closing. The curvature is 

concave outward and designated with a 

negative sign. 

(b) Rigid arms are in neutral position. 

(c) Rigid arms are opening. The curvature 

concave inward and designated with a 

positive sign. 
6 

2.2 GRIP DESIGN 

Each grip consists of a V-shaped block integrated into the rigid arm of the U-frame, and a 

V-shaped end-block (Figure 3). The surfaces of the blocks serve as contacts to the specimen 

assembly and have been equipped with linear roller bearing sets embedded to ensure the free 

axial movement of the specimen when the force is being transferred. The rod specimen is 

secured by opening and closing the end-blocks in two grips.4,5,6,7 A partly exposed view on one 

end of the specimen is shown in Figure 4 with the end-block removed. The specimen has a pair 

of rigid sleeves on both holding areas. The sleeves protect the specimen from any contact 

damage through a compliant layer between the specimen and sleeves. The rigid sleeves also 

enable the specimen to move freely in the axial direction while being set against the embedded 

linear roller bearing sets. The free axial movement of the specimen is a critical requirement in 

pure bending, as discussed elsewhere in the literature.3

 

Figure 3. Assembled grip (left) and two views (right) of detached grip body. 
7 

Figure 4. Image showing the grip design of CIRFT with one end-block removed. 

The mounting of rigid sleeves onto the rod is achieved using casting epoxy. The cast epoxy, 

therefore, serves as the compliant layer whose size depends on the dimensions of the rigid 

sleeve and rod. The current rigid sleeves have an inside diameter (ID) of 15 mm, an outside 

diameter (OD) of 25 mm, and a length of 50 mm. For fuel rods with an OD of 10.70 mm, taken 

from the H. B. Robinson Nuclear Plant (HBR) reactor in South Carolina, the cast epoxy layer 

would have a thickness of 2.15 mm and the same length of rigid sleeves. Related physical and 

mechanical properties of the cast epoxy are listed in Table 1, along with those of the stainless 

steel in the U-frame setup. The use of cast epoxy as a compliant layer has been demonstrated 

to be effective in meeting requirements for both specimen protection and pure bending 

boundary conditions.4,7 

Table 1. Physical and mechanical properties of related materials 

Materials Density (kg/m3) Young’s modulus 

(GPa) Poisson’s ratio 0.2% yield strength (MPa) 

Stainless steel 8,030 193 0.30 290 

Epoxy 1,200 3.5 0.25 45 

Casting of the epoxy has been proven feasible in the hot cell environment. A vise mold has 

been designed and built for that purpose, and the procedures to cast the epoxy and mount the 

rigid sleeves into the rod specimen have been developed.4,5 Several important modifications 

were made to adapt the vise mold to hot cell testing.9

 The images in Figure 5 (a) show different 

views of the vise mold, and those in Figures 5 (b) and 5 (c) illustrate the process of epoxy 

casting in the hot cell. The robust mold design enables two important functionalities for casting 

the CIRFT specimen in a hot cell environment: (1) the rod ends up in the center every time, and 

(2) the compliant layer is applied the same way each time so as not to introduce a variable. 

The amount of fresh epoxy for each rigid sleeve is determined as needed in the hot cell based 

on the results from an out-of-cell study. In the case of the HBR rods, 5 grams (g) of epoxy can 

fill the gap properly. The rod released from the vise mold is generally examined visually, and the 

gap will be refilled if a substantial cavity is observed between the rod and the rigid sleeves. 

Rod specimen 

Linear roller bearing sets (blocked 

by specimen) embedded in the 

end-face of the rigid arm, enabling 

free axial movement of specimen 

Three LVDTs 

for curvature 

measurement 

Rigid sleeves 
8 

(a) 

 

 (b) (c) 

Figure 5. (a) Vise mold for assembling octagonal rigid sleeves into a rod; (b) rod (Demo1, 

606B2) to be inserted; (c) one pin was turned away and a rigid sleeve loaded with epoxy 

inserted into the lower chamber. 

Vise mold 

Rod 

Left pin 

Right pin 

Upper 

chamber 

Octagonal insert or rigid sleeve 

Handle 
9 

2.3 LVDT MEASUREMENTS AND CURVATURE CALCULATION 

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

the traveling displacement at the loading points of the rigid arm. The displacement measured, 

however, contains the contribution of the compliant layers, depending on the materials used in 

the compliant layers and the level of loading. 

To address this issue, direct measurement of the specimen displacement at three adjacent 

points along the rod method was proposed12 and has been implemented to evaluate the 

curvature of a bending rod in this study.4,7 

Given the deflections from three LVDTs, d1, d2, and d3, as shown in Figure 6, the curvature  of 

the bending rod can be evaluated as follows: 

2/12

0

2

20 [( ])

   ydx , (1)

)(2

2 )()( 32 21

0

ab

aba b

mm

ddmddmhmm

x 

 

,

22

1 21

0 0

hdd

x

m

y

a

 



 



  

,

where 

12 dd

h

ma  

,

23 dd

h

mb  

,

and h is the sensor distance, 12 mm. 

Figure 6. Determination of the bending curvature of the rod by use of deflections 

measured at three points. 

The arrangement of the three LVDTs and their installation in the setup can be seen in Figure 1 

and Figure 4. A series of static tests was performed to confirm the measurement method by 

using a surrogate rod 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.7

h h 

1 2 3 

x

y 

d1 d2

d3
10 

The effect of contact between the LVDT probe and the rod on the curvature measurement 

became a concern, especially when loading during the hot-cell testing approached very small 

levels. The current LVDTs use flat, disk-shaped probes. The sensor spacing h in the curvature 

calculation would be affected by the contact position of the probes. As shown in the upper view 

of Figure 7 (a), the contacts will be shifted outward and the h increases correspondingly when 

the rod is concave toward the sensors; in the lower view, the h decreases when the rod is 

concave away from the sensors. 

In an out-of-cell study, the effect of the probe on the curvature measurement in dynamic testing 

has been investigated by using LVDT probes with a chisel contact.13 The results showed that 

the effect of probe geometry on the curvature measurement is negligible in dynamic testing 

within a low load range. Thus, the initial data reduction for the hot-cell testing did not envision 

the correction needed to account for the effect of the probe’s width. However, CIRFT testing in 

the hot cell reveals the nonsymmetry curvature measurements for the clad tension and clad 

compressive loading cycles. This is partly because the SNF rod has a very limited pellet-pellet 

interface bond as compared to that of epoxied surrogate rod with well-bonded pellet-pellet 

interface structure. The potential sensor probe contact also has some effect. Thus, the SNF 

pellet-pellet debond interface–induced stress concentration effect cannot be readily observed or 

revealed from a surrogated rod tested under low loading amplitude. In order to further clarify the 

potential sensitivity of curvature measurement with regard to LVDT probe contact, a control 

experiment with a uniform polycarbonate surrogate rod was conducted, and the test results are 

illustrated in Figs. 7 (b) and (c). The test results of LVDTs located at the clad compression site 

show that the measurement with the disk probe overestimates the curvature by 50% compared 

to measurements with the chisel head probe. An adjustment factor of h = 2.9 mm was applied 

to develop consistent results. For LVDTs located at the clad tensile site, the disk probe 

underestimates curvature by 40%, and a correction factor of h = 2.4 mm is needed. 

 

(a) 
11 

(b)

(c)

Figure 7. (a) Diagrams showing the effect of the LVDT probe with disk shape on 

curvature measurement, where the contact points are shifted. Comparisons of curvature 

measurement between disk and chisel head probe for LVDTs located at clad 

compression and tension site are shown in (b) and (c), respectively. (c). 
12 

The adjustment factors presented above are only applied to a uniform rod system. Further 

investigation of using the surrogate rod with the debond pellet-pellet interfaces of a nonuniform 

rod system is needed to take into account the stress concentration effect and the increased 

curvature resulting from segment pellet structure. This research will help to finalize adjustment 

factor development. 

2.4 TEST SEGMENTS 

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 were taken from rods at locations E04, E14, F07, G10 and R05 

in assembly S-15H. The irradiation history of these rods and details can be found elsewhere.1

 

Although the goal of this project was to test static and cyclic bending performance, some rod 

segments were used for instrumentation demonstration, machine tuning, and benchmarking. 

The specimen cutting plan and the specimen locations in the father rods can be found in 

Appendix A. 

A summary of specimen information for hot cell testing is provided in Table 2. 

According to the plan, 6 in. segments were cut. Figure 8 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.14 The measurement 

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

rod and 3 (0, 120, and 240o

) at each reading distance. The mean value and standard deviation 

of the diameter readings were found to be 10.772 mm (0.4241 in.) and 0.008 mm (0.0003 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 8. Segment 608C4B for DL2 taken from rod G10, with the storage tube. 
13 

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

Specimen 

label 

ORNL 

inventory 

label 

(“Seg. ID”) 

Span Burnup* (GWd/MTU) Estimated oxide* (μm) Estimated hydrogen of span** (wppm) Estimated local hydrogen (wppm) Outer diameter (mm) Wall*** thickness (mm) 

Static tests 

S1 606C3C 4 66.8 70–100 550–750 650 10.748 0.751 S2 605D1E 2–3 66.5 40–70 360–550 400 10.746 0.750 S3 609C5 4 66.5 70–100 550–750 550 10.736 0.745 S4 609C6 4 66.5 70–100 550–750 550 10.740 0.747 

Dynamic tests

D0 605D1F 2–3 66.5 40–70 360–550 360 10.746 0.750 D1 607C4B 4 63.8 70–100 550–750 700 10.834 0.794 D2 608C4B 4 63.8 70–100 550–750 700 10.781 0.768 D3 605C10A 4 66.5 70–100 550–750 550 10.774 0.764 D4 605D1C 2–3 66.5 40–70 360–550 500 10.750 0.752 D5 605D1B 2–3 66.5 40–70 360–550 550 10.771 0.763 D6 609C4 4 66.5 70–100 550–750 700 10.748 0.751 D7 609C3 4 66.5 70–100 550–750 750 10.777 0.765 D8 606C3E 4 66.8 70–100 550–750 550 10.739 0.746 D9 609C7 4 66.5 70–100 550–750 550 10.738 0.746 

D10 606C3A 4 66.8 70–100 550–750 750 10.759 0.756 D11 607C4A 4 63.8 70–100 550–750 700 10.832 0.793 D12 608C4A 4 63.8 70–100 550–750 700 10.821 0.788 D13 606B3E 5 66.5 100–110 750–800 750 10.765 0.760 D14 606B3D 5 66.5 100–110 750–800 750 10.765 0.760 D15 606B3C 5 66.5 100–110 750–800 750 10.768 0.761 

* Reference (1)

** References (15,16) *** Cladding wall thickness was based on the nominal inner diameter (9.256 mm, 0.364 in) provided by EPRI (2001) and the outer diameter measured by 

ORNL.
14 

2.5 TEST PROTOCOL 

The test system was calibrated under both static and dynamic testing conditions using three 

surrogate rods consisting of stainless steel cladding with alumina pellets. The system was tuned 

and benchmarked by testing a number of specimens. The related results can be found in 

Appendix C. 

2.5.1 Static Testing 

The purpose of static testing is to generate bending strength data for SNF rods. In particular, the 

static tests measure the bending strength of fueled SNF rods so that the contribution of the fuel 

to the SNF rod bending strength can be evaluated. The analysis of the static test results also 

provides a reference yield strength to establish the dynamic testing matrix. Static testing is 

carried out using displacement control. It involves ramping both loading arms of the U-frame at 

0.1 mm/s up to 12 mm, where 12 mm displacement is the machine stroke capacity. The 

following procedure is followed. 

 Perform the standard static bending test beyond SNF rod yielding to failure or up to device 

displacement or loading capacity. 

 If the machine capacity is reached before specimen failure or specimen maximum strength 

is reached, repeat unidirectional static testing using the same condition for a few more 

loading/unloading cycles, or to specimen failure, whichever comes first. 

 If the SNF rod does not fail after 3–4 loading/unloading cycles, follow a dynamic or cyclic 

test procedure to fracture the tested specimen to support postmortem examination. Such a 

dynamic test is called a follow-up test. 

 Collect and weigh any fuel fragments that may have dislodged during the test. 

The loading processes to be used after the first loading cycle were suggested because the 

device might not have sufficient stroke to test these particular specimens to failure. 

2.5.2 Dynamic Testing 

Dynamic testing consists of two major activities—dynamic real-time online monitoring and 

periodic quasi-static deformation measurements (see Figure 9). The procedure is as follows: 

 Perform the dynamic cyclic test under constant load control using a sine wave input in 

reverse bending mode. 

 Set the cycle frequency at 5 Hz and select the amplitudes for individual cycle tests 

considering the target cycles to be achieved with each test. 

 Monitor the SNF fatigue evolution with defined intervals: perform static measurements of the 

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

amplitude under displacement control. 

 Stop the dynamic test when failure/clad fracture is detected or the preselected number of 

cycles is reached. 
15 

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

Figure 9. Flowchart for cyclic testing of spent fuel rod. 

2.6 DATA PROCESSING 

Measurement data and 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, such as 

those illustrated in Figure 10 for D1. The uniform moment applied to the rod is derived simply 

from the following equation. 

M = F × L (2) 

where F is the averaged value of applied loads (load1 and load2) from the Bose dual motors, 

and L is the loading arm length, 101.60 mm. The computation of curvature is described in 

Section 2.3. 

An equivalent stress-strain curve can be obtained under the assumption that the SNF rod can 

be idealized as a linear elastic homogeneous material without consideration of the effects 

induced by pellet-cladding interaction. The equivalent stress was calculated using: 

 = M × ymax/I (3) 

where I is the moment of inertia, I = Ic+Ip, Ic and Ip are moments of inertia of cladding and pellet, 

respectively, and ymax is the maximum distance to the neutral axis of the test rod section and is 

set equal to the radius of the cladding. The measured ODs of rod segments are given in Table 

2. The calculation of stress disregards the difference of elastic moduli between cladding and 

pellets. 

The equivalent strain is then: 

ɛ =  × yman . (4) 
16 

2.6.1 HBU SNF Characteristics 

The SNF system contains multiphase components such as the cladding, pellets, and oxide. 

Thus, an SNF rod is far from an ideal homogeneous material, especially in HBU fuel, where the 

irradiated cladding and pellets and the associated interfaces are modified significantly during 

irradiation. Assuming the HBU rod has a perfect interface bond at pellet-cladding and pelletpellet interfaces, the associated flexural rigidity EI can be expressed as 

EI = EcIc + EpIp , (5) 

where Ic and Ip are moments of inertia of the cladding and pellet, respectively, and the value of Ip

is based on the reference.1 Ec and Ep are the Young’s modulus of the cladding and pellet. The 

associated stress and strain evaluations of Eqs. (3) and (4) can still be used. 

However, because of the imperfect interface cohesive bonding, the effective EI will be less than 

that of the perfect bonding property, which can be written as 

EI = EcIc + EpIp  F(BE, cladding, fuel), (6) 

where F is the correction factor, which depends on interface bonding efficiency (BE) and the fuel 

pellet and cladding aging properties, as well as cyclic loading amplitudes. This hypothesis was 

further validated in Reference 17.

Figure 10. 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. 
17 

Based on the static measurements of CIRFT testing, the EI data trend indicates that the 

cohesive bonding at the pellet-pellet interface is fairly weak. The resulting degradation of 

flexural rigidity, which is associated with the debonding at the pellet-pellet interface, is reflected 

by the LVDT measured curvatures evidenced by a significant variation between the cladding 

tensile stress and cladding compressive stress regions shown in Fig. 10. This suggests that the 

fuel pellet reinforcement for SNF rod flexural deformation at the cladding tensile stress region is 

no longer available due to pellet-pellet debonding under flexural tensile loading. 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. 

The above mentioned phenomenon was also observed from the CIRFT test online monitoring 

data, where the tensile cladding stress stage shows a higher curvature reading than the 

compressive stress stage, as shown in Fig. 24 (d). The EI values also change accordingly in 

each moment reversal cycle, as shown in moment-curvature plot of Fig. 10, where the EI (the 

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

cycle. 

Furthermore, a detailed three-dimensional finite element analysis with a 6.25 N·m uniform 

moment also reveals that the localized stress in the cladding at the pellet-pellet-cladding 

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

parameters)17,20 the LVDT measurement for the tensile cladding stress region, and the 

associated localized stress is about 2.6 times the average tensile cladding stress away from the 

pellet-pellet interface region. 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.17 Therefore, the conventional 

approach, as stated in Eq. (3) and (4) from a global M-k consideration, is no longer valid for 

describing the cladding failure mechanism associated with localized flexural rigidity degradation 

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

pellet-pellet interface translated from global measurements is beyond the current scope of the 

project. Therefore, only the mechanical response derived from direct measurements (the 

LVDTs) will be utilized in this study and discussion. 

In order to generate the stress-strain relationship associated with M-k, a first order 

approximation approach, commonly used for an isotropic/homogeneous structure, was adopted 

to translate the global CIRFT M- data into equivalent stress and strain data of cladding. Where 

the SNF rod’s moment of inertia is written as I = Ic+c  Ip, Ic and Ip are moments of inertia of 

cladding and pellet, respectively, and c = Ep/Ec (because the baseline Young’s modulus of the 

oxide fuel is twice that of the cladding, c is likely greater than 1). In the proposed equivalent 

stress and strain approach, c is set to 1, which implies a 50% reduction in pellet Young’s 

modulus, to take into account the radiation induced degradation of the HBU fuel pellet, where 

the pellets are assumed to be perfectly bonded with the cladding at the pellet-pellet interface. 

Furthermore, setting c = 1 implies the same E properties for the fuel and the cladding, which 

allows the use of Eq. (3) and (4) under the homogenous properties hypothesis for an equivalent 

stress and strain evaluation. 

Generally, the moment range (M), curvature range (), and flexural rigidity EI are used in 

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

M = Mmax  Mmin, (7)

  = max  min, 
18 

EI = M / 

where the subscripts “max” and “min” represent the maximum and minimum waveforms (Figure 

10). 

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

one cycle of loading as will be discussed in Section 4.1. The offset of the M- loop on the  axis 

with respect to the origin can be described by a mean value of curvatures, m: 

 m = 0.5  (max + min); (8) 

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

 ),max(  max  max  min . (9) 

For a given specimen, the ||max given by Eq. (9) 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 11. The CIRFT system resistance was tested and evaluated by using empty runs in 

which the specimen was not loaded on testing machine. This is conducted in the specified 

period or whenever the calibration is needed. 

The effect of CIRFT system resistance on the dynamic loading of the rod was also investigated 

by using a calibration rod in an out-of-cell study. It was shown that due to low loading amplitude 

the net effect of system resistance and dynamic inertia on the load applied to the rod is 

negligible within the tested range. Since the out of cell condition that was evaluated is 

equivalent to the condition used in hot cell cyclic testing, no correction procedure needs to be 

applied to the data acquired during dynamic tests. 
19 

Figure 11. “Empty-run” conducted on HBR S1, for which the load and displacement were 

plotted. 

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

21 

3. STATIC TESTING 

3.1 RESULTS 

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

Four static tests were subsequently conducted. This chapter focuses on these four static tests; 

refer to Appendix C for the static benchmarking tests. 

S1 (606C3C, 66.8 GWd/MTU burnup, 70–100 m 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 12 (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 3.2 m-1. The slope changes of the 

linear portion of moment-curvature curves could be seen approximately near moment levels 12 

and 59 N·m, respectively. It is noted here that due to pellet end dish-in design, the pellet-pellet 

interface bonding strength would be fairly weak, and the pellet-pellet de-bond phenomenon 

should be observed in the early phase of bending loading. The slope changed points indicate 

that the flexural rigidity changed in the SNF rod system within the linear elastic range. The 

changes of slope and associated deflection points are quantified in Section 3.2. The equivalent 

stress-strain curve based on Eqs. (3) and (4) for the initial cycle is presented in Figure 12 (b), 

where Ymax and I are based on the measured clad OD provided in Appendix B.. 

Since the machine capacity was reached and no failure occurred to the specimen, a follow-up 

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

 moment reversals. The 

failure occurred near motor 1 (right side) with a fuel release of 0.9 grams. Optical images of the 

fracture surfaces are shown in Figure 13. The pellet end-face can be identified from the haze 

zone characterized by a brown-blue color (indicated by the arrow in Figure 13a). This indicates 

that the rod failure may have initiated at the pellet-to-pellet interface. At the same time, a 

serrated shear fracture surface of the cladding was observed. The serrated fracture surface was 

believed to be the end stage of the crack propagation upon the fatigued CIRFT sample 

fractured. Note that this specimen appears to have been loaded into the test machine at 90° 

from the specimen ID marks. 

The cladding surface (Figure 13 [c] and [d]) near the neutral axis of the bending rod is free of 

visible surface oxide spalling, which is expected since these regions experienced little stress. In 

contrast, one of the stressed lateral surfaces which experienced the maximum flexural stress is 

covered with significant visible circumferential surface oxide spalling. 
22 

(a) (b) 

Figure 12. (a) Curve of moment versus curvature, and (b) equivalent strain–stress curve1

for S1 (606C3C). 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. 

 

(a) (b) 

(c)

 

1

 Based on Eq. (3) and Eq. (4). 
23 

(d) 

Figure 13. Mating fracture surfaces (a) and (b); lateral sides near the neutral axis of the 

bending rod for S1 (606C3C) (c) and (d). The rod was subjected to four repeat static 

cycles to 24 mm maximum relative displacement followed by dynamic cyclic loading 

±25.40 to 30.48 N·m at 5 Hz; 1.4×104

 failure cycles, 0.9 gram fuel particles collected. Note: 

This specimen appears to have been loaded into the test machine at 90° off from the ID 

marks 

Specimens S2 and S3 were both tested under static condition for three loading/unloading cycles 

without any sign of failure. They failed in follow-up dynamic tests under ±30.48 N·m at 5Hz. The 

amplitude of applied moment in the dynamic tests was 35 to 36% of the maximum applied 

moments used in the respective unidirectional bending process, located in the early second 

stage of moment-curvature curves. It is noted here that the selection of the applied moment 

amplitudes was to generate the failure of the tested specimens within a reasonable number of 

cycles. The bending fatigue response of the as-received fuel rods will be discussed next. 

Finally, S4 was tested under static conditions for three cycles and failed during the fourth cycle. 

Details of these static tests can be found in Appendix D. 

It should be noted that, in the static tests reported above, loading/unloading cycles did not result 

in apparent or catastrophic failure of specimens because of the limited stroke of the CIRFT 

machine. The failure of specimens occurred in the initial loading cycle of two benchmarking 

tests, Scal1 and Scal2. The peak moments of the two tests were 78.2 and 80.7 N·m, 

respectively. Details of these benchmarking tests can be found in Appendix C. 

A summary of the moment-curvature curves is shown in Figure 14 for the four static tests. The 

top and right axes present the scale of the converted maximum strain and stress of the bending 

specimen according to Eqs. (3) and (4). 

Several important observations can be made with respect to the responses of the test 

specimens during the first loading cycle. 

 The static test results of the four specimens prior to the follow-up dynamic testing all show 

relatively similar elastic behavior. 

 The initial moment-curvature response is characterized by a flexural rigidity of 52 to 63 

N·m2

. 

 There is a change in the flexural rigidity at the moment of 12 to 21 N·m. This change will be 

quantified in Section 3.2. 
24 

 After this change, the moment-curvature response is characterized by a degraded flexural 

rigidity of 37 to 41 N·m2

. 

 Each static test segment experienced 1–2% plastic strain without failure. 

 The unloading flexural rigidity is similar to the loading flexural rigidity in the second stage 

after the moment reaches 12 to 21 N·m. 

Overall, the rods failed within the gage sections in all of the tested rods. The majority of the rod 

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

location seemed away from pellet-pellet interface, as no end face of the pellet can be identified 

from the fracture surface of the fuel. A summary of static test results is provided in Table 3. 

Figure 14. Moment-curvature curves measurements in static tests (Corresponding stress 

and strain displayed on right and top axes, respectively). 
25 

Table 3. Summary of static tests

Specimen 

ID 

Specimen 

burnup 

Testing condition and result Fracture surfaces Side view of fractured specimens 

S1, 

606C3C 

66.8 GWd/MTU Four static cycles to 24 mm 

max. relative displacement, 85.5 N·m max. moment, followed by dynamic loading ±25.40 to 30.48 N·m, 5 Hz; 1.4×104 failure 

cycles, 0.9 g fuel fragments collected 

S2, 

605D1E 

66.5 GWd/MTU Three static cycles to 24 

mm max. relative displacement, 87.2 N·m max. moment, followed by dynamic cyclic loading ±30.48 N·m, 5Hz; 7.2×103

failure cycles, 0.6 g fuel particles collected 

S3, 609C5 66.5 GWd/MTU Three static cycles to 24 

mm max. relative displacement, 85.8 N·m max. moment, followed by dynamic loading: ±30.48 N·m, 5Hz; 9.6 × 103 failure 

cycles, 0.2g fuel particles collected 
26 

Table 3. Summary of static tests (continued)

Specimen 

label/ ID 

Specimen 

condition 

Testing condition and result Fracture surfaces Lateral surfaces of fratured specimens 

S4, 609C6 66.5 GWd/MTU Three repeat static loadings 

to 24 mm max. relative displacement and failed at 85.5 N·m in the 4th cycle, 1.3g fuel particles collected 
27 

3.2 DISCUSSION 

3.2.1 Characteristics of Moment-Curvature Curve 

The moment-curvature responses of all four static test specimens were similar. They are 

characterized by two linear constituent behavior responses, followed by a nonlinear response 

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

the rod followed the unloading curve linearly and proceeded with nonlinear response after 

passing the previously achieved maximum load. Such loading/unloading/reloading responses 

can be found in many mechanical systems where strain hardening prevails.7

An effort was made to characterize the moment-curvature response based on the characteristic 

points to facilitate understanding test results. The flexural rigidities EI1, EI2, and EI3 were 

obtained, corresponding to the slopes of the first and second linear segments and of the 

unloading segment by using curve fitting with the first order polynomial (Figure 15). 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 and horizontal axis intercept 0.37 m1

. The quantities corresponding to points A, B, 

and C are designated by A, B, C, and MA, MB, and MC. 

The results are summarized in Table 4. For a given specimen, EI1, EI2, and EI3 are generally in 

decreasing order with a marginal difference between the latter two. The characteristics derived 

from equivalent stress-strain curves are provided in Table 5. The E1 in the initial stage of the 

stress-strain curve was 83 to 101 GPa, and the 0.2% yield strength (C) was 687 to 727 MPa. 

The ranges of the elastic modulus and the 0.2% yield strength appear consistent with the range 

of HBU HBR cladding.18 However, the observation should not be overemphasized, for the 

results observed here reflect a comprehensive global response of fuel rods with both pellets and 

cladding included. Moreover, substantial simplifications were made with regard to the elasticity 

and contributions of individual components, where the localized stress riser in cladding at the 

pellet-pellet interface region was not considered. 
28 

Figure 15. Characteristic points of moment-curvature curve. 
29 

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

Spec 

label Seg. ID 

EI1 EI2 EI3 A B C D 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·m 

S1 606C3C 63.057 40.122 35.016 0.267 1.435 2.659 3.242 12.293 59.142 80.105 85.503 S2 605D1E 52.032 41.462 36.067 0.385 1.419 2.719 3.084 19.651 62.496 84.696 87.155 S3 609C5 51.627 41.457 32.618 0.378 1.373 2.864 3.413 17.684 58.946 81.322 85.843 S4 609C6 52.286 36.736 28.657 0.347 1.146 3.288 3.532 21.306 50.656 83.602 85.515 Avg. 54.751 39.944 33.090 0.344 1.343 2.883 3.318 17.734 57.810 82.431 86.004 Std. Dev. 5.544 2.230 3.289 0.054 0.134 0.284 0.196 3.918 5.040 2.093 0.783 

Table 5. Characteristic points and quantities based on equivalent stress-strain curves 

Spec 

label Seg. ID E1 E2 E3 ɛA ɛB ɛC ɛD A B C D GPa GPa GPa % % % % MPa MPa MPa MPa 

S1 606C3C 100.577 63.995 55.851 0.144 0.771 1.429 1.742 105.388 507.039 686.75 733.032S2 605D1E 83.121 66.235 57.616 0.207 0.762 1.461 1.657 168.669 536.428 726.971 748.081S3 609C5 82.796 66.486 52.31 0.203 0.737 1.537 1.832 152.237 507.458 700.088 739.008S4 609C6 83.721 58.822 45.886 0.187 0.616 1.766 1.896 183.201 435.575 718.86 735.312Avg. 87.554 63.885 52.916 0.185 0.722 1.548 1.782 152.374 496.625 708.167 738.858Std. Dev. 8.691 3.556 5.180 0.029 0.072 0.152 0.104 33.781 42.962 18.183 6.623 
30 

3.2.2 Comparison of Static Results with PNNL Cladding Data 

In order 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 PNNL19. 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 

Using these values in the PNNL database, the Young’s modulus was given as 9.15x1010 Pa, the 

yield stress 919 MPa, the uniform elongation 0.00393, and ultimate tensile strength 976 MPa. 

From these values the engineering stress and strain values can be calculated. 

It is noted that the CIRFT LVDTs measured the global rod deformation within the gage section. 

Thus, the stress-strain estimate is an average evaluation throughout the gage section that 

covers several pellet lengths with several pellet-pellet interfaces. For the linear response of the 

cladding region of the rod away from the pellet-pellet interface, evaluation of the stress and 

strain can be a straightforward composite flexural rigidity formulation due to relatively good bond 

at the pellet-cladding interface. However, due to the composite structure segmented pellets, the 

localized stress riser or curvature magnification at the pellet-pellet interface region cannot be 

estimated directly from the equivalent stress and strain approach; instead, they need to be 

evaluated using FEA17, 19. 

In order to further quantify the fuel pellet support or reinforcement to the SNF system 

mechanical property with respect to bending flexural response, the comparison based on CIRFT 

M- data was completed as illustrated below. 

For comparison, the PNNL data ( curve) were converted to a moment-curvature curve by 

using the following equations, with the consideration of stress distribution that is appropriate for 

a thin-walled tube under uniform bending, 

 M = ·I/ymax, (10) 

and 

 = /ymax, (11) 

where I = Ic, and other quantities have same meanings as in Eq. (2) and (3). 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 16, where the SNF rod shows 

much higher bending moment resistance 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 

Error! Reference source not found.. HBR’s rod used in the flexural rigidity calculation has the 

following cladding dimension: cladding ID 0.364”, OD 0.423”. 

In the data comparison above, a distinct difference from that of the four HBU fuel rod systems 

was observed in the slope of the PNNL data . The slopes of the four HBU fuel rod systems are 

greater than that of the PNNL data. By focusing on the initial slope, between 0-15 N*m, the 

slope (flexural rigidity) of HBU fuel rods was approximately twice that of the PNNL data for 

cladding alone. The increase in flexural rigidity can be attributed to the presence of fuel. 

Nevertheless, a detailed analysis is required to determine exactly how the presence of fuel 
31 

accounts for this increase in flexural rigidity, as well as the net effect of fuel support on SNF rod 

system fatigue lifetime. 

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

fuel’s geometry property of the moment inertia. However, the measured flexural rigidity of a fuel 

rod system is much less than that estimated from a direct summation of EcIc + EpIp . This is 

primarily due to interface bonding de-efficiency at fuel-cladding and fuel pellet-pellet interfaces, 

as well as existing fuel-clad gap potential. All these will affect the degree to which the presence 

of fuel results in increased rod flexural rigidity. This research effort did not attempt to account for 

and quantify each of these influences. Thus, the preliminary conclusion based on CIRFT SNF 

static bending testing is that the flexural rigidity was approximately twice what would be 

obtained if cladding properties alone were used to predict behavior. 

From a flexural rigidity perspective, the fuel pellets indeed increase the SNF rod’s system 

flexural rigidity; however, due to segmental pellet structure, numerous stress concentration sites 

within cladding at pellet-pellet interfaces in a SNF rod are created. In general the stress 

concentration event will cause the structure accelerated aging or reduced lifetime as compared 

to the very same system w/o stress concentration sites or discontinuous materials interfaces. 

The stress concentration 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-pellet interfaces. 

Therefore, the pellet and clad mechanical interaction at material discontinuity interface regions 

seems to reduce the benefit gained from the stiffness increase of a HBU fuel rod system. 

Moreover, the intensity of the stress concentration or the pellet-clad interaction is strongly 

dependent on the loading intensity. Thus, at low loading level, cladding alone is still the key that 

dictates the SNF rod flexural reliability. From CIRFT dynamic testing, it also confirms that as 

long as the clad fracture threshold was not reached, a SNF system can survive millions of 

vibration cycles. 

Figure 16. Comparison of CIRFT global data with PNNL moment-curvature curve 

converted from PNNL cladding stress-strain data. 
32 

Table 6. Comparison of flexural rigidity results in CIRFT tests and PNNL data 

EI1 (NM2) EI2 (NM2) EI3 (NM2) 

CIRFT testing 54.751 39.944 33.090 

PNNL data 26.933 

3.2.3 Effect of LVDT Probe Contact on Large Curvature Measurement 

The LVDT probe contact design with respect to monitoring specimen curvature deformation was 

found to affect curvature measurement results. 

An experimental study was performed using two surrogate rods made of identical materials and 

configurations (stainless steel [SS] tubes and ten alumina pellets in each surrogate rod). The 

same flat disk probe contact was used to test each specimen. The flat disk probe forms a linecontact when the probe initially contacts the rod surface. However, upon loading, the contact will 

shift from the line-contact to point-contact at one side of the probe disk. As shown in Figure 17, 

under similar loading levels, the maximum curvature obtained from the LVDT probes located at 

a cladding compression site is significantly larger than that of probes located at cladding tension 

sites. The discrepancy due to the probe contact effect can be mitigated or reconciled by 

adjusting sensor spacing, h, of Eq. (1) as illustrated in Figure 18. It has been shown that under 

progressive adjustments the moment-curvature curves in both cases seem to converge. The 

actual sensor spacing adjustment can be calibrated with a chisel head probe that has point 

contact instead of line contract to the rod, as shown in Section 2.3. 

Due to LVDT probes located at cladding compression sites during the static testing, the 

estimated curvature is expected to be overestimated as demonstrated in Figure 18 of dh=0 mm 

case. The resultant higher curvature trend seems to be consistent with the observations in 

Figure 17, where modified equivalence data corresponds to a higher peak tensile strain 

compared to that of the PNNL data. 

(a) (b) 

Figure 17. Moment-curvature curves based on stainless steel alumina pellet (SSAP) rod 

testing when rod is bending with respect to three LVDTs (a) concavely and (b) convexly. 
33 

Figure 18. Moment-curvature curves as a function of sensor spacing adjustment when 

rod is bending with respect to three LVDTs (a) concavely and (b) convexly. The unit of dh 

is mm. 

(a) (b) 

35 

4. DYNAMIC TESTING 

4.1 RESULTS 

Sixteen dynamic tests were conducted on the HBU HBR fuel in the hot cell, including one for 

benchmark testing and fifteen according to dynamic testing plan. Load 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 m 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 gram of fuel 

fragments was collected at the failure site. 

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

conducted using two relative displacement levels: 0.8 and 1.2 mm, at the target intervals 

described in Section 2.5.2. Time series of moment and curvature and moment-curvature loops 

obtained at the 1st cycle and after approximately 111,000 cycles are shown in Figure 19. The 

loops became slender with accumulated cycles, and the amplitude of the loops decreased. The 

relations of moment-range versus curvature-range and flexural rigidity are illustrated in Figure 

20. Most of the rigidity degradation occurred in the first 1,000 cycles. Variations of these 

quantities as a function of number of cycles are provided in Figure 21. Although the curvature 

range stayed quite consistent under displacement control, a decrease in moment range was 

observed. The rigidity of the measurements at two displacements, converged before the failure 

while exhibiting a slightly declining trend. 

The curvature, moment, and flexural rigidity based on online monitoring data are presented in

Figure 22. The online monitoring showed a flexural rigidity of about 50 N·m2

, a little lower than 

that observed in measurements. This occurred because different loading conditions were used 

in measurement and cycling. A curvature range of less than 0.3 m-1 was used in the quasi-static 

measurement, which is lower than that used in the cyclic test to ensure measurement data 

didn’t affect the dynamic cycle data. In general, the flexural rigidity tends to increase with 

decreasing curvature. This is probably due to a better interface bond at relatively low load 

resulting in less stiffness reduction. Overall, a stable rod response was exhibited before the final 

failure. The curvature time history shown in Figure 24 (d), it clearly indicated a non-symmetric 

deformation under reverse loading, where the cladding tension site has much higher 

deformation (about 1.6 times) compared to that at cladding compression site. This phenomenon 

could be the consequence of debonding at the pellet-pellet interface and the stress 

concentration occurring at the pellet-pellet interface region at the tension side of the cladding, in 

addition to sensor probe sensitivity. In general flexural hysteresis history remains quite uniform 

throughout the reversal bending test, as shown in Fig. 24 (f). Under relatively low loading 

amplitude 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 non-linear response 

associated with segment pellets induced stress concentration at interface regions as well as 

uncertainly resulting from sensor probe sensitivity. 

Moment and curvature time history and moment-curvature loops based on online monitoring at 

26 and 1.10 × 105

 cycles are shown in Figure 23. The curvature data appear to have a much 

higher noise level compared to that of SSAP out-of-cell test data, as shown in Figure 10. This is 
36 

primarily due to there being two different calibration ranges for the LVDT set-ups used in the 

out-of-cell testing. The large range set-up is designated for static testing due to the large 

deformation required in that type of test, and the small range set-up which has higher sensitivity 

is designed for dynamic testing due to its lower test specimen deformation level. In a hot-cell 

environment, the current CIRFT device is only equipped with the large range calibration LVDT 

set-up, which is used to cover both static and dynamic testing capabilities. Thus, the sensitivity 

of the small range set-up for the low load dynamic testing was sacrificed, resulting in the higher 

noise level in the curvature data of Figure 23. It is also noted that the symmetry curvature 

responses were observed at the beginning of cycles, which may indicate good interface bonds 

at the beginning of test cycle, while at higher cycling, the nonsymmetry characteristic of 

curvatures under clad tension and compression cycles was observed in Fig. 25. 

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

setup). The failure occurred at the pellet-to-pellet interface as illustrated in Figure 24. The end 

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

surfaces of the rod were found to have been covered with equally spaced circumferential cracks 

throughout the gage section. Spalling only occurred on the local area near the fracture. The 

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

from the damage experienced 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 (for example, see S3 [609C5] in Table 3). 
37 

 

 

(a) (b) 

Figure 19. 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. 
38 

 

(a) (b) 

Figure 20. (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. 

·

(a) (b) 

(c) 

Figure 21. 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. 
39 

 

(a) (b) 

 

(c) (d) 

(e) (f) 

Figure 22. 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 = 1.1×105

 cycles under ±15.24 N·m, 5 Hz. Fuel particles collected <1.0 g. 
40 

 

 

 

(a) (b) 

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

(a) 26 and (b) 1.10×105

 cycles; results are based on online monitoring. 
41 

 

(a) (b) 

(c)

(d) 

Figure 24. (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. 

The details of the other dynamic tests can be seen in Appendix E. The dynamic testing results 

for HBR rods are summarized in Table 7, along with one rod from benchmarking (D0). The 

result for specimen D0 is considered to be effective in characterizing the cyclic fatigue of rods, 

as the test followed the dynamic testing procedure. Table 7 shows the failure modes of rods 

from the dynamic tests. 
42 

Table 7. Summary of dynamic test results 

Spec label Seg. ID 

Applied Load 

Applied Moment amp. 

Number of cycles 

Failure** /2 /2 ||max m min max

N N·m % m1 m1 m1 m1 m1

D0 605D1F 250 25.4 2.50×104 1 0.3073 0.5712 0.6169 0.0457 0.5254 0.6169 D1 607C4B 150 15.24 1.10×105 1 0.1707 0.3173 0.3844 0.0689 0.3844 0.2465 D2 608C4B 50 5.08 6.40×106 0 0.0706 0.1313 0.1629 0.0288 0.1054 0.1629 D3 605C10A 100 10.16 1.00×106 1 0.1202 0.2235 0.3305 0.0968 0.3305 0.1369 D4 605D1C 75 7.62 1.10×107 0 0.0964 0.1791 0.2312 0.0582 0.1149 0.2312 D5 605D1B 90 9.144 2.30×106 1 0.1100 0.2045 0.2260 0.0216 0.2260 0.1829 D6 609C4 125 12.7 2.50×105 1 0.1625 0.3021 0.3243 0.0222 0.2799 0.3243 D7 609C3 200 20.32 6.50×104 1 0.2607 0.4847 0.5023 0.0176 0.5023 0.4671 D8 606C3E 87.5 8.89 1.28×107 0 0.1065 0.1979 0.2125 0.0184 0.1756 0.2125 D9 609C7 350 35.56 7.10×103 1 0.4882 0.9074 1.1903 0.2829 0.6245 1.1903 D10 606C3A 125 12.7 1.80×105 1 0.1628 0.3026 0.4015 0.0990 0.2036 0.4015 D11 607C4A 300 30.48 5.50×103 1 0.3528 0.6558 0.8533 0.1975 0.4582 0.8533 D12 608C4A 110 11.176 3.86×105 1 0.1311 0.2436 0.2845 0.0409 -0.2845 0.2027 D13 606B3E 135 13.716 1.29×105 1 0.1458 0.2710 0.3873 0.1164 0.1546 0.3873 D14 606B3D 87.5 8.89 2.74×105 1 0.1130 0.2100 0.2848 0.0748 0.2848 0.1351 D15 606B3C 75 7.62 5.11×106 0 0.0926 0.1721 0.2445 0.0724 0.0997 0.2445 

** 1-failure, 0-without failure; 
43 

Table 8. Failure modes of rods from dynamic tests 

Spec label 

Applied 

Moment 

amplitude 

 Number of 

cycles to failure Cross section Lateral view 

N·m m1

D1, 607C4B 

15.24 0.6346 1.10×105

D2, 608C4B 

5.08 0.2626 6.40×106 w/o 

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

to failure 
44 

Table 8. Failure modes of rods from dynamic tests (continued) 

Spec label 

Moment 

amplitude  Number of cycles to failure Cross section Lateral view N·m m1

D3, 605C10A 

10.16 0.4469 1.00×106

D5, 605D1B 

9.144 0.4089 2.30×106

D6, 609C4 12.7 0.6042 2.50×105
45 

Table 8. Failure modes of rods from dynamic tests (continued) 

Spec label 

Moment 

amplitude  Number of cycles to failure Cross section Lateral view N·m m1

D7, 609C3 20.32 0.9693 6.50×104D16, 605D1F 

25.4 1.1423 2.50×104
46

The equivalent strain amplitudes (a, half of strain range  /2, and equivalent strain  is based 

on Eq. [4]) as a function of the number of failure cycles are shown in Figure 25. 

 For those specimens that failed during testing, the fatigue life decreases with increasing 

strain amplitude, resulting in a defined -N curve. The data points can be effectively fit by 

using the power function y = 3.5693·x0.252 with a correlation coefficient as high as 0.8722, 

where x is the number of cycles to failure and y is the strain amplitude (%). 

 An endurance limit is likely located between 0.106 and 0.110% strains if it is defined at 107

cycles. In other words, below 0.1% strain amplitude there appears to be no number of 

cycles that will cause failure. 

 A large variation of hydrogen content existed in the cladding of the test specimens 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. 

The curve fitting to the strain amplitude versus number of cycles was extended to include the 

data points without failure by using two power functions, as presented in Figure 26. For N ≤106, 

the power function was that based on failure data points, y = 3.5693•x0.252; for N >106

, y = 

0.3234•x0.076. It can be seen that the correlation efficient was quite lower for the second power 

function. This is because the amplitude is close to the endurance limit, and the number of failure 

cycles is likely independent of the applied load. 

Figure 25. Equivalent strain amplitudes (/2) as a function of number of cycles. Solid 

markers represent tests with specimens failure; open markers indicate the tests stopped 

without failure. 
47 

Figure 26. Equivalent strain amplitudes (/2) as a function of number of cycles with 

curve fitting extended to include the data points without failure. Solid markers represent 

tests with specimens failure; open markers indicate the tests stopped without failure. 

4.2 DISCUSSION 

4.2.1 -N Curve 

The equivalent strain amplitude has been 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-pellet interfaces cannot be readily represented by this simplified approach. In order 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, it is expected that HBU fuel has a 

fairly weak interface bond at the pellet-pellet interface. This phenomenon was also observed in 

CIRFT test results, where the flexural rigidity dropped off quickly in the initial cycles of testing 

and where a non-symmetric curvature response was revealed in the reverse loading cycle. 

Asymmetrical deformation response is an important signature of the HBU SNF due to pelletpellet-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 27 to serve as another important index for the HBU SNF fatigue life. 

The observations based on curvature amplitude (a, half of curvature range  /2) are similar to 

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

in Eq. (4). 

 The fatigue life decreases with increasing curvature amplitude with a defined -N curve. The 

curve fitting generated a power function y = 6.6345·x0.252 with a correlation coefficient 

0.8722, where x is the number of cycles to failure and y is the curvature amplitude (m-1). 
48 

 An endurance limit may be located between 0.198 and 0.204 m1 if it is defined at 107

cycles. 

It is worth noting that the dynamic tests were conducted under load control, and the curvature 

response of the rods was not necessarily symmetric (Figure 22 [d]). The mean values of the 

monitored mean curvatures m were generally not at the zero level as seen in Table 7. 

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

Section 2.6) was pursued as an alternative quantity for describing the loading imposed on fuel. 

This is significant 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 

27 (a). The -N curve has been demonstrated in terms of maxima of cladding curvature ||max to 

be similar to that of the curvature amplitude a; the exponents of the curve-fit power function are 

in fact the same. 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 m. 

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

 For those specimens that failed during the tests, the fatigue life decreases with increasing 

curvature maxima according to a well-defined -N curve. The data points can be fit by using 

the power function y = 8.1941·x0.252 with a correlation coefficient as high as 0.8926, where x

is the number of cycles to failure and y is the curvature maxima (m-1). 

 A fatigue limit may be located between 0.226 and 0.245 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 vs. the number of failure cycles as shown in Figure 27 (b). 

The curve fitting to the fatigue lifetime as a function of curvature amplitudes was extended to 

include the data points without failure by using two power functions, and the result is presented 

in Figure 28(a). For N ≤106

, the power function was that based on failure data points, 

y = 6.6345·x0.252; for N >106

, y = 0.6011·x0.076. The curve-fitting to the maxima of absolute 

curvatures resulted in the following power functions shown in Fig. 30(b). For N ≤ 8×105

, 

y = 8.1941·x0.252; for N > 8×105

, y = 0.9762·x0.093. The curve-fitting to the maxima of absolute 

strains resulted in the following power functions, shown in Fig. 30(c). For N ≤ 8×105

, 

y = 4.4084·x0.252; for N > 8×105

, y = 0.5252·x0.093. 
49 

(a) 

(b) 

Figure 27. (a) Maxima of absolute curvature extremes and curvature amplitudes as a 

function of number of cycles, (b) Maxima of absolute strain extremes and strain 

amplitudes as a function of number of cycles. (Solid markers represent tests with 

specimens failure; open markers indicate tests without failure.) 
50 

(a)

(b) 

Figure 28. (a) Curvature amplitudes, (b) maxima of absolute curvature extremes, and 

(c) maxima of absolute strain extreme as a function of number of cycles with curve-fitting 

extended to include the no-failure data points. Solid markers represent tests with 

specimens failure; open markers indicate tests without failure. 
51 

(c) 

4.2.2 Post-Irradiation Examination (PIE) 

The objective of the PIE is to gather meaningful information to identify underlying failure 

mechanisms. The primary focuses are: 

 To generate information on cladding features such as hydrides and incipient cracks near 

PPIs and near the mid-pellet. 

 To reveal information about the fuel/cladding bond. 

 To perform PIE with the fuel pellet intact to ensure that the SNF damage mechanism 

associated with P-C interaction and fuel failure can be preserved at the PPI region. 

PIE was conducted on five fuel segments that were 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 PPI – S2/ 605D1E 

 Tested but non-failed under dynamic loading condition– D4/ 605D1C 

 Tested and failed under dynamic loading at a PPI – D3/ 605C10A 

y = 4.4084x‐0.252

R2 = 0.8926

y = 0.5252x‐0.093

R2 = 0.2246

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

1.00E+03 1.00E+04 1.00E+05 1.00E+06 1.00E+07

Max. Strain (%)

Number of Cycles or Cycles to Failure

Failure

No failure

N ≤ 8x105 cycles

N > 8x105 cycles
52 

 Tested and failed under dynamic loading but not at PPI – D5/ 605D1B 

PIEs were generally conducted on axial sections, while one PIE was carried out on a transverse 

section. For the tested and failed specimen, the preparation of axial 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 the 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 finish for 

metallography. Examinations were then performed on an optical microscope.16 The section 

preparation for the other segments/specimens was similar to the above procedure. 

PIE on the transverse section was conducted only on D3. The section was prepared using a 

piece of fuel rod approximately a quarter inch long cut away from the remaining half of 

specimen D3. 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. The 

section was then 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 in the following. 

An optical image of the axial section for un-tested segment 605D is shown in Figure 29. 

The three dished fuel pellets can be easily identified. Two primary fractures with sizes of pellet 

length and pellet radius were developed along and normal to the axial direction, respectively. 

Detailed study showed that the pellet-to-pellet interface near the edge of the dish and the pelletto-cladding interface were fairly close without any visible gaps. Secondary fractures had 

developed within pellets near the peripheral area close to the pellet-to-cladding interfaces. A 

cavity is also observed around some triple boundaries of pellets and cladding. The 

circumferential hydride layers can be clearly seen over the cladding thickness. The 

circumferential hydrides are widely spaced in the large middle part of the wall. The density of 

the hydrides increases significantly towards the outside of the cladding and in the claddingpellet interface area. The circumferential hydride layers are very sizable and, in the middle of 

the cladding wall, the layers can run more than 1000 μm. The outside surface of cladding is 

covered with a continuous thinner oxide layer that 

measured about 75 μm. 

 (a) (b) 

Figure 29. Optical images of untested segment 605D for (a) overall axial cross section 

and (b) enlarged area A. 

Optical images based on the axial section of S2 specimen are shown in Figure 30. Fuel failure 

occurred at the PPI as seen from the left side images. Unlike those in the untested fuel 

segment, the primary axial fractures did not align but tilted toward the axial direction. 

A 
53 

The pellet-to-cladding interface appeared to be in good contact. The fracture surface in the 

cladding exhibited a zigzag pattern both on the back and front sides of the specimen. The 

unique pattern resulted from the combination of brittle fracture of hydrides and ductile failure of 

the metal matrix under tension. Delamination can be seen over the hydride layers near the 

fracture surface. In addition, spalling and cracking of the surface oxide layer can be seen on the 

back and front sides. 

(a) 

 (b) 

 (c) 

Figure 30. (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, B – cracking of oxide layer. 

A

A

B
54 

PIEs of the other fuel segments can be found in Appendix F, including five axial sections and 

one transverse section. 

Some technical difficulty was encountered when preparing the PIE polished specimens. The 

foreign particles from the polished specimen (such as fine fuel particle due to possible micro 

fracturing of the fuel during dynamic testing) were easily brought into the working surface, 

introducing scratches. 

The following observations were made based on the PIE. 

 PPIs and PCIs near the edges or peripheral area 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 irradiated fuel. These fractures were 

observed in both untested and tested fuel segments. 

 Specimen failures occurred primarily at PPI. 

 The density of the hydride layers increases significantly toward the outside of cladding 

wall. The cladding outer wall was covered with an oxide layer of 75 m thick. 
55 

5. CONCLUSIONS 

High burnup H. B. Robinson 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 the 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-LVDTs 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. Nevertheless, the 

segment composite structure of an HBU rod also introduces numerous stress concentration 

sites into an HBU rod system, ultimately resulting in HBU specimen fractured at the pellet-pellet 

interface regions under dynamic CIRFT fatigue testing. 

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. The 

following conclusions can be drawn: 

 The HBU HBR rods survived static unidirectional bending to a maximum curvature of 3.1 to 

3.5 m1

 or a maximum moment of 86 to 87 N·m (Table 4). The maximum equivalent strain 

was 1.7 to 1.9%, corresponding to an equivalent stress of 733 to 748 MPa (Table 5). 

 The HBU 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 52 to 63 N·m2

, 

corresponding to an elastic modulus of 83 to 101 GPa (Table 5). 

 The equivalent 0.2% yield strength of the HBU cladding was estimated at 687 to 727 MPa 

(Table 5), depending on the level of hydrogen content or magnitude of oxidation thickness. 

 A large majority of the specimen failures of the HBU HBR rods under static unidirectional 

loading occurred near the PPI as validated by the post-test or PIEs. 

 The fatigue life of HBU 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 ±0.07 to ±0.49% at 5 Hz (Table 

7), the fatigue life ranged from 5.5 × 103

 to 2.3×106 cycles. 

 The -N curve of the HBU HBR rods can be described by a power function of y = 

3.5693·x0.252, 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 0.106 and 0.110% if it is 

defined at 107

 cycles. 

 Maxima of the imposed curvature in dynamic tests ranged from ±0.16 to ±1.19 m1

 at 5 Hz 

(Table 7). The -N curve of the HBU HBR rods can be described by a power function of y = 

8.1941·x0.252, where x is the number of cycles to failure and y is the maxima of cladding 

tensile curvature |max (m-1). A fatigue limit is likely located between 0.226 and 0.245 m1

 if it 

is defined at 107

 cycles. 
56 

 The failure of HBU HBR rods under cyclic reverse loading was primarily due to pelletcladding interaction and generally resided at the pellet-to-pellet interface, as validated by the 

PIE.
57 

6. REFERENCES 

 

1. 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. EPRI, 

Palo Alto, CA: 2001.1001558 

2. B. Hanson et al., Gap Analysis to Support Extended Storage of Used Nuclear Fuel Rev. 0. 

U.S. Department of Energy, Washington, DC: January 2012. 

3. J.-A. J. Wang et al., High Burn-up Spent Fuel Vibration Integrity Study Progress Letter 

Report (Out-of-Cell Fatigue Testing Development–Task 2.1). Oak Ridge National 

Laboratory, Oak Ridge, TN: January 2011. ORNL/TM-2010/288 

4. J.-A. J. Wang, et al., Progress Letter Report on U Frame Test Setup and Bending Fatigue 

Test for Vibration Integrity Study (Out-of-Cell Fatigue Testing Development–Task 2.2).

Oak Ridge National Laboratory, Oak Ridge, TN: January 2012. ORNL/TM-2011/531 

5. J.-A. J. Wang, et al., Progress Letter Report on U-Frame Test Setup and Bending Fatigue 

Test for Vibration Integrity Study (Out-of-Cell Fatigue Testing Development–Task 2.3.” 

Oak Ridge National Laboratory, Oak Ridge, TN: August 2012. ORNL/TM-2012/417 

6. J.-A. J. Wang, et al., An Innovative Dynamic Reversal Bending Fatigue Testing System for 

Evaluating Spent Nuclear Fuel Rod Vibration Integrity or Other Materials Fatigue Aging 

Performance, ORNL Invention Disclosure 201102593, DOE S 124,149, Oak Ridge 

National Laboratory, Oak Ridge, TN: April 2011. Patent in review, 13/396,413: February 

2012. 

7. H. Wang, et al., Development of U-frame Bending System for Studying the Vibration 

Integrity of Spent Nuclear Fuel. Journal of Nuclear Materials: 2013. 440, 201–213 

8. J.-A. J. Wang et al., SNF Test System for Bending Stiffness and Vibration Integrity,” 

International High-Level. Radioactive Waste Management Conference, Albuquerque, NM: 

April 2013. 

9. J.-A. J. Wang, et al., Progress Letter Report on Bending Fatigue Test System 

Development for Spent Nuclear Fuel Vibration Integrity Study (Out-of-Cell Fatigue Testing 

Development–Task 2.4). Oak Ridge National Laboratory, Oak Ridge, TN: July 2013. 

ORNL/TM-2013/225 

10. J.-A. J. Wang, et al., Reversible Bending Fatigue Test System for Investigating Vibration 

Integrity of Spent Nuclear Fuel During Transportation,” 17th International Symposium on 

the Packaging and Transportation of Radioactive Materials (PATRAM 2013), San 

Francisco, CA: August 2013. 

11. J.-A. J. Wang and H. Wang, Progress Letter Report on Reversal Bending Fatigue Testing 

of Zry-4 Surrogate Rod (Out-of-Cell Fatigue Testing Development–Task 2.4). Oak Ridge 

National Laboratory, Oak Ridge, TN: August 2013. ORNL/TM-2013/297 

12. G. Bjorkman, High Burnup Spent Fuel Testing Program Objectives. NRC Program Review 

Meeting, Oak Ridge, TN: August 2011. 

13. J.-A. J. Wang and H. Wang, 2014 Semi-Annual Progress Letter Report on Used Nuclear 

Fuel Integrity Study in Transportation Environments. Oak Ridge National Laboratory, Oak 

Ridge, TN: April 2014. ORNL/TM-2014/63 
58 

 

14. C. Baldwin, Oak Ridge National Laboratory, Mar. 6, 2013. Reproduced in Appendix B. 

15. Cladding Embrittlement during Postulated Loss-of coolant Accidents. U.S. Nuclear 

Regulatory Commission, Office of Nuclear Regulatory Research, Rockville, MD: July 2008. 

NUREG/CR-6967/ANL-07/04 

16. M.C. Billone, et al., Baseline Properties and DBTT of High-Burnup PWR Cladding Alloys,” 

17th International Symposium on the Packaging and Transportation of Radioactive 

Materials (PATRAM 2013), San Francisco, CA: August 2013. 

17. J-A Wang, H. Jiang, “Quantification of CIRFT System Biases and Uncertainties When 

Testing High-Burnup Spent Nuclear Fuel. Oak Ridge National Laboratory, Oak Ridge, TN: 

September 2014. DOE FCRFD-2014-000604 and ORNL/TM-2014/288 

18. R.S. Daum, et al., Experimental and analytical investigation of the mechanical behavior of 

high-burnup Zircaloy-4 fuel cladding. Journal of ASTM International. Paper ID JAI101209. 

19. J-A Wang, et al., Using Finite Model Analysis and Out of Hot Cell Surrogate Rod Testing to 

Analyze High Burnup Used Nuclear Fuel Mechanical Properties. Oak Ridge National 

Laboratory, Oak Ridge, TN: August 2014. ORNL/TM-2014/257 and DOE FCRD-UFD-2014-

000603, 
APPENDIX A 

ROD SEGMENTS AND CUTTING PLAN 

 

A-1 

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 

* 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-2 

(a) 

S2 S1 DL3 DM1 DM2 

Used for 

drill 

practice 
A-3 

(b) 

Demo 

1 

Demo 

2 R5 R4 R3 S5 S4 S3 DM3 
A-4 

(c) 

DL1 R1 
A-5 

(d) 

DL2 R2 
A-6 

(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 
APPENDIX B 

METROLOGY OF ROD SEGMENTS 

B-1 

Table B.1. Diameter measurements on fuel rod segment R5 (606B3C) (unit: inch)* 

Distance 

from End 

(inch) 

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 0.42368 B - (120) 0.42428 0.42416 0.42400 0.42416 0.42392 0.42396 0.42370 0.42404 0.42402 0.42374 0.42382 C - (240) 0.42432 0.42430 0.42386 0.42412 0.42374 0.42390 0.42446 0.42506 0.42474 0.42384 0.42382 Maximum 0.42450 0.42436 0.42450 0.42440 0.42402 0.42396 0.42446 0.42506 0.42474 0.42392 0.42382 Minimum 0.42428 0.42416 0.42386 0.42412 0.42374 0.42390 0.42370 0.42400 0.42402 0.42374 0.42368 Mean 0.42437 0.42427 0.42412 0.42423 0.42389 0.42392 0.42401 0.42437 0.42435 0.42383 0.42377 Std. Dev. 0.00012 0.00010 0.00034 0.00015 0.00014 0.00003 0.00040 0.00060 0.00036 0.00009 0.00008 OverallMaximum 0.42506Minimum 0.42368Mean 0.42410 Std. Dev. 0.00032

*Based on C. Baldwin, ORNL, Mar. 6, 2013. 
B-2 

Table B.2. Diameter measurements on HBR fuel rod segments (unit: inch)* 

Spec 

Label 

ORNL 

Label 

NRC 

Original 

Label 

Seg. 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 0.00012 S2 S2 S2 605D1E 0.42322 0.42298 0.42298 0.42310 0.42310 0.42310 0.42308 0.00009 S3 DCAL DCAL 609C5 0.42222 0.42248 0.42272 0.42298 0.42288 0.42288 0.42269 0.00029 S4 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 0.00021 D2 DL2 DL2 608C4B 0.42442 0.42422 0.42410 0.42482 0.42462 0.42454 0.42445 0.00026 D3 DL3 DL3 605C10A 0.42398 0.42402 0.42422 0.42444 0.42418 0.42420 0.42417 0.00016 D4 DM1 DM1 605D1C 0.42314 0.42336 0.42352 0.42302 0.42310 0.42316 0.42322 0.00019 D5 DM2 DM2 605D1B 0.42376 0.42412 0.42396 0.42426 0.42424 0.42410 0.42407 0.00019 D6 DH1 DH1 609C4 0.42338 0.42336 0.42286 0.42288 0.42310 0.42322 0.42313 0.00023 D7 DH2 DH2 609C3 0.42464 0.42404 0.42394 0.42430 0.42446 0.42426 0.42427 0.00026 D8 DM3 DM3 606C3E 0.42250 0.42280 0.42268 0.42284 0.42292 0.42292 0.42278 0.00016 D9 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 0.00020 D11 R1 R1 607C4A 0.42642 0.42690 0.42676 0.42614 0.42616 0.42644 0.42647 0.00031 D12 R2 R2 608C4A 0.42630 0.42588 0.42618 0.42622 0.42574 0.42592 0.42604 0.00022 D13 R3 R3 606B3E 0.42402 0.42396 0.42408 0.42334 0.42356 0.42400 0.42383 0.00030 D14 R4 R4 606B3D 0.42416 0.42404 0.42378 0.42368 0.42368 0.42358 0.42382 0.00023 D15 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 0.00012 Scal2 S4 S4 606C3B 0.42376 0.42354 0.42340 0.42390 0.42376 0.42374 0.42368 0.00018 

 Average 0.42391  Std. Dev. 0.00112 

*Based on C. Baldwin, ORNL, Mar. 6, 2013.
APPENDIX C 

BENCHMARKING OF TEST SYSTEM 

C-1 

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 m oxide layer, 750 ppm H content) was 

initially tested to the relative displacement of 2 mm, at rate 0.1 mm/s at each loading point on 

the U-frame. Such a displacement level was quite low so as 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, same applies in the following unless specified 

otherwise), and the rod appeared to be quite 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 above loading processes are given 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 gage 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 given 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 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 –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. 

It thus can be seen from testing of the first fuel rod that: 

 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 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. 

 The testing setup worked as designed, confining the final failure within the gage section of 

the rod. 
C-2 

 

(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. 

(a) (b) 

(c) 

Figure C.2. Variations in (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for Demo1 (606B2); Nf = 3.96×103

 cycles under 

±30.48 N·m 5 Hz. The specimen was subjected to repeats of static loading before the cycle test. 
C-3 

 

(a) (b) 

Figure C.3. Mating fracture surfaces of Demo1 (606B2); the reference marker indicates 

the frontal maximum stress line (facing the operator); Nf = 3.96×103

 cycles under ±30.48 

N·m 5 Hz. The specimen had been subjected to repeat static loading cycles before the dynamic 

cycle test. 

2. STATIC TEST ON SCAL1/ DEMO2/ 606B3B 

The test on Scal1/ Demo2 (606B3B, 66.5 GWd/MTU burnup, 100–110 m 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 gage 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. So, there is a substantial mechanical property difference 

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 gage 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. It is 

worthwhile noting that circumferential cracks and spalling also occurred on Demo1 and 

several other rods, as discussed below. 
C-4 

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. 
C-5 

 

(a) (b) 

(c) 

(d) 

Figure C.5. (a, b) Mating fracture surfaces, (c) front side (compressive), and (d) back 

side (tensile) of Scal1/ Demo2 (606B3B). The rate was 0.1 mm/sec at the U-frame loading 

points. The rod broke at peak moment = 78.2 N·m. 

3. DYNAMIC TEST ON DCAL/ S3/ 605D1F (±25.4 N·M 5 HZ) 

Dynamic cycle testing was conducted on rod Dcal/ S3 (605D1F, 66.5 GWd/MTU burnup, 40–70 

m oxide layer, 360 ppm H content) as a part of system benchmarking. The cycle test was 

conducted under ±250 N or 25.40 N·m 5 Hz. The system has been demonstrated to meet the 

testing requirements for loading and monitoring. A lifetime of 2.5×104

 cycles was observed. 

Flexural rigidity was shown to be fairly stable as the cycles accumulated prior to final fracture, 

as shown in Figure C.6. The associated measurements are shown in Figures C.7 and C.8. 

Failure was observed in the gage section near motor 2. A clear cut can be seen from the 

exposed end faces of neighboring pellets as shown in Figure C.9. 
C-6 

 Rod failure was dominated by the pellet-to-pellet interface. 

 The lateral surface of the rod was shown to be free of the cracks or spalling observed in 

Demo1, Scal1, and Scal2 (will be discussed below). It is not clear if the absence of circular 

cracks was related to the small loading amplitude or to the lesser degree of preexisting 

oxidation and hydrides in Dcal. 

(a) (b) 

(c) (d) 

(e) 

Figure C.6. Variations in (a) curvature range, (b) applied moment range, (c) flexural 

rigidity, (d) maximum and minimum values of curvature, and (e) maximum and minimum 
C-7 

values of moment as a function of number of cycles for Dcal/ S3 (605D1F); Nf = 2.5×104

cycles under ±25.4 N·m 5 Hz. 

 

(a) (b) 

Figure C.7. (a) Moment-curvature relation, and (b) moment-flexural rigidity relation at 

various numbers of cycles for Dcal/ S3 (605D1F); Nf = 2.5×104

 cycles under ±25.4 N·m 

5 Hz. 

 

(a) (b) 

(c) 

Figure C.8. Variations in (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for Dcal/ S3 (605D1F); Nf = 2.5×104

 cycles under 

±25.4 N·m 5 Hz. 
C-8 

 

(a) (b) 

(c) 

(d) 

Figure C.9. (a, b) Mating fracture surfaces, (c) front side, and (d) back side of Dcal/ S3 

(605D1F); Nf = 2.5×104

 cycles under ±25.40 N·m 5 Hz. 

4. STATIC TEST ON SCAL2/ S4/ 606C3B 

Scal2/ S4 (606C3B, 66.8 GWd/MTU burnup, 70–100 m oxide layer, 700 ppm H content) 

fractured at the peak moment near 80.7 N·m at a loading of 0.1 mm/s. A large variety of 

fragment sizes were revealed in the postmortem analysis (Figure C.10). 

 The end face of a pellet can be seen clearly from the brownish haze zone. So, the failure 

took place near a pellet-to-pellet interface with the fracture involving a neighboring pellet. 
C-9 

 Besides the circumferential cracks, a sharp attachment of cladding was outstanding on the 

frontal or compressive side. On the other side, spalling was not as extensive as in Demo1 

and Scal1. 

 Compared to Scal1 from the same fuel rod but a different segment (606B), the slightly 

higher failure moment and less spalling in Scal2 are coincident with the lower levels of 

oxide-layer thickness and H content of the rod segment. 

 

(a) (b) 

(c) 

(d) 

Figure C.10. (a, b) Mating fracture surfaces, (c) front (compressive) side, and (d) back 

(tensile) side of Scal2/ S4 (606C3B). The rate at the U-frame loading points was 0.1 mm/sec. 

The rod broke at peak moment = 80.7 N·m. 

APPENDIX D 

STATIC TEST RESULTS 

D-1 

Static tests were conducted on October 7–8, 2013. A brief description of four static tests and 

their results is given below. 

1. S1/ 606C3C (+ DYNAMIC) 

S1 (606C3C, 66.8 GWd/MTU burnup, 70–100 m 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 [Figure D.2 (c, d)] near the rod’s neutral axis of bending were free of 

circumferential cracks to a great extent. On the other hand, one of the stressed lateral surfaces 

was covered with a large number of circumferential cracks and related subsurface spalling 

[Figure D.2 (d)]. It is uncertain if 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. 
D-2 

 

(a) (b) 

(c) 

(d) 

Figure D.2. (a, b) Mating fracture surfaces and (c, d) lateral sides near the rod’s neutral 

axis of bending for S1 (606C3C), which was subjected to four repeated static cycles to 

24 mm maximum relative displacement followed by dynamic cyclic loading ±25.40 to 

30.48 N·m 5 Hz; 1.4×104

 failure cycles, 0.9 gram of fuel particles collected. 

2. S2/ 605D1E (+ DYNAMIC) 

S2 (605D1E, 66.5 GWd/MTU burnup, 40–70 m oxide layer, 400 ppm H content) was tested by 

following a routine similar to that of S1. The specimen survived three cycles of large 

displacement loading and 7.2×103

 cycles under ±30.48 N·m 5 Hz before failure at the middle of 

the gage section. 

The moment-curvature curve and equivalent stress-strain curve for the initial cycle are 

presented in Figure D.3. Again, the turning points were revealed near 20 and 62 N·m. 
D-3 

Fracture occurred on a pellet-to-pellet interface as can be seen from Figure D.4, in which the 

smooth brownish end face of the adjacent pellet can be seen clearly. One of the attachments 

[Figure D.4 (b)] is convex-shaped and covers almost 1/3 of the end face, and several small ones 

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 agrees 

with the results of Dcal/ S3 as they were prepared from same segment (605D) with the same 

level of preexisting oxides and hydrides. 

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. 
D-4 

 

(a) (b) 

(c) 

(d) 

Figure D.4. (a, b) Mating fracture surfaces, (c) front side, and (d) back side of S2 

(605D1E), which was subjected to three repeated static cycles to 24 mm maximum 

relative displacement followed by dynamic cyclic loading ±30.48 N·m 5 Hz; 7.2×103

 failure 

cycles, 0.6 gram of fuel particles collected. 

3. S3/ DCAL/ 609C5 (+ DYNAMIC) 

S3/Dcal (609C5, 66.5 GWd/MTU burnup, 70–100 m oxide layer, 550 ppm H content) was 

subjected to three repeated static cycles of 24 mm displacement loading followed by cyclic 

loading ±30.48 N·m 5 Hz. The S3 failed around 9.6×103

 cycles at the middle of the gage section 

with about 0.2 gram of fuel released. 
D-5 

A three stage moment-curvature curve was obtained as shown in Figure D.5, in which two 

turning points can be seen near 18 and 60 N·m. 

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 whose size was about a quarter the diameter of a pellet. 

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 those rods 

whose failure actively involves the pellet-to-pellet interface, where 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, 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. 
D-6 

 

(a) (b) 

(c) 

(d) 

Figure D.6. (a, b) Mating fracture surfaces, (c) front side, and (d) back side of S3/ Dcal 

(609C5), which was subjected to three repeated static loadings to 24 mm maximum 

relative displacement followed by dynamic loading: ±30.48 N·m 5 Hz; 9.6×103

 failure 

cycles, 0.2 gram of fuel particles collected. 

4. S4/ SCAL/ 609C6 

S4/ Scal (609C6, 66.5 GWd/MTU burnup, 70–100 m oxide layer, 550 ppm H content) 

sustained three cycles of loading with relative displacement of 24 mm on the U-frame and an 

additional load with 25 mm before failure. Failure occurred around the peak moment 86.1 N·m. 

The curves for the initial cycle exhibited a smooth transit from linear to nonlinear as shown in
D-7 

Figure D.7. However, the turning point near 20 N·m is still appreciable based on the 

discontinuous change in the slope of the moment-curvature curve. 

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 lot 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. 
D-8 

 

(a) (b) 

(c) 

(d) 

Figure D.8. (a, b) Mating fracture surfaces and (c, d) lateral sides near the rod’s neutral 

axis of bending for S4/ Scal (609C6), which was subjected to repeated static loading to 

24 mm maximum relative displacement and failed at 86.1 N·m in the 4th cycle; 1.3 grams 

of fuel particles collected. 
APPENDIX E 

DYNAMIC TEST RESULTS

E-1 

Dynamic testing began on October 22, 2013, and is still ongoing. A description of completed 

dynamic tests and their results is given below. 

1. D1/ DL1/ 607C4B (±15.24 N·M 5 HZ) 

The test on D1/ DL1 (607C4B, 63.8 GWd/MTU burnup, 70–100 m 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, 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. It is interesting that the loops actually became quite slender with the accumulated 

cycles, whilst the amplitude of the loops indeed 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 number of cycles are shown in Figure E.3. While the curvature stayed 

at a quite 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 on-line 

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 quite higher than those used in measurements as part of the experiment 

design. Accordingly, the flexural rigidity based on on-line monitoring was about 50 N·m2

, 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 quite stable 

response prior to its final breakage. 

Failure was observed in the gage 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 quite clean. At the same time, the fracture surfaces of the cladding were 

serrated. Some striations had developed in the circumferential direction, but they are hardly 

differentiated at this magnification. 

Both lateral stressed sides of the rod were found covered with equally spaced circumferential 

cracks throughout the gage section. Spalling only occurred at the local area near the fracture. 

The 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 but less on the other 

side, as seen, for example, in S3/ Dcal (see 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. It is thus believed that the parallel equally spaced cracks were related to a preexisting 

condition as D1/ DL1 actually had a thicker layer of oxidation and a higher H content. 
E-2 

 

(a) (b) 

Figure E.1. Moment-curvature loops based on measurements when (a) N = 1 and (b) N = 

111000 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. 
E-3 

(a) (b) 

(c) 

Figure E.3. Variations of (a) curvature range, (b) moment range, and (c) flexural rigidity 

as a function of number of cycles for D1/ DL1 (607C4B); Nf = 1.1×105

 cycles under ±15.24 

N·m 5 Hz. Fuel particles collected < 1.0 gram. 
E-4 

 

(a) (b) 

 

(c) (d) 

(e) 

Figure E.4. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for D1/ DL1 (607C4B); Nf = 1.1×105

cycles under ±15.24 N·m 5 Hz. Fuel particles collected < 1.0 gram. 
E-5 

 

(a) (b) 

(c) 

(d) 

Figure E.5. (a, b) Mating fracture surfaces, (c) front side, and (d) back side of D1/ DL1 

(607C4B), Nf = 1.1×105

 cycles under ±15.24 N·m 5 Hz. Fuel particles collected < 1.0 gram. 

2. D2/ DL2/ 608C4B (±5.08/35.56 N·M 5 HZ) 

D2/ DL2 (608C4B, 63.8 GWd/MTU burnup, 70–100 m oxide layer, 700 ppm H content) was 

tested under ±5.08 N·m 5 Hz. The measurement was performed using 0.4 mm relative 

displacement. The rod sustained more than 6.4×106

 cycles of loading without showing any 

changes in flexural rigidity in both the on-line monitoring and measurements as shown in 

Figures E.6 and E.7 (an earlier stage of this test is not shown because the LVDT3 channel did 

not generate reliable data owing to loose wires). 
E-6 

 

(a) (b) 

 

(c) (d) 

(e) 

Figure E.6. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for D2/ DL2 (608C4B); 6.4×104

 cycles 

completed under ±5.08 N·m 5 Hz without failure. 
E-7 

 

(a) (b) 

(c) 

Figure E.7. Variations of (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for D2/ DL2 (608C4B) based on measurements 

with maximum relative displacement 0.4 mm; 6.4×106

 cycles completed under ±5.08 N·m 

5 Hz without failure. 

A subsequent cycle test using ±35.56 N·m 5 Hz was conducted to investigate the failure mode 

of the tested specimen. The specimen failed near motor 2 (end cap A of the rod specimen) 

around 1.8×103

 cycles. The clean fracture surfaces revealed that failure occurred at the pelletpellet interface (Figure E.8). The lateral view showed that the fracture surface was not a flat cut 

with a sharp attachment on the front side of the rod with extensive spalling. In addition, 

circumferential cracks covered the stressed surfaces of the rod. 
E-8 

 

(a) (b) 

(c) 

(d) 

Figure E.8. (a, b) Mating fracture surfaces, (c) front side, and (d) back side of D2/ DL2 

(608C4B); follow-up test was conducted with increased amplitude ±35.56 N·m 5 Hz with Nf

= 1.8×103

 cycles. 

3. D3/ DL3/ 605C10A (±10.16 N·M 5 HZ) 

D3/ DL3 (605C10A, 66.5 GWd/MTU burnup, 70–100 m oxide layer, 550 ppm H content) was 

tested with ±10.16 N·m 5 Hz at a lifetime of 1.0×106

 cycles. The monitoring data revealed a 

slight increase in flexural rigidity prior to failure, as seen in Figure E.9. The measurements using 

0.4 and 0.8 mm demonstrated a trend of fall and rise as shown in Figure E.10. However, the 

final rigidity values were still lower than the respective pre-fatigue levels. The specimen failed 

near motor 2 (left side, end cap A of the rod specimen). The fracture surface indicates that 

failure occurred at the pellet-pellet interface as the end faces of involved pellets can be seen 

clearly (Figure E.11). No circumferential crack was observed on the stressed surfaces of the 

rod. 
E-9 

 

(a) (b) 

 

(c) (d) 

(e) 

Figure E.9. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for D3/ DL3 (60510A); Nf = 1.0×106

cycles under ±10.60 N·m 5 Hz. 
E-10 

 

(a) (b) 

(c) 

Figure E.10. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

rigidity as a function of number of cycles for D3/ DL3 (60510A) based on measurements 

with maximum relative displacement at 0.4 and 0.8 mm; Nf = 1.0×106

 cycles under ±10.16 

N·m 5 Hz. 
E-11 

 

(a) (b) 

(c) 

(d) 

Figure E.11. (a, b) Mating fracture surfaces, (c) front side, and (d) back side of D3/ DL3 

(605C10A); Nf = 1.0×106

 cycles under ±10.16 N·m 5 Hz. 

4. D4/ DM1/ 605D1C (±7.62 N·M 5 HZ) 

D4/ DM1 (605D1C, 66.5 GWd/MTU burnup, 40–70 m oxide layer, 500 ppm H content) was 

tested under ±7.62 N·m 5 Hz as the result of an effort to define a possible fatigue limit for the 

SNF rod. DM1 finished the target of 1.1×107 cycles without any sign of significant fatigue during 

the cyclic test, and the test was stopped. 

On-line monitoring demonstrated that the flexural rigidity fluctuated between 42 and 49 N·m2

with a marginal decrease near the end of the test (Figure E.12). The measurements using 0.4 
E-12 

and 0.6 mm also demonstrated a certain amount of drop in rigidity near the end-of-cycle test, as 

shown in Figure E.13. 

 

(a) (b) 

 

(c) (d) 

(e) 

Figure E.12. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for D4/ DM1 (605D1C); N = 1.0×107

cycles under ±7.62 N·m 5 Hz. 
E-13 

 

(a) (b) 

(c) 

Figure E.13. Variations of (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for D4/ DM1 (605D1C) based on measurements 

with maximum relative displacement at 0.4 and 0.6 mm; N = 1.0×107

 cycles under 

±7.62 N·m 5 Hz. 

5. D5/ DM2/ 605D1B (±9.14 N·M 5 HZ) 

The cycle test on D5/ DM2 (605D1B, 66.5 GWd/MTU burnup, 40–70 m oxide layer, 550 ppm H 

content) was conducted under ±9.14 N·m 5 Hz. DM2 fractured near 2.3×106

 cycles. 

On-line monitoring demonstrated that the flexural rigidity fluctuated between 47 and 51 N·m2

with a marginal decrease near the end of test (Figure E.14). The measurements using 0.4 and 

0.8 mm also showed a similar degree of drop in rigidity near the end of the test, as shown in 

Figure E.15. 

DM2 fractured in the gage section near motor 2 (left side) as shown in Figure E.16. The 

involvement of the pellet-pellet interface was not quite clear because no end face of a pellet can 

be identified. No circumferential cracking can be seen on either of the stressed surfaces of the 

rod, but spalling was observed near the fracture. 
E-14 

 

(a) (b) 

(c) (d) 

(e) 

Figure E.14. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for D5/ DM2 (605D1B); Nf = 2.3×106

cycles under ±9.14 N·m 5 Hz. 
E-15 

 

(a) (b) 

(c) 

Figure E.15. Variations of (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for D5/ DM2 (605D1B) based on measurements 

with maximum relative displacement at 0.4 and 0.8 mm; Nf = 2.3×106

 cycles under ±9.14 

N·m 5 Hz. 
E-16 

 

(a) (b) 

(c) 

(d) 

Figure E.16. (a, b) Mating fracture surfaces, (c) front side, and (d) back side of D5/ DM2 

(605D1B); Nf = 2.3×106

 cycles under ±9.14 N·m 5 Hz. 

6. D6/ DH1/ 609C4 (±12.7 N·M 5 HZ) 

The cycle test on D6/ DH1 (609C4, 66.5 GWd/MTU burnup, 70–100 m oxide layer, 700 ppm H 

content) was conducted under ±12.7 N·m 5 Hz. D6/ DH1 fractured around 2.4×105

 cycles. 

On-line monitoring demonstrated that the flexural rigidity remained between 42 and 45 N·m2

over the cycle test with a slight increase (Figure E.17Figure E.17). The measurements using 0.4 

and 0.8 mm also showed a similar variation in rigidity as shown in Figure E.18. 
E-17 

DH1 fractured in the gage section near motor 2 (left side) as shown in Figure E.19. The 

involvement of the pellet-pellet interface was not quite clear because no end face of a pellet can 

be seen. However, equally spaced circumferential cracks exist on both stressed surfaces of the 

rod. 

 

(a) (b) 

(c) (d) 

(e) 

Figure E.17. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for D6/ DH1 (609C4); Nf = 2.4×105

cycles under ±12.7 N·m 5 Hz. 
E-18 

(a) (b) 

(c) 

Figure E.18. Variations of (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for D6/ DH1 (609C4) based on measurements 

with maximum relative displacement at 0.4 and 0.8 mm; Nf = 2.4×105

 cycles under ±12.7 

N·m 5 Hz. 
E-19 

 

(a) (b) 

(c) 

(d) 

Figure E.19. (a, b) Mating fracture surfaces, (c) front side, and (d) back side of D6/ DH1 

(609C4); Nf = 2.4×105

 cycles under ±12.7 N·m 5 Hz. 

7. D7/ DH2/ 609C3 (±20.32 N·M 5 HZ) 

The cycle test on D7/ DH2 (609C3, 66.5 GWd/MTU burnup, 70–100 m oxide layer, 750 ppm H 

content) was conducted under ±20.32 N·m 5 Hz. DH2 fractured around 6.5×104

 cycles. 

On-line monitoring displayed a steady decrease throughout the cycle test with a sudden drop 

near the end (Figure E.20). The measurements using 0.8 and 1.2 mm also showed a similar 

rigidity variation as given in Figure E.21. 
E-20 

DH2 fractured in the middle of the gage section on a pellet-pellet interface as can be seen from 

the clean cut in Figure E.22. Equally spaced circumferential cracks can be observed over both 

stressed surfaces of the rod with some spalling near the fracture. 

(a) (b) 

(c) (d) 

(e) 

Figure E.20. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for D7/ DH2 (609C3); Nf = 6.5×104

cycles under ±20.32 N·m 5 Hz. 
E-21 

 

(a) (b) 

(c) 

Figure E.21. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

rigidity as a function of number of cycles for D7/ DH2 (609C3) based on measurements 

with maximum relative displacement at 0.8 and 1.2 mm; Nf = 6.5×104

 cycles under ±20.32 

N·m 5 Hz. 
E-22 

 

(a) (b) 

(c) 

(d) 

Figure E.22. (a, b) Mating fracture surfaces, (c) front side, and (d) back side of D7/ DH2 

(609C3); Nf = 6.5×104

 cycles under ±20.32 N·m 5 Hz. 

8. D8/ DM3/ 606C3E (±8.89 N·M 5 HZ) 

D8/ DM3 (606C3E, 66.8 GWd/MTU burnup, 70–100 m oxide layer, 550 ppm H content) 

sustained quite a long testing period under ±8.89 N·m 5 Hz. Testing was stopped after 1.3×107

cycles. 

On-line monitoring displayed a stable flexural rigidity response, followed by a fluctuation or falland-rise near 4×106

 cycles (Figure E.23). The fluctuation was limited within the range of 45 to 

50 N·m2

. The measurements using 0.8 and 1.2 mm also revealed a variation of a quite similar 

pattern in rigidity, as shown in Figure E.24. 
E-23 

 

(a) (b) 

(c) (d) 

(e) 

Figure E.23. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for D8/ DM3 (606C3E); N = 1.3×107

cycles under ±8.89 N·m 5 Hz. 
E-24 

(a) (b) 

(c) 

Figure E.24. Variations of (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for D8/ DM3 (606C3E) based on measurements 

with maximum relative displacement at 0.4 and 0.8 mm; N = 1.3×107

 cycles under ±8.89 

N·m 5 Hz. 

9. D9/ DH3/ 609C7 (±35.56 N·M 5 HZ) 

The cycle test on D9/ DH3 (609C7, 66.5 GWd/MTU burnup, 70–100 m oxide layer, 550 ppm H 

content) was conducted under ±35.56 N·m 5 Hz. The specimen failed near 7.13×103

 cycles 

within the gage section. 

On-line monitoring exhibited a relatively flat response in flexural rigidity before 6.5×102

 cycles, 

and then fluctuated or rose substantially as shown in Figure E.25. Although the cause of the 

fluctuation was not clear, it can be seen that a sudden drop occurred prior to rod failure. The 

measurements using 0.8, 1.2, and 1.6 mm relative displacements also revealed a considerable 

increase in rigidity at 103

 cycles as shown in Figure E.26. The increase observed in both on-line 

monitoring and measurement remains to be studied. 
E-25 

(a) (b) 

(c) (d) 

(e) 

Figure E.25. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of as a function of number of cycles for D9/ DH3 (609C7); Nf = 7.13×103

 cycles 

under ±35.56 N·m 5 Hz. 
E-26 

 

(a) (b) 

(c) 

Figure E.26. Variations of (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for D9/ DH3 (609C7); Nf = 7.13×103

 cycles under 

±35.56 N·m 5 Hz. 

10. D10/ S5/ 606C3A (±12.70 N·M 5 HZ) 

The cycle test on D10/ S5 (606C3A, 66.8 GWd/MTU burnup, 70–100 m oxide layer, 750 ppm 

H content) was conducted under ±12.70 N·m 5 Hz. The testing condition was the same as that 

used in DH1, where failure took place near the edge of the rigid sleeve. The test was repeated 

to confirm the lifetime and failure position under the same testing condition. S5, in fact, failed 

around 1.81×105

 cycles within the gage section. The lifetime was thus a little shorter than that of 

DH1. The difference may be related to the preexisting condition as a little higher estimated H 

content in S5 was seen. 

The equivalent decrease in flexural rigidity prior to final failure in both monitoring and 

measurement can be seen clearly from Figures E.27 and E.28, respectively. Detailed 

examination showed a certain degree of curvature waveform drift toward the positive direction. 

The positively biased curvature wave corresponded to an absolute extreme value as great as 

0.4 m-1. The level of extreme curvature in D10/ S5 was clearly higher than that observed in 

D6/ DH1, which was also partly responsive to the shorter lifetime in the former case. 
E-27 

 

(a) (b) 

 

(c) (d) 

(e) 

Figure E.27. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for D10/ S5 (606C3A); Nf = 1.81×105

cycles under ±12.70 N·m 5 Hz. 
E-28 

 

(a) (b) 

(c) 

Figure E.28. Variations of (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for D10/ S5 (606C3A); Nf = 1.81×105

 cycles 

under ±12.70 N·m 5 Hz. 

11. D11/ R1/ 607C4A (±30.48 N·M 5 HZ) 

The cycle test on D11/ R1 (606C3A, 63.8 GWd/MTU burnup, 70–100 m oxide layer, 700 ppm 

H content) was conducted under ±30.48 N·m 5 Hz. R1 failed around 5.5×103

 cycles within the 

gage section. Significant fall and rise were observed in on-line monitoring between 1×103

 and 

4×103

 cycles as shown in Figure E.29. On the other hand, the measurements using 0.8, 1.2, 

and 1.6 mm relative displacement at loading points demonstrated a defined decrease in flexural 

rigidity as seen in Figure E.30. Such lifetime was obviously shorter than that expected from the 

result of D9/ DH3, where a longer lifetime was obtained with an even higher moment level. 

Considering that the H content of D9/ DH3, 550 ppm, was much lower than that of D11/ R1, it is 

believed that the preexisting condition of specimens played a certain role in controlling failure in 

high burnup, irradiated rods as can be seen from the estimated H contents in both rod 

segments. 
E-29 

 

(a) (b) 

 

(c) (d) 

(e) 

Figure E.29. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for D11/ R1 (607C4A); Nf = 5.5×103

cycles under ±30.48 N·m 5 Hz. 
E-30 

 

(a) (b) 

(c) 

Figure E.30. Variations of (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for D11/ R1 (607C4A); Nf = 5.5×103

 cycles under 

±30.48 N·m 5 Hz. 

12. D12/ R2/ 608C4A (±11.18 N·M 5 HZ) 

The cycle test on D12/ R2 (608C4A, 63.8 GWd/MTU burnup, 70–100 m oxide layer, 700 ppm 

H content) was conducted under ±11.18 N·m 5 Hz. R2 failed around 3.86×105

 cycles within the 

gage section. 

A certain amount of fluctuation was observed in on-line monitoring and measurement data sets 

as illustrated in Figures E.31 and E.32, respectively. This can be seen around 104 cycles. The 

peak/valley monitoring indicated that the curvature waveform down shifted a little throughout the 

cycle test process. An abrupt drop in flexural rigidity was captured prior to the final failure in both 

monitoring and measurement. 
E-31 

(a) (b) 

(c) (d) 

(e) 

Figure E.31. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for D12/ R2 (608C4A); Nf = 3.86×105

cycles under ±11.18 N·m 5 Hz. 
E-32 

 

(a) (b) 

(c) 

Figure E.32. Variations of (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for D12/ R2 (608C4A); Nf = 3.86×105

 cycles 

under ±11.18 N·m 5 Hz. 

13. D13/ R3/ 606B3E (±13.72 N·M 5 HZ) 

The cycle test on D13/ R3 (606B3E, 66.5 GWd/MTU burnup, 100–110 μm oxide layer, 750 ppm 

H content) was conducted under ±13.72 N·m 5 Hz. R3 failed around 1.29×105 cycles within the 

gage section. 

Testing revealed that the on-line monitoring rigidity sustained a certain increase prior to the final 

drop at failure because of the continuous decrease in curvature range as shown in Figure E.33. 

The monitoring of the peak and valley of curvature indicated that the decreased curvature range 

is mainly due to the drop in peak value. 

The cause of such single-side change in curvature waveform during the cyclic test is unknown, 

but decreased rigidity was indeed seen in the measurement shown in Figure E.34. 
E-33 

 

(a) (b) 

 

(c) (d) 

(e) 

Figure E.33. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for D13/ R3 (606B3E); Nf = 1.29×105

cycles under ±13.72 N·m 5 Hz. 
E-34 

(a) (b) 

(c) 

Figure E.34. Variations of (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for D13/ R3 (606B3E); Nf = 1.29×105

 cycles 

under ±13.72 N·m 5 Hz. 

14. D14/ R4/ 606B3D (±8.89 N·M 5 HZ) 

The cycle test on D14/ R4 (606B3D, 66.5 GWd/MTU burnup, 100–110 m oxide layer, 750 ppm 

H content) was conducted under ±8.89 N·m 5 Hz. This testing condition was the same as that of 

DM3. The test on R4 was used to generate an additional observation on the same testing 

condition as DM3. R4 exhibited a relatively short lifetime with 2.7×105

 cycles to failure. The 

short lifetime may be due to a different preexisting specimen condition as can be seen from the 

estimated H content. The H content of R4 was around 750 ppm, whilst that of DM3 was as low 

as 550 ppm. 

On-line monitoring demonstrated a quite flat response of rigidity with a slight drop prior to final 

failure as shown in Figure E.35. The measurement showed that rigidity increased slightly, also 

as seen in Figure E.36. 
E-35 

 

(a) (b) 

(c) (d) 

(e) 

Figure E.35. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for R4 (606B3D); Nf = 2.7×105

 cycles 

under ±8.89 N·m 5 Hz. 
E-36 

 

(a) (b) 

(c) 

Figure E.36. Variations of (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for R4 (606B3D); Nf = 2.7×105

 cycles under 

±8.89 N·m 5 Hz. 

15. D15/ R5/ 606B3C (±7.62 N·M 5 HZ) 

The cycle test on R5 (606B3C, 66.5 GWd/MTU burnup, 100–110 m oxide layer, 750 ppm H 

content) was conducted under ±7.62 N·m 5 Hz. The moment range used for R5 is same as that 

used in DM1. The test on R5 thus provides the opportunity to examine the variation of fatigue 

response between different rods. The test on R5 was stopped after 2.23×107

 cycles had been 

accumulated because no sign of failure could be seen. 

The analysis based on the monitoring data indicated that the flexural rigidity was fairly stable 

and limited within a range of 40 to 50 N•m2

. A drop was observed after rod was taken off and remounted to the testing machine. The reason was not clear at this moment; however the rod 

appeared to be quite stable in subsequent cycling process. 

The on-line monitoring and measurement data are given In Figures E.37 and E.38, respectively. 
E-37 

 

(a) (b) 

 

(c) (d) 

(e) 

Figure E.37. Variations of (a) curvature range, (b) applied moment range, (c) flexural 

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

values of moment as a function of number of cycles for R5 (606B3C); N = 2.23×107

 cycles 

under ±7.62 N·m 5 Hz. 
E-38 

(a) (b) 

(c) 

Figure E.38. Variations of (a) curvature range, (b) applied moment range, and (c) flexural 

rigidity as a function of number of cycles for R5 (606B3C); N = 2.23×107

 cycles under 

±7.62 N·m 5 Hz. 
APPENDIX F 

POST IRRADIATION EXAMINATION (PIE)

F-1 

1. PREPARATION OF SPECIMENS 

Axial section 

PIE of the axial section was conducted on one untested rod segment, on a fractured specimen 

from static test S2, and on three specimens tested in dynamic tests [one un-failed (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 the 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 

finish for metallography. Examinations were then performed on 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 and with the same 

estimated amount of hydrogen content as those in other PIE specimens. The section for the 

tested and un-failed specimen was taken from the gage section of the specimen. 

Transverse section 

PIE on the transverse section was conducted only on D3. 

The section was prepared by using an approximately 0.25 in. 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). 

2. UNTESTED SEGMENT 605D1D 

The untested section was prepared from one inch of fuel 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.1 (a). 

The three dished fuel pellets can be easily identified. Two primary fractures had developed 

approximately along and normal to the axial direction. Detail study [Figure F.1 (b)] showed that 

the pellet-to-pellet interface near the edge of the dish and the pellet-to-clad interface were fairly 

close without any visible gaps. Secondary fractures had developed within pellets near the 

peripheral area close to the pellet-to-clad interfaces. A cavity is also observed around some 

triple boundaries of pellets and clad. The circumferential hydride layers can be seen clearly over 

the whole thickness. The circumferential hydrides are widely spaced in the large middle part of 

the wall. The density of the hydrides increases significantly towards the outside of the clad and 

in the clad-pellet interface area. The circumferential hydride layers are very sizable and, in the 

middle of clad wall, the layers can run more than 1000 μm. The outside surface of clad is 

covered with a continuous thinner oxide layer that measured about 75 μm. 
F-2 

(a) (b)

Figure F.1. Optical images for (a) overall axial cross section and (b) enlargement of area 

indicated in (a). 

3. FRACTURED IN STATIC TEST S2/ 605D1E 

S2 was based on segment 605D1E, whose estimated hydrogen and oxide layer thickness were 

400 ppm and 40–70 μm. S2 failed in a dynamic test (±30.48 N·m, 5 Hz) after three cycles of 

unidirectional loading. 

The failure took place at the pellet-to-pellet interface as seen in Figure F.2. Primary fractures in 

the same range of pellet size can be seen clearly. Unlike those in the untested fuel segment, the 

axial fractures did not align but tilted toward the axial direction. 

The pellet-to-clad interface appeared in good contacts. The fracture surface in the clad exhibited 

a zigzag pattern both on the back and front sides of the specimen. This unique pattern resulted 

from the combination of brittle fracture of hydrides and ductile failure of the metal matrix under 

tension. Delamination can be seen over the hydride layers near the fracture surface. 

In addition, spalling and cracking of the surface oxide layer can be seen on the back and front 

sides. 
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. A – delamination; B – cracking of oxide layer. 

4. TESTED AND UN-FAILED IN DYNAMIC TEST D4/ DM1/ 605D1C 

D4/ DM1 was based on segment 605D1C, whose estimated hydrogen and oxide layer thickness 

were 500 ppm and 40–70 μm. D4 survived 1.1×107 cycles of dynamic testing under ±7.62 N·m 

5 Hz 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 the untested fuel segment. Also visible is that the closely spaced 

fractures interacted and resulted in slender fragments aligned in the axial direction. One of the 

fragments was pulled out in the grinding process. 

The pellet-to-clad interface appeared to be in good contact. The circumferential hydrides are 

similar to those observed in the untested fuel segment. This is understandable because the two 

were neighboring segments in the same father fuel rod. 

In addition, the surface oxide layer seems to be 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 

5. FRACTURED IN DYNAMIC TEST D3/ DL3/ 605C10A 

Axial section 

D3/ DL3 was based on segment 605C10A with an estimated hydrogen contact of 550 ppm and 

an oxide layer of 70–100 μm. As mentioned earlier, D3 was tested under ±10.16 N·m 5 Hz with 

a lifetime of 106

 cycles. 

An optical image of the axial cross section is given in Figure F.4. The fracture surface was 

located at the left side, or 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. 

Part of the clad fracture surface near the front side was not preserved. The fracture surface near 

the back side of the specimen was highly serrated. The cavity visible near the fracture perhaps 

resulted from hydride cracking because it is coincident with the circumferential hydride layers. 

Spalling of the outside oxide layer was again seen near the fracture surface. 

(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. A – spalling of 

oxide layer; B – cavity aligning with circumferential hydride layer. 

A 

B
F-6 

Transverse section 

The set of SEM images presented in Figure F.5 (a–d) shows the different areas of the 

transverse cross section. The images corresponding to the 12 and 6 o’clock positions represent 

the maximum alternating stress areas of the rods subjected to potential crack initiation. The 

cladding was flattened apparently at the maximum stress areas after the fracture, and the origin 

of failure was not preserved. On the other hand, the areas near the 3 and 9 o’clock positions of 

the fracture surface revealed a clear laminate structure [Figure F.5 (e)] that 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 de-fueled D3: (a) 12 o’clock 

(reference mark), (b) 3 o’clock, (c) 6 o’clock, (d) 9 o’clock, and (e) enlarged area A. 

A 
F-7 

6. FRACTURED IN DYNAMIC TEST D5/ DM2/ 605D1B 

D5/ DM2 was based on segment 605D1B with an estimated hydrogen contact of 550 ppm and 

an oxide layer of 40–70 μm. As mentioned earlier, D5 was tested under ±9.14 N·m 5 Hz with a 

lifetime of 2.3×106

 cycles. 

An optical image of the axial cross section is given in Figure F.6. The fracture surface was 

located at the left side within a pellet. The primary fractures (whose size was of the same scale 

as the pellet) were similar to those revealed in the specimens examined above. The primary 

fractures were preexisting and may be a radiation effect because they were similar to those 

observed in untested fuel segments in terms of size and pattern. One primary fracture system 

normal to the axial direction of fuel corresponded to the fracture surface of the fuel and thus 

served as a main contributor to failure. 

Part of the clad fracture surface near the front side was flattened, resulting from impacts to the 

fracture surfaces before the machine reached the predetermined termination condition. Part of 

the fracture surface near the back side of the specimen was preserved and exhibited a highly 

serrated appearance. The delamination and connected cavities can be seen near the fracture, 

which may be attributed to hydride cracking because they are coincident with the circumferential 

hydride layers. 

At the same time, spalling of the outside oxide layer was seen near the fracture surface. 

 
F-8 

 

(a) (b)

 

(c) 

Figure F.6. (a) Axial cross section of D5/ 605D1B, (b) fracture surface near front side of 

rod, and (c) fracture surface near back side of rod and an enlarged area. A – spalling of 

oxide layer; B – delamination; C – connected cavities aligning with hydride layers. 

C

B

A 






NUREG/CR

-7198 Mechanical Fatigue Testing of High-Burnup Fuel for

Transportation Applications