Assessment of Current 

Test Methods for PostLOCA Cladding Behavior

Office of Nuclear Regulatory Research

NUREG/CR-7139

ANL-11/52
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Assessment of Current 

Test Methods for PostLOCA Cladding Behavior

Manuscript Completed: October 2011

Date Published: August 2012

Prepared by

M. C. Billone

Nuclear Engineering Division

Argonne National Laboratory

Argonne, IL 60439

H. H. Scott, NRC Project Manager

NRC Job Code V6199

Office of Nuclear Regulatory Research

NUREG/CR-7139

ANL-11/52

 

iii 

ABSTRACT 

Test methods to assess fuel-rod cladding behavior following a loss-of-coolant accident (LOCA) are 

compared and evaluated. For non-deformed cladding regions with uniform levels of hydrogen content 

and oxidation, the three-point bend test (3-PBT) is a very good test for ductility determination of asfabricated and pre-hydrided cladding subjected to LOCA oxidation and quench. For irradiated cladding, 

the ring compression test (RCT) has clear advantages over the 3-PBT because the required sample length 

is only about one-tenth of that needed for the 3-PBT. Overall, the RCT is the best test method for 

generating ductility data for assessing the effects of irradiation and hydrogen pickup on embrittlement 

oxidation threshold. However, neither 3-PBTs nor RCTs are useful for evaluating the performance of 

ballooned and ruptured cladding with significant axial gradients in cladding geometry, oxidation level, 

and hydrogen content, as well as circumferential gradients in wall thickness and oxidation level within the 

rupture region. Partially restrained axial contraction tests are useful for determining the fracture/nofracture boundary for ballooned, ruptured, oxidized, and quenched cladding as a function of hydrogen 

content and oxidation level. The four-point bend test (4-PBT) is best for determining three cladding 

performance metrics: maximum bending moment (measure of strength), failure energy (measure of 

toughness), and offset displacement (measure of plastic deformation). 

 

v 

FOREWORD 

The purpose of this document is to support the technical basis for the choice of test methods to assess 

fuel-rod cladding behavior following a loss-of-coolant accident (LOCA). For non-deformed cladding 

regions with uniform levels of hydrogen content and oxidation, and for irradiated cladding, the ring 

compression test (RCT) has been selected. The four-point bend test (4-PBT) has been selected for 

evaluating the performance of ballooned and ruptured regions of a fuel rod in LOCA analysis. The 

technical basis is founded on the results of the NRC’s LOCA research program, which was designed to 

measure the mechanical behavior of both non-deformed and ballooned and ruptured cladding following 

LOCA conditions. Integral LOCA tests were conducted at Argonne National Laboratory and Studsvik 

Laboratory in Sweden. The Japanese Atomic Energy Agency (JAEA) has also performed LOCA integral 

experiments. The results and observations of these experimental programs have been combined to 

develop considerations regarding the impact of oxidation and hydrogen content on the mechanical 

behavior of ballooned and ruptured cladding following LOCA conditions. 

Brian W. Sheron, Director 

Office of Nuclear Regulatory Research 

U.S. Nuclear Regulatory Commission 

 

vii 

CONTENTS 

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

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

CONTENTS................................................................................................................................................vii

FIGURES..................................................................................................................................................... ix

TABLES .....................................................................................................................................................xii

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

ACKNOWLEDGMENTS ........................................................................................................................xvii

ACRONYMS AND ABBREVIATIONS .................................................................................................. xix

SYMBOLS................................................................................................................................................. xxi

1 BACKGROUND ................................................................................................................................ 1

2 TEST METHODS FOR NON-DEFORMED CLADDING ............................................................... 3

2.1 Ring-compression tests (RCTs)................................................................................................. 3

2.2 Three-point-bend tests (3-PBTs) ............................................................................................... 7

2.3 Four-point-bend Tests (4-PBTs) ............................................................................................. 10

2.4 Axial tensile Tests ................................................................................................................... 18

2.5 Ring-stretch Tests.................................................................................................................... 19

2.6 Comparison of Test Methods for Non-deformed Cladding..................................................... 21

3 TEST METHODS FOR BALLOONED AND RUPTURED CLADDING ..................................... 25

3.1 Axial Tensile Tests.................................................................................................................. 28

3.2 Axial Restraint Tests ............................................................................................................... 28

3.3 Axial Bend Tests ..................................................................................................................... 32

3.4 Comparison of ANL 4-PBT and JAEA Axial-restraint Tests ................................................. 46

4 CONCLUSIONS............................................................................................................................... 53

REFERENCES ........................................................................................................................................... 55

 

ix 

FIGURES 

Figure 1. Schematic of ring-compression test sample and loading. The displacement (δ) rate is 

controlled and the response force (P) is measured. For two-sided-oxidized samples, 

failure occurs at one or both of the indicated locations. For one-sided-oxidized 

samples, failure may occur at locations ±90° from the loading direction. ................................ 4

Figure 2. Load-displacement curve from a RT ring compression test with an 8-mm-long 15×15 

Zry-4 cladding ring. The displacement rate was 2 mm/minute................................................ 4

Figure 3. Difference between offset and permanent strains as a function of oxidation level for asfabricated, pre-hydrided, and high-burnup cladding alloys oxidized at 1200°C and ringcompressed at 0.0333 mm/s. The dataset includes: 15×15 Zry-4; 17×17 Zry-4, M5 and 

ZIRLO; and 10×10 Zry-2 compressed at RT and 135°C to permanent strains of 1.0-

2.3%........................................................................................................................................... 6

Figure 4. Embrittlement oxidation limit (CP-ECR in %) as a function of metal hydrogen content 

for as-fabricated, pre-hydrided, and high-burnup cladding alloys oxidized at a peak 

temperature of 1200°C, quenched at 800°C, and ring-compressed at 135°C. Oxidation 

values above data points and/or trend lines resulted in brittle behavior. For hydrogen 

contents in the range of 540-600 wppm, peak cladding temperatures (1130-1180°C) 

occurred during the heating ramp at <1200°C. ......................................................................... 6

Figure 5. Schematic of the 3-PBT loading and support locations. The total length of the sample 

is L and the distance between supports is the span Ls. .............................................................. 8

Figure 6. CEA RCT and 3-PBT data generated at 135°C for pre-hydrided (≈600 wppm) 17×17 

Zry-4 cladding samples oxidized at 1200°C to about 3 to 6% measured ECR and 

cooled at either ≤0.4°C/s or 10°C/s to quench temperatures of 800°C or 600°C or 

cooled slowly with no quench: (a) offset displacement and (b) offset strain. ......................... 11

Figure 7. Schematic of four-point-bend test (4-PBT) loading. The bending moment [M = (P/2) 

a] is constant within the span Ls. The displacement (δ) rate is controlled and the force 

(P) is measured by the load cell............................................................................................... 12

Figure 8. Results for 4-PBT conducted at RT with an as-fabricated 15×15 Zry-4 sample: (a) total 

load (P) vs. displacement (δ) at loading points and (b) deformed shape of post-bend 

sample. Scale of 0.828:1 refers to original drawing size and is not relevant to Fig. 8b. ....... 13

Figure 9. Results of 4-PBT conducted at 135°C and 1 mm/s with as-fabricated ZIRLO cladding 

oxidized (one-sided) at 1200°C to 17% CP-ECR and quenched at 800°C (Test 

OCZL#24): (a) load-displacement curve showing expected low ductility (0.4 mm 

offset displacement) and (b) post-bend-test appearance of central region of sample 

showing severing at two locations on either side of sample mid-span marked in ink on 

this specimen. The severed cross section to the right in Fig. 9b appears to be the 

primary failure location. The severing of the cross section to the left may have 

occurred during impact of the sample with the bottom of the test fixture............................... 15
 

x 

Figure 10. Load-displacement curve for Ring 1B sectioned from post-LOCA-bend sample 

OCZL#24 oxidized (one-sided) at 1200°C to 17% CP-ECR. Sample had one tight 

crack at the 6 o’clock position and 1.2% permanent strain (ductile)....................................... 16

Figure 11. Load-displacement curve for Ring 1C sectioned from post-LOCA-bend sample 

OCZL#24 oxidized (one-sided) at 1200°C to 17% CP-ECR. Sample had one tight 

crack at the 12 o’clock position and 0.7% permanent strain (brittle)...................................... 16

Figure 12. Load-displacement curve for Ring 3B sectioned from post-LOCA-bend sample 

OCZL#24 oxidized (one-sided) at 1200°C to 17% CP-ECR. Sample had one tight 

crack at the 6 o’clock position and 0.8% permanent strain (brittle)........................................ 17

Figure 13. Load-displacement curve (a) and engineering stress-strain response (b) for 15×15 Zry4 sample subjected to axial tensile loading at room temperature. The slopes of both 

curves are about one-third of the expected slopes based on the sample gauge length, 

cross-sectional area, and Young’s modulus. ........................................................................... 20

Figure 14. Low magnification image of cross section through rupture mid-span for Test 

OCZL#18 sample with 43% mid-wall circumferential strain oxidized to 12% CP-ECR 

based on a pre-oxidation average wall thickness of 0.40 mm. ................................................ 25

Figure 15. Axial profiles of (a) cladding diametral strains and (b) hydrogen content for the 

OCZL#18 test sample with 43% maximum circumferential strain oxidized to 12% CPECR at T ≤ 1200°C, and quenched at 800°C. Do was 9.50 mm............................................. 26

Figure 16. Axial profiles of (a) cladding diametral strains and (b) hydrogen and oxygen content 

for the OCZL#19 test sample with 24% maximum circumferential strain oxidized to 

17% CP-ECR at T ≤ 1200°C and quenched at 800°C. Do was 9.50 mm. .............................. 27

Figure 17. Shape of the ANL OCZL#29 post-quench sample with maximum circumferential 

strain of 49% and maximum CP-ECR of 17%........................................................................ 29

Figure 18. JAEA LOCA apparatus (a) showing the gripping device used to restrain sample during 

cooling and (b) load vs. time curves for fully restrained samples and partially 

restrained (maximum loads of 735 N, 540 N, and 390 N) samples [17]................................. 30

Figure 19. Appearance [17] of JAEA A 3-1 sample with about 170-wppm pre-test hydrogen, 

oxidized at 1176°C to 29.3% BJ-ECR (23% CP-ECR), and partially constrained during 

cooling to a maximum axial tensile load of 540 N. Measured failure load was 498 N.......... 32

Figure 20. Maximum bending moment as a function of maximum oxidation level (CP-ECR) for 

post-LOCA samples subjected to 4-PBTs with the rupture region in tension for all but 

one test. 4-PBTs were performed at 135°C and 2 or 1 mm/s to 14-mm maximum 

displacement. One bend test was conducted at 30°C. ............................................................ 37

Figure 21. Maximum (for 0% CP-ECR) and failure (for ≥10% CP-ECR) energy as a function of 

oxidation level (CP-ECR) for post-LOCA samples subjected to 4-PBTs with the 

rupture region in tension for all tests but one. 4-PBTs were performed at 135°C and 2 

or 1 mm/s to 14-mm maximum displacement. One bend test was conducted at 30°C. ......... 38
 

xi 

Figure 22. Four-point-bend test (4-PBT) load-displacement curve for ramp-to-rupture sample 

OCZL#9. The as-fabricated ZIRLO sample ruptured at 875±15°C with a mid-wall 

rupture strain of 33%. The 4-PBT was conducted at 135°C and 2 mm/s displacement 

rate to 14-mm maximum displacement. .................................................................................. 40

Figure 23. Post-bend characterization of OCZL#19 sample that was subjected to bending at 

135°C with the rupture region under maximum tensile stress: (a) hydrogen-content 

profile; (b) measured values at failure locations; and (c) low-magnification image of 

severed cross section at 24 mm below rupture mid-span. ....................................................... 41

Figure 24. Post-bend characterization of the OCZL#18 sample oxidized to 12% CP-ECR, 

quenched, and subjected to bending at 135°C with the rupture region under maximum 

tensile stress: (a) hydrogen-content profile; (b) failure location; and (c) lowmagnification image of the severed cross section. .................................................................. 42

Figure 25. Metallographic images for the OCZL#18 severed cross section following oxidation to 

12% CP-ECR, quench, and bending at 135°C with the rupture region under maximum 

tension: (a) rupture tip with 0.14-mm average metal wall and (b) back side of cross 

section with 0.44-mm metal wall thickness. ........................................................................... 43

Figure 26. Post-bend characterization of OCZL#12 sample that was subjected to bending at 

135°C with the rupture region under maximum compressive stress: (a) hydrogencontent profile; (b) measured values at failure locations; and (c) low-magnification 

image of severed cross section at 40 mm below rupture mid-span. ........................................ 44

Figure 27. Post-bend results for OCZL#15 sample (18% CP-ECR and quench) with rupture node 

in compression (reverse bending), 135°C test temperature and 2 mm/s displacement 

rate: (a) load-displacement curve for 1st loading; (b) load-displacement curve for 2nd

loading; and (c) sample appearance after 2nd loading.............................................................. 45

Figure 28. 4-PBT results for Test Sample OCZL#17 oxidized to 13% CP-ECR, quenched, and 

tested at 135°C and 2 mm/s displacement rate: (a) load-displacement curve for 1st 4-

mm displacement (out of 10-mm total displacement); (b) post-test image of side view; 

and (c) post-test image of rupture view................................................................................... 47

Figure 29. Test Sample OCZL#29 oxidized to 17% CP-ECR, quenched, and subjected to 4-PBTs 

at 135°C and 1 mm/s displacement rate: (a) load-displacement curve for 1st loadingunloading sequence and (b) sample appearance after 1st loading-unloading........................... 48

Figure 30. Test Sample OCZL#29 oxidized to 17% CP-ECR, quenched, and subjected to 4-PBTs 

at 135°C and 1 mm/s displacement rate: (a) load-displacement curve for 2nd loadingunloading sequence and (b) sample appearance after 2nd loading-unloading.......................... 49
 

xii 

TABLES 

Table 1. Determination of normalization length (Ln) to convert offset displacement to maximum 

plastic strain for ANL 4-PBT samples with uniform geometry and material properties 

along the axial direction. Results are for as-fabricated Zry-4 cladding samples 

subjected to 4-PBTs (see Fig. 7) at RT.................................................................................... 14

Table 2. Results of characterization and ring-compression testing (at 135°C and 2 mm/minute 

displacement rate) for the post-bend OCZL#24 sample that was oxidized (one-sided) at 

1200°C to 17% CP-ECR and quenched at 800°C. .................................................................. 18

Table 3. Comparison of test methods for determination of ductile-to-brittle-transition oxidation 

level of post-LOCA samples at 135°C. Sample stiffness values are given qualitatively 

relative to machine stiffness values. 3-PBT and 4-PBT represent 3- and 4-point bend 

tests, respectively. PSU is Penn State University. EDC is expansion due to 

compression. R&D is research and development................................................................... 23

Table 4. Comparison of ANL and JAEA [13-16] LOCA integral sample and experimental 

parameters. AF is as-fabricated, PH is pre-hydrided, and Irr is irradiated. ............................. 29

Table 5. Summary of results for LOCA integral and post-LOCA bend tests with as-fabricated 

ZIRLO cladding. Reference LOCA test conditions were: 600- or 1200-psig fill 

pressure at 300°C, 5°C/s heating rate to 1200°C, 1200°C hold temperature, 3°C/s 

cooling rate to 800°C, and quench at 800°C. Reference conditions for 4-PBTs were 

135°C test temperature and 2 mm/s displacement rate to 14-mm maximum 

displacement. The displacement rate was lowered to 1 mm/s after the OCZL#21 bend 

test. .......................................................................................................................................... 34

Table 6. Dimensions, properties, and failure bending moment at severed location for test sample 

OCZL#19 with 17% maximum CP-ECR in the rupture node and 12% CP-ECR at the 

severed location....................................................................................................................... 50
 

xiii 

EXECUTIVE SUMMARY 

Current LOCA acceptance criteria in 10 CFR 50.46(b) limit the peak cladding temperature to 2200°F and 

the maximum oxidation level to 17% of the cladding wall thickness, assuming that all of the oxygen 

picked up by the cladding is in the form of ZrO2. These criteria are intended to assure that cladding 

retains some ductility during and following quench. According to the 1973 Commission: “Our selection 

of the 2200°F limit results primarily from our belief that retention of ductility in the zircaloy is the best 

guarantee of its remaining intact during the hypothetical LOCA.” 

In the absence of a credible analysis of loads, cladding stresses, and cladding strains for a degraded 

LOCA core, there are no absolute metrics to determine how much ductility or strength would be needed 

to “guarantee” that fuel-rod cladding would maintain its geometry during and following LOCA quench. 

It is also not clear what impact severing of some fuel rods into two pieces would have on core coolability. 

Fragmentation of brittle fuel-rod cladding with very low fracture toughness would be more detrimental to 

core coolability than severing of rods into two pieces. However, it is well recognized that materials that 

can deform plastically prior to failure – and hence are classified as ductile – are more resistant to 

fragmentation than brittle materials with low fracture toughness. Materials that retain ductility can relax 

secondary thermal and differential-expansion stresses through plastic flow. Therefore, the intent to 

maintain ductility is beneficial even without adequate knowledge of LOCA loads. 

Two test methods are currently being used as ductility screening tests for non-deformed cladding after 

exposure to LOCA oxidation and quench: ring compression tests (RCTs) and three-point bend tests (3-

PBTs). Two other test methods are currently being used to assess the performance of ballooned and 

ruptured cladding following oxidation: post-quench four-point bend tests (4-PBTs) and partial-to-full 

axial restraint tests during quench. Other test methods have been employed to determine plastic stressstrain properties and ductility for cladding materials under conditions relevant to normal operation and 

reactivity-initiated accidents: axial tensile tests, hoop tensile tests, and tests with combined axial and hoop 

tensile stresses. These test methods are discussed and their relative advantages and disadvantages are 

summarized for determining cladding performance following LOCA oxidation. 

Tests that require samples with end grips (e.g., axial tension) and/or machined gauge sections (e.g., axial 

and hoop tension) are not practical for LOCA-oxidized samples. Samples with high stiffness values 

relative to machine stiffness values are not desirable if one wants to compute failure energy, as well as 

ductility, from load-displacement curves. Tests that give load-displacement curves with a loading 

stiffness equal to the unloading stiffness are preferred if the simpler offset displacement method is used to 

determine ductility. Also, tests with smaller sample length are more practical for irradiated cladding 

because of limited availability and ease of defueling. Taking all these factors into account, the 3-PBT is 

the best choice for post-LOCA embrittlement determination of as-fabricated and pre-hydrided samples, 

which are not subjected to pre-oxidation deformation and are uniformly oxidized. Under the same 

conditions, the RCT is preferred for irradiated samples. However, it is desirable to have a single test 

method for as-fabricated, pre-hydrided, and irradiated cladding to eliminate systematic errors when 

comparing the behavior of these materials. For such comprehensive studies, the RCT is preferred. 

Long pressurized cladding samples subjected to a LOCA transient can balloon and rupture, resulting in 

local areas with significant axial variation in diameter, wall thickness, hydrogen content, and oxidation 

level, as well as large circumferential variations in wall thickness and oxidation level within cross 

sections containing the rupture region. RCTs with samples sectioned from the balloon region are not 

appropriate because of the steep variation in outer diameter. Axial tension and bend tests are more 

appropriate for studying ballooned specimens. 
 

xiv 

To produce ballooned specimens for axial bend tests and axial-restraint tensile tests, Argonne (ANL) and 

the Japanese Atomic Energy Agency (JAEA) have performed LOCA integral experiments with lengths of 

as-fabricated and pre-hydrided cladding. In addition, JAEA has conducted experiments with defueled 

cladding sectioned from irradiated fuel rods. 

Traditional axial tensile tests could be performed on post-LOCA integral samples, and pulling the sample 

to failure would give an accurate measure of failure load. However, if the load-displacement curve 

exhibited any offset displacement, it would not be clear if plastic displacement had occurred within the 

ballooned and ruptured region, just above and below this region where the temperature drops, or outside 

the middle region where the cladding is annealed, and lightly oxidized with negligible hydrogen pickup. 

The JAEA LOCA integral samples were not tested in the traditional way, but were fully or partially 

restrained from contracting during quench in the integral experiment. For samples that failed during 

quench, these tests generated data for axial failure load, failure location, and failure temperature. 

Although neither ductility nor failure energy could be determined, the JAEA test results do indicate that 

cladding can withstand high axial tensile loads without severing if the pre-oxidation hydrogen content and 

oxidation level are limited. 

ANL has used four-point-bend tests (4-PBTs) to determine post-LOCA sample failure location, maximum 

bending moment (measure of strength), failure energy (measure of toughness), and offset displacement 

(measure of plastic deformation). The 4-PBT has a major advantage over the 3-PBT for ballooned and 

ruptured samples because 4-PBT loading does not bias the failure location. For 3-PBT loading, the 

bending moment varies from maximum at the load application point to zero at the two supports. For 4-

PBT loading, the bending moment is constant between the load application points, which are spaced far 

enough apart to include the regions with non-uniform hydrogen content, oxidation level, and geometry. 

With analysis, the 4-PBT bending moment can be related to an equivalent axial tensile load. 

In the ANL 4-PBTs, samples with rupture strains <32% failed at an axial location between the rupture 

edge and the hydrogen peak where the cladding was fully brittle. With a few exceptions, samples with 

>40% rupture strain failed in the rupture node location where the crack propagated from very brittle 

cladding (rupture tips) through brittle-ductile-transition material and through high-ductility cladding (back 

side of the balloon). The maximum bending moment and the failure energy were found to be strong 

functions of the oxidation that occurred after rupture of as-fabricated cladding. Equations A and B, 

respectively, show best-fit linear correlations for the maximum bending moment (Mmax) and failure 

energy (Emax) as a function of calculated oxidation levels (CP-ECR, where CP refers to the use of the 

Cathcart-Pawel weight gain correlation) from 10% to 19%. 

Mmax = 13.92 – 1.073 (CP-ECR – 10%), N•m (A) 

Emax = 1.225 – 0.1236 (CP-ECR – 10%), J (B) 

Between 19% and 23% (maximum tested) CP-ECR, both Mmax and Emax are relatively small and constant. 

Equations A and B suggest that the oxidation level in the balloon region should be limited to enable the 

cladding to survive forces and moments during quench beyond the thermal stresses calculated for 

unrestrained cladding. At 17% CP-ECR, Eq. A predicts Mmax = 6.4 N•m, which corresponds to an axialrestraint failure load of about 2600 N for samples with <31% rupture strain. This value is comparable to 

the maximum loads measured by JAEA (1200-2400 N) for ballooned, ruptured, oxidized and quenched 

samples that survived conditions of full restraint from contraction during quench from 700°C to ≈100°C. 
 

xv 

Offset displacement can also be determined from 4-PBT load-displacement curves for ballooned, 

ruptured, oxidized, and quenched samples. However, for oxidation levels in the range of 10% to 18%, no 

offset strain was observed prior to the first significant crack severing >50% of the cross section 

containing the rupture flaw or all of the cross section within or outside the rupture region. These results 

apply to bend samples tested in the standard way with the brittle rupture tips subjected to axial tension 

and the ductile back region subjected to axial compression. Although it is clear that more than half the 

cladding cross section containing the rupture flaw had the capability of deforming plastically, most of the 

samples that failed in this region exhibited no offset displacement prior to cracking. Based on unflawed 

RCT samples, the ductile-to-brittle transition oxidation level is 19% for as-fabricated ZIRLO rings. 

Consistent with the RCT results, one 4-PBT sample at 18% CP-ECR subjected to “reverse” bending with 

the back region under axial tension exhibited a large offset displacement and did not even sever through 

14-mm displacement. It was concluded that offset displacement is not a useful metric for assessing postquench cladding performance of ballooned, ruptured and oxidized samples. 

Based on the ANL results, the 4-PBT maximum bending moment and failure energy can be used to 

determine the resistance to fracturing and fragmentation of cladding subjected to ballooning, rupture, 

oxidation, and quench. The ANL data suggest that the current limits of 17% CP-ECR and 1204°C for 

fresh and very low-burnup fuel cladding are adequate for protecting the cladding during quench not only 

from fragmentation but also from severing into two pieces under a wide range of loading conditions. 

None of the samples tested to failure in the ANL program fragmented or failed in a “low-toughness” 

mode. Test samples severed at one or two locations. By contrast, glass and ceramic rods with much 

lower toughness than the post-LOCA samples sever into as many as 5 to 10 pieces when subjected to 4-

PBT loading. 

 

xvii 

ACKNOWLEDGMENTS 

The author would like to express his appreciation to the Office of Nuclear Regulatory Research (RES) for 

support and management of this program. As Project Manager, Harold Scott provided programmatic and 

technical guidance. John Voglewede and Michelle Flanagan provided technical guidance with regard to 

data needs relevant to licensing issues. John also performed a detailed technical review of this report. 

The LOCA integral test data presented in this report were generated by ANL lead experimenters Yong 

Yan (now at Oak Ridge National Laboratory – ORNL) and Tatiana Burtseva. ANL technicians David 

McGann and Kevin Byrne are also acknowledged for their careful efforts in test setup, test conduct, and 

post-test measurements. Ralph Meyer provided valuable technical guidance for this test program and a 

detailed technical review of this report. He is also acknowledged for his in-depth assessment of ANL data 

and the comparison of these data to the international dataset for cladding behavior during and following a 

hypothetical loss-of-coolant accident. 

The author would like to express his appreciation to international colleagues Jean-Christophe Brachet and 

Valerie Vandenberghe of the Commissariat a l’Energie Atomique (CEA) at Saclay and Fumihisa Nagase 

of the Japanese Atomic Energy Agency (JAEA) for sharing details and insights regarding test 

methodologies, data, and embrittlement mechanisms. Members of the international ad hoc Mechanical 

Properties Experts Group are also acknowledged for sharing test methodologies and data for ring tension 

tests. The group was formed in October 2000 and has included members from Anatech, CEA, Electricitie 

de France (EdF), Electric Power Research Institute (EPRI), Institute for Radiological Protection and 

Nuclear Safety (IRSN), JAEA, ORNL, Pennsylvania State University (PSU), Russian Research Center’s 

Kurchatov Institute (RRC-KI), and Studsvik Nuclear AB. 

 

xix 

ACRONYMS AND ABBREVIATIONS 

3-PBT 3-point bend test 

4-PBT 4-point bend test 

AF As-fabricated 

ANL Argonne National Laboratory 

BJ Baker-Just weight gain correlation 

CEA Commissariat a l’Energie Atomique 

CFR Code of Federal Regulations 

cm Centimeters 

CP Cathcart-Pawel weight gain correlation 

°C Degrees Celsius 

ECR Equivalent cladding reacted 

EDC Expansion due to compression 

EdF Electricitie de France 

EPRI Electric Power Research Institute 

FEA Finite Element Analysis 

GWd/MTU Gigawatt-days per metric tonne of uranium 

ID Inner diameter 

IRSN Institute for Radiological Protection and Nuclear Safety 

J Joule, measure of energy 

JAEA Japanese Atomic Energy Agency 

kN Kilonewton 

LOCA Loss-of-coolant accident 

m Meter, measure of length 

mm Millimeter 

MPa Megapascal, measure of pressure and stress 

N Newton, measure of force 

NRC Nuclear Regulatory Commission 

OCZL Out-of-cell oxidation test for ZIRLO 

OD Outer diameter 

ORNL Oak Ridge National Laboratory 

PH Pre-hydrided 

ppm (weight) parts per million 

psig Pounds per square inch gauge pressure 

PSU Pennsylvania State University 

PWR Pressurized water reactor 

RCT Ring compression test 

RES Office of Nuclear Regulatory Research 

RRC-KI Russian Research Center’s Kurchatov Institute 

RT Room temperature 

s seconds 

wppm weight parts per million 

Zry-2 Zircaloy-2 

Zry-4 Zircaloy-4 

 

xxi 

SYMBOLS 

a Distance between applied load and support in 4-PBT 

A Cross-sectional area 

Aeq Equivalent cross-sectional area relative to the metal 

Ag Cross-sectional area for sample gauge section in axial tension test 

Am Cross-sectional area of metal 

Aox Cross-sectional area of oxide 

(Aox)i Cross-sectional area of inner-surface oxide 

(Aox)o Cross-sectional area of outer-surface oxide 

dap Permanent displacement at loading positions in 4-PBTs 

Di Inner diameter 

Dmi Inner diameter of metal 

Dmo Outer diameter of metal 

Do Outer diameter 

dp Permanent displacement of RCT ring in loading direction 

E Young’s modulus 

Eg Young’s modulus for gauge material in axial tension test 

Em Young’s modulus for metal 

Emax Maximum applied energy in 4-PBT; equivalent to failure energy for failed samples 

Eox Young’s modulus for oxide 

h wall thickness 

hm Wall thickness of metal 

I Area moment of inertia for bending 

Ieq Equivalent area moment of inertia relative to the metal 

Im Area moment of inertia for metal 

(Iox)i Area moment of inertia for inner-surface oxide 

(Iox)o Area moment of inertia for outer-surface oxide 

K Stiffness; linearized loading slope of load-displacement curve 

Kg Stiffness of gauge section in response to axial load 

Km Measured stiffness for RCTs and 4-PBTs; machine stiffness for axial tension tests 

L Length 

Lg Gauge length for axial tension test sample 

Ln Normalization length to convert bending displacement to axial strain 

Ls Distance between supports in 3-PBT and distance between applied loads in 4-PBT 

M Bending moment 

Mmax Maximum bending moment 

P Load measured by load cell 

Pmax Maximum load 

Rmo Outer radius of metal 

T Temperature 

TR Rupture temperature 

δ Applied displacement 

ΔC/Cm Circumferential strain at the mid-wall location 

δe Elastic displacement 

δg Elastic displacement of gauge section in axial tension test 

δm Elastic displacement of machine, including all components outside gauge length 

(δox)i Thickness of inner-surface oxide 
 

xxii 

(δox)o Thickness of outer surface oxide 

δp Offset displacement based on load-displacement curve 

εe Elastic strain 

εp Plastic strain 

ρp Post-bend-test radius of curvature of the sample 
 

1 

1 BACKGROUND

Current LOCA acceptance criteria in 10 CFR 50.46(b) limit the peak cladding temperature to 2200°F and 

the maximum oxidation level to 17% of the cladding wall thickness, assuming that all of the oxygen 

picked up by the cladding is in the form of ZrO2. These criteria are intended to assure that cladding 

retains some ductility during and following quench. According to the 1973 Commission: “Our selection 

of the 2200°F limit results primarily from our belief that retention of ductility in the zircaloy is the best 

guarantee of its remaining intact during the hypothetical LOCA.” 

In the absence of a credible analysis of loads, cladding stresses, and cladding strains for a degraded 

LOCA core, there are no absolute metrics to determine how much ductility or strength would be needed 

to “guarantee” that fuel-rod cladding would maintain its geometry during and following LOCA quench. 

It is also not clear what impact severing of some fuel rods into two pieces would have on core coolability. 

Fragmentation of very brittle fuel rod cladding would be more detrimental to core coolability than 

severing of rods into two pieces. It is well recognized that materials that can deform plastically – and 

hence are classified as ductile – are more resistant to fragmentation than brittle materials. In particular, 

materials that retain ductility can relax secondary thermal and differential-expansion stresses through 

plastic flow. Therefore, the intent to maintain ductility is beneficial even without adequate knowledge of 

LOCA loads. 

The current LOCA acceptance criteria were based on ring-compression tests to assess oxidation and 

temperature limits at which embrittlement occurs. Ring-compression loading induces circumferential 

bending stresses in post-LOCA cladding samples. Under the load and above the support, the hoop 

stresses across the cladding wall vary from maximum tensile stress (inner surface) to maximum 

compressive (outer surface). At ±90° from the loading direction, the hoop stress distribution is reversed 

(i.e., compressive inner- and tensile outer-surface hoop stresses). However, ring-compression loading is 

not prototypical of other anticipated loads on the cladding during quench: axial stresses due to bending, 

axial stresses due to restricted thermal contraction of the cladding, and possible impact loading in the 

balloon region. As such, the ring-compression test should be viewed as a ductility screening test. Axial 

bending, axial tension, and hoop tension tests could also be used as ductility screening tests. These test 

methods are reviewed and assessed for post-LOCA mechanical tests of non-deformed and deformed-andruptured cladding samples following oxidation and quench. 

 

3 

2 TEST METHODS FOR NON-DEFORMED CLADDING 

Several types of tests have been used as ductility screening tests to assess oxidation and temperature 

limits at which embrittlement occurs for cladding samples that are not deformed after exposure to LOCA 

conditions. Ring compression tests are described in Section 2.1. Three-point and four-point axial bend 

tests are discussed in Sections 2.2 and 2.3, respectively. Axial tensile and ring-stretch tests are discussed 

in Sections 2.4 and 2.5, respectively. The relative advantages and disadvantages of these test methods are 

summarized in Section 2.6. 

2.1 Ring-compression tests (RCTs) 

Argonne National Laboratory (ANL) has exposed as-fabricated, pre-hydrided, and high-burnup cladding 

samples to two-sided oxidation and quench. Short (about 8-mm-long) rings sectioned from these samples 

have been subjected to ring-compression tests to determine the ductile-to-brittle transition oxidation level 

as a function of hydrogen content. Results of such tests have been reported [1,2]. With the limited 

amount of high-burnup cladding available for such tests, there is a practical advantage in using ring 

compression testing in that a large number of tests can be performed with a relatively small inventory of 

cladding. Another practical advantage comes from lower gamma and beta-gamma worker dose 

accumulation per test for short samples. 

Figure 1 shows a schematic of the ring cross section, loading, and displacement. The displacement (δ) is 

applied at a constant rate and the force (P) is measured by the load cell. All samples are compressed to 

failure, which may manifest itself as a single crack or multiple cracks through the sample wall and along 

the sample length. Ductility is determined from the permanent change (dp) in ring diameter after 

unloading and/or by the offset displacement (δp) method. For two-sided-oxidized samples, cracking 

through the wall and along the length of the sample is equally probable at inner-cladding surface locations 

under the load (12 o’clock position) and above the support plate (6 o’clock position). The maximum 

bending moment and tensile hoop stress occur at these two locations. The bending moment and tensile 

hoop stress at positions ±90° (3 and 9 o’clock) from the loading axis are about 40% less within the elastic 

deformation regime. Also, the RCT stresses at ±90° from the loading axis vary from compressive at the 

inner cladding surface to tensile at the outer cladding surface. However, failure data are seldom generated 

at these locations because of the lower stress. 

The ring in Fig. 1 acts like a spring in the elastic deformation regime with spring constant K such that P = 

K δ. For as-fabricated cladding alloys, the spring constant K can be expressed in terms of the Young’s 

modulus E, length L, outer diameter Do, and wall thickness h. For 8-mm-long, as-fabricated cladding 

rings tested at ANL, calculated stiffness values (Kc) ranged from 1440 N/mm (15×15 Zry-4 with Do = 

10.77 mm and h = 0.76 mm) at room temperature (RT) to 800 N/mm (17×17 Zry-4 with Do = 9.50 mm 

and h = 0.57 mm) at 135°C. These values were calculated using the isotropic and recrystallized-annealed 

correlation for E given in MATPRO [3] for Zry-2 and Zry-4. Measured values (Km) determined from 

load-displacement curves were generally within ±10% of Kc for Zry-4 (see Fig. 2) and other cladding 

alloys (e.g., Zry-2, ZIRLO, and M5). The deviation was reduced to ±5% when MATPRO anisotropic 

values for Eθ were used to determine Kc. MATPRO also has a correction for cold-work that would 

further reduce E (by about 5% for 25% CW). However, the as-fabricated cold-work levels for Zry-4 and 

ZIRLO alloys supplied to ANL by vendors were not provided with the material description. The 

observation that the measured loading slope agrees with the calculated loading slope is important as it 

indicates that the RCT load-displacement curve represents the response of the cladding ring independent 

of the compliance of the test machines used at ANL. 
 

4 

Figure 1. Schematic of ring-compression test sample and loading. The displacement (δ) rate is 

controlled and the response force (P) is measured. For two-sided-oxidized samples, 

failure occurs at one or both of the indicated locations. For one-sided-oxidized 

samples, failure may occur at locations ±90° from the loading direction. 

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 0.5 1 1.5 2 2.5

Displacement (mm)

Load (kN)

Initial Loading Slope = 1.34 kN/mm

Offset Displacement = 1.32 mm

Permanent Displacement = 1.19 mm

Figure 2. Load-displacement curve from a RT ring compression test with an 8-mm-long 15×15 

Zry-4 cladding ring. The displacement rate was 2 mm/minute. 
 

5 

Rings with ductility will exhibit nonlinear load vs. displacement response beyond the elastic deformation 

regime. As shown in Fig. 2, the offset displacement (δp) was 1.32 mm while the permanent displacement 

(dp) was 1.19 mm. The inherent error in δp comes from the assumption that the unloading slope (i.e., 

stiffness at the end of loading) is the same as the loading slope. The permanent displacement shown in 

Fig. 2 was in excellent agreement with the permanent change (pre-test minus post-test) in diameter 

measured directly from the sample. This subject is discussed in detail in Ref. 4. For post-LOCA rings 

tested to failure, the unloading slope of a ring just prior to failure cannot be determined. Thus, the loading 

slope must be used to determine δp. This approach is adequate to determine the ductile-to-brittle 

transition as long as the inherent error in δp is quantified. 

It is common practice to normalize plastic displacement by the initial diameter (Do) to determine relative 

plastic displacement. This is referred to as “strain” although it does not represent an average or maximum 

strain in the material. For example, the maximum elastic strain (εe) for an as-fabricated 17×17 Zry-4 ring 

is only about a third (0.29 δe/Do) of the normalized elastic displacement. However, considering the 

complexity of oxidized and quenched rings, the determination of relative plastic displacement is accurate 

enough for a ductility screening test. 

Because rings are compressed to failure in these tests, the measured permanent change in diameter also 

contains an error due to the unloading of a cracked ring. Based on experience and data trends, this error is 

<1% permanent strain. Therefore, rings exhibiting >1% permanent strain are classified as ductile, while 

rings that fail at <1% permanent strain are classified as brittle. For non-oxidized, as-fabricated cladding 

rings, the error between offset and permanent strain is negligible for small offset strain values (e.g., <3%). 

However, for oxidized-and-quenched cladding, the error increases with oxidation level. This deviation is 

shown in Fig. 3, which gives offset minus permanent strain (δp/Do - dp/Do) as a function of oxidation level 

(in % CP-ECR). By setting the permanent strain (dp/Do) to the 1% ductility limit and adding it to the 1-σ

upper bound in Fig. 3, the offset-strain ductility criterion can be expressed as: 

δp/Do ≥ 1.41% + 0.1082 CP-ECR (1) 

For new RCT data sets, the measured average offset strain (rounded to the nearest 0.1%) at a particular 

hydrogen content and oxidation level should be greater than or equal to the right-hand side of Eq. 1 

(rounded to nearest 0.1%) for the material to be classified as ductile. 

Ring compression tests have been conducted at many laboratories. However, ring support and loading 

surfaces (e.g., flat vs. curved), test conditions (e.g. test temperature and displacement rate), test 

procedures (e.g., stopping test after first significant load drop), and data interpretation (e.g., ductility 

based on total displacement vs. offset displacement vs. permanent displacement) vary considerably. Also, 

most laboratories did not search for the ductile-to-brittle-transition oxidation level as a function of 

hydrogen content. Thus, a one-to-one comparison could not be made between the dataset generated by 

ANL (see Fig. 4) and data generated at other laboratories. 

Ductility and determination of ductile-to-brittle transition oxidation level have been the focus of the 

material presented on ring-compression testing. However, it has also been demonstrated that the loaddisplacement results are independent of machine compliance. Thus, one could determine failure energy 

(normalized to length) of the ring from the test results. 

At their research center in Saclay, the French Atomic Energy Commission (CEA) performed ringcompression tests with as-fabricated and pre-hydrided Zry-4 and M5 cladding samples oxidized at 

1200°C, as well as at lower and higher temperatures. RCT results for 1100°C and 1200°C oxidation 
 

6 

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

0 2 4 6 8 10 12 14 16 18 20 22

CP-Predicted ECR (%)

Offset Minus Permanent Strain (%)

Zircaloy

ZIRLO

M5

Best Fit

1-σ Upper Bound

Figure 3. Difference between offset and permanent strains as a function of oxidation level for asfabricated, pre-hydrided, and high-burnup cladding alloys oxidized at 1200°C and ringcompressed at 0.0333 mm/s. The dataset includes: 15×15 Zry-4; 17×17 Zry-4, M5 and 

ZIRLO; and 10×10 Zry-2 compressed at RT and 135°C to permanent strains of 1.0-2.3%. 

0

2

4

6

8

10

12

14

16

18

20

0 100 200 300 400 500 600 700 800

Hydrogen Content (wppm)

Embrittlement Oxidation Limit (%) )

17x17 Zry-4

HBR-type (old) 15x15 Zry-4

17x17 ZIRLO

17x17 M5

15x15 Zry-4

10x10 Zry-2

Linear Fit for ≤ 400-wppm H

Linear Fit for > 400-wppm H

63 GWd/t

70 GWd/t 70 GWd/t

64 GWd/t

Figure 4. Embrittlement oxidation limit (CP-ECR in %) as a function of metal hydrogen content for 

as-fabricated, pre-hydrided, and high-burnup cladding alloys oxidized at a peak 

temperature of 1200°C, quenched at 800°C, and ring-compressed at 135°C. Oxidation 

values above data points and/or trend lines resulted in brittle behavior. For hydrogen 

contents in the range of 540-600 wppm, peak cladding temperatures (1130-1180°C) 

occurred during the heating ramp at <1200°C. 
 

7 

temperatures are presented in Refs. 5 and 6. Differences between CEA and ANL oxidation-quench tests 

include: heats (i.e., fabrication lots) of Zry-4 and M5 samples; 150-mm vs. 25-mm sample length; onesided vs. two-sided oxidation, fast (20°C/s) vs. slow (2-3°C/s) temperature ramps within 100°C of 

oxidation hold temperature; and direct quench at oxidation temperature vs. about 13°C/s cooling rate from 

oxidation temperature to 800°C quench temperature. However, CEA also ran some tests with cooling 

rates of 0.4°C/s and 10°C/s from 1200°C to 800°C with quench temperatures of 800°C, 700°C, and 600°C 

[5,6]. ANL conducted a few tests with quench temperatures of 700°C and 600°C [1]. With regard to 

RCTs, differences between CEA and ANL include: 10-mm vs. 8-mm sample length; curved vs. straight 

support plate; 0.5 vs. 2 mm/minute displacement rate; and fixed displacement resulting in multiple cracks 

vs. stopping tests at the first significant load drop. The CEA approach yielded valuable data for offset 

strain, but data for permanent strain could not be determined for CEA test samples because of multiple 

cracks or wide single cracks due to continued displacement beyond the point at which the first throughwall crack developed. 

For non-oxidized, as-fabricated 17×17 Zry-4, the RCT stiffness for CEA rings is calculated to be 1000 

N/mm at 135°C. As CEA has not published load-displacement curves for such material, it cannot be 

determined if the slope of the load-displacement curves agrees with the calculated ring stiffness. Such 

agreement does not affect the determination of offset strain. However, if machine and fixture compliance 

is low enough, the measured loading slope could be lower than calculated for the ring, and the failure 

energy determined from the load-displacement curve could be higher than the failure energy for the ring. 

The stiffness of the CEA machine loading and support fixtures would have to be >10 kN/mm for the 

measured loading stiffness to be within 10% of the calculated ring loading stiffness. 

In general, ANL and CEA RCT results were consistent in terms of ductile vs. brittle behavior for cases in 

which the pre-oxidation hydrogen content, oxidation level, cooling rate, and quench temperature were 

similar (see Fig. 7 in Ref. 6). 

2.2 Three-point-bend tests (3-PBTs) 

Figure 5 is a schematic of the three-point bend test (3-PBT). The applied load at the mid-span distance 

between supports results in a bending moment distribution that is maximum (P•Ls/4) at the applied 

loading position (Ls/2) and decreases linearly to zero from the loading position to the two support 

positions. The top and bottom surfaces in Fig. 5 are subjected to axial compressive and tensile stresses, 

respectively. Thus, the maximum axial tensile stress occurs at the axial location at which the load is 

applied. With the load applied to the top surface, the maximum axial tensile stress occurs at the bottom 

surface. 

CEA conducted both ring-compression tests (RCTs) and three-point-bend tests (3-PBTs) using asfabricated and pre-hydrided cladding samples oxidized (outer surface only) at 1200°C and either 

quenched at 1200°C, slow cooled to lower temperatures and quenched, or cooled at 10°C/s to lower 

temperatures and quenched [5,6]. For the CEA bend samples, the span Ls between supports was 80 mm 

and the total sample length was 90 mm. The controlled displacement rate at the loading position was 0.5 

mm/minute, the same as used in the CEA RCTs. As cladding inventory and dose rate are not issues for 

these materials, the use of long samples for 3-PBTs is not a disadvantage. Also, the 3-PBT has an 

advantage over the RCT inasmuch as loading and unloading stiffness values in bend tests should be 

essentially the same whereas they are different for RCT samples. Another advantage of the 3-PBT is that 
 

8 

Figure 5. Schematic of the 3-PBT loading and support locations. The total length of the sample is 

L and the distance between supports is the span Ls. 

failure is characterized by a single crack through the cross section at the loading position. In the absence 

of multiple cracks that form in RCT samples, it is straightforward to determine offset-displacement values 

from load-displacement curves. One disadvantage of the 3-PBT is that permanent displacement cannot be 

determined from pre- and post-test measurements for samples that sever into two pieces. 

Because axial stresses are likely to occur during LOCA quench, the 3-PBT loading, which induces axial 

tensile stresses on the convex surface (bottom surface in Fig. 5) of the bent sample, is LOCA-relevant. 

However, CEA concluded that no significant difference exists between embrittlement oxidation levels 

determined from ring tests (hoop bending stress) and axial bend tests (axial bending stress). This implies 

that if the prior-beta layer is brittle in the hoop direction it will also be brittle in the axial direction [5,6]. 

However, it is difficult to compare displacement and strain data sets directly because the lengthnormalization factor to convert displacement to strain has not been determined for the 3-PBT sample. 

Also, unlike the RCT, the offset displacement and strain uncertainty levels have not been determined for 

the 3-PBT. The displacement uncertainty is estimated to be low but non-zero. The uncertainty in offsetdisplacement measurement and the length-normalization factor are estimated in this section for CEA 3-

PBT results to allow a direct comparison of these results with CEA RCT results. 

CEA RCT and 3-PBT results in Refs. 5 and 6 are for pre-hydrided (≈600-wppm H) 17×17 Zry-4 oxidized 

to about 3% to 6% measured ECR and ring-compression tested at 135°C. The displacement rate for both 

test types was 0.5 mm/minute. There is no a priori reason why the same displacement rate would give 

the same material strain rate for these tests. However, for as-fabricated, non-oxidized cladding, the 

maximum elastic strain rate for the CEA 3-PBT sample is about 60% of the strain rate for the RCT 

sample at the 12 and 6 o’clock positions and about equal to the maximum RCT elastic strain rate at the 3 

and 9 o’clock positions. Thus, CEA’s use of the same displacement rate for both types of tests is 

justified. 

The calculated stiffness for the 80-mm-span CEA bend sample loaded at mid-span is 1300 N/mm at 

135°C for as-fabricated, non-oxidized 17×17 Zry-4 with 9.50-mm OD and 0.57-mm wall thickness. This 

stiffness is 30% higher than the stiffness for RCT samples sectioned from the same material. Stiffness 

values for oxidized samples may be higher due to the higher Young’s modulus for the oxide as compared 

to the metal. On the other hand, the addition of 600-wppm of hydrogen would tend to lower the Young’s 

modulus of the cladding metal. 
 

9 

The 3-PBT loading biases the failure location to coincide with the loading location, at which the bending 

moment, axial tensile stress, and axial tensile strain all reach their maximum value. The bending moment 

decreases linearly from the mid-span to the supports. For uniform cladding geometry, hydrogen content 

and oxidation level, failure should occur at the loading location. 

An advantage of the 3-PBT is that the displacement is measured at the probable failure location. Thus, 

the load-displacement curve can be used to determine directly the offset displacement at the failure 

location. For the axial bend sample, the unloading slope is essentially equal to the loading slope. Thus, 

unlike the ring-compression results, there is no significant error in the determination of offset 

displacement using the loading slope to “unload” the sample prior to severing of the cross section. 

Unlike the ring-compression test, there is no obvious choice of a dimension to normalize the 3-PBT offset 

displacement to calculate relative plastic displacement or strain. Intuitively, it should be the span length 

or some fraction of the span length. CEA could perform a finite-element analysis (FEA) to calculate 

maximum tensile plastic strain as a function of displacement for as-fabricated, non-oxidized cladding, as 

well as benchmark 3-PBTs with such cladding. In the absence of such an analysis, ANL has performed 

an elastic analysis to determine the appropriate length (Ln) to convert applied displacement to strain 

equivalent to the maximum axial tensile strain. This length was determined to be: Ln = Ls

2

/(12 Do), where 

Ls is the cladding sample length between supports and Do is the cladding sample OD. For the CEA 

sample geometry, Ln = 56 mm, which is considerably larger than the 9.50 mm (sample OD) that CEA 

used to normalize 3-PBT offset displacements to determine offset strain (Fig. 5b in Ref. 6). Based on 

ANL FEA and benchmark data for the 4-PBT (see next section), the normalization length for offset 

displacement is less than the length for elastic displacement. As 56 mm is closer to the half span (40 mm) 

than the full span (80 mm) for the CEA tests, a preliminary recommendation is that Ln = (Ls/2) = 40 mm 

be used to convert CEA 3-PBT offset displacement to offset strain. However, ANL and CEA agree that 

there is no clean and simple way to compare RCT and 3-PBT results in terms of offset strain because the 

3-PBT normalization length changes with displacement and oxidation level. 

In terms of offset displacement uncertainty for bend tests, ANL has conducted 4-PBT loading-unloadingreloading experiments to determine loading and unloading slopes as a function of plastic displacement. 

Within the range of interest, there was no significant difference. Thus, for homogeneous material, one 

would expect no inherent error in the determination of offset displacement. However for nonhomogeneous material, such as oxidized cladding, there may be an uncertainty of about 0.1 mm in the 

determination of offset displacement. In Ref. 6 (see Fig. 6), CEA has published 3-PBT load-displacement 

curves for 600-wppm-H Zry-4 oxidized to about 3% to 6% measured ECR, cooled at different rates, and 

quenched at different temperatures. The results range from near-zero (3% ECR oxidized and quenched at 

1200°C) offset displacement to values as high as 2.7 mm (6% ECR slow cooled to 600°C and quenched). 

Based on fractography using scanning electron microscopy (SEM) images, it is clear that the sample with 

near-zero 3-PBT offset displacement is brittle. Yet the 3-PBT offset displacement was measured by ANL 

to be 0.06 mm based on the load displacement curve given in Fig. 6 of Ref. 6. Thus, it would be prudent 

to place an uncertainty limit of about 0.1 mm on 3-PBT offset displacement determination and assessment 

of plastic deformation. The corresponding RCT limit for samples oxidized to 3-6% ECR is 0.19 mm (see 

Fig. 3). These limits are shown in Fig. 6a for CEA RCT offset displacement vs. 3-PBT offset 

displacement (note: values were calculated from data plotted in Figs. 5b and 7b of Ref. 6). Eight of the 

data points plotted showed excellent agreement with respect to ductile vs. brittle behavior determined by 

both test methods: 5 ductile data points and 3 brittle data points. Two of the data points were brittle 

based on the RCT criterion (<0.19 mm) and ductile based on the 3-PBT criterion (≥0.1 mm). The 

agreement is quite good given the different stress states (hoop vs. axial) induced by the two test methods. 
 

10 

The offset displacement data shown in Fig. 6a were normalized by 9.50 mm for RCT and 40 mm for 3-

PBT to generate the offset strain data shown in Fig. 6b. The conclusions are the same with regard to the 

very good agreement between the two test methods in distinguishing ductile vs. brittle behavior. 

CEA also plotted “spent energy” or failure energy in Figs. 5d and 7d of Ref. 6. Depending on the 

effective stiffness of the CEA 3-PBT fixtures and tensile machine, these energy values may overestimate 

the energy stored in the samples prior to failure. With an ANL-calculated sample stiffness of 1300 

N/mm, the machine stiffness would have to be >12.6 kN/mm in order for the energy determined from the 

linear part of the 3-PBT load-displacement curve to be within 10% of the elastic energy stored in the 

sample. 

2.3 Four-point-bend Tests (4-PBTs) 

The four-point-bend test (4-PBT, see Fig. 7) has advantages over the 3-PBT if there is axial variation in 

cladding geometry, hydrogen content, and/or oxidation level. A uniform bending moment is applied 

along the span Ls between the two loading points. The sample will fail at its weakest location or 

locations. If the sample is very uniform and brittle (e.g., glass or ceramic rod), then simultaneous failure 

at multiple locations is likely to occur. 

For the ANL loading and support fixtures, Ls = 150 mm, a = 50 mm, and the distance between supports is 

250 mm. The overhang regions beyond the supports are each 25 mm, but this length does not affect the 

loading. Thus, the uniform bending moment (in N•m) is given in terms of the measured load (in N) by M 

= (0.025 m) P. This bending moment results in maximum tensile stresses and strains at the bottom 

surface of the sample shown in Fig. 7, which would be convex during bending. The stresses and strains 

transition from maximum axial tension at the bottom surface to zero at the cross-section neutral axis to 

maximum axial compression at the top surface. The maximum stresses and strains are uniform within the 

span Ls for uniform geometry and material properties. 

The 4-PBT sample is much more flexible than the RCT sample or the CEA 3-PBT sample. Benchmark 

tests have been conducted at RT and 135°C with 15×15 Zry-4 (Do = 10.77 mm and h = 0.76 mm) and 

17×17 M5 (Do = 9.50 mm and h = 0.57 mm). Calculated stiffness values varied from 246 N/mm (15×15 

Zry-4 at RT) to 121 N/mm (17×17 M5 at 135°C). Measured values (248±2 N/mm and 121 ±1 N/m, 

respectively) were in excellent agreement with calculated values. FEA calculations were also performed 

for 15×15 and 17×17 Zry-4 based on ANL-measured tensile properties for 15×15 Zry-4. The FEAcalculated load-displacement curves for the 15×15 Zry-4 benchmark tests at RT were in excellent 

agreement with the measured load-displacement curves. 

Based on a comparison of the maximum elastic strain vs. displacement (δ) for the 4-PBT sample and the 

normalized RCT displacement (δ/Do), the displacement rate for the 4-PBT was increased from 2 

mm/minute to 2 mm/s. This was later reduced to 1 mm/s based on a direct comparison between 

calculated maximum elastic strain values for the RCT sample (εθ = 0.0305 mm-1 δ) and the 4-PBT sample 

(εz = 0.00104 mm-1 δ). 

For one of the RT benchmark tests, 15×15 Zry-4 was subjected to a displacement of 10 mm at the loading 

points, which was large enough to induce plastic flow. The load-displacement curve for this test is shown 

in Fig. 8a. Although loads were not measured during unloading, the subsequent reloading stiffness was 

measured and found to be within 1% of the original loading stiffness. From this it was concluded that: 

the unloading stiffness was also the same and there would be no inherent error in determining 4-PBT 

offset displacement as there is for the RCT. 
 

11 

0.0

0.5

1.0

1.5

2.0

0.0 0.5 1.0 1.5 2.0 2.5 3.0

3-PBT Offset Displacement (mm)

RCT Offset Displacement (mm)

RCT Brittle Regime

3-PBT Brittle Regime

(a) Offset displacement 

0

1

2

3

4

5

6

7

8

9

10

11

12

13

14

012345678

3-PBT Offset Strain (%)

RCT Offset Strain (%)

RCT Brittle Regime

3-PBT Brittle Regime

(b) Offset strain 

Figure 6. CEA RCT and 3-PBT data generated at 135°C for pre-hydrided (≈600 wppm) 17×17 Zry-4 

cladding samples oxidized at 1200°C to about 3 to 6% measured ECR and cooled at 

either ≤0.4°C/s or 10°C/s to quench temperatures of 800°C or 600°C or cooled slowly 

with no quench: (a) offset displacement and (b) offset strain. 
 

12 

Figure 7. Schematic of four-point-bend test (4-PBT) loading. The bending moment [M = (P/2) a] is 

constant within the span Ls. The displacement (δ) rate is controlled and the force (P) is 

measured by the load cell. 

From Fig. 8a, the offset displacement was determined to be δp = 1.9 mm at the sample loading points. 

The post-test deformed sample is shown in Fig. 8b. Accurate determination of the permanent 

displacement at the loading points requires enlargement of Fig. 8b. The approximate value of permanent 

displacement at the loading points is dap = 2.4 mm based on a reduced-size (0.828:1) photograph and dap = 

2.1 mm based on an enlarged (1.40:1) photograph. The 2.1-mm permanent displacement is in good 

agreement with the 1.9-mm offset displacement. It should be noted that permanent displacement cannot 

be measured accurately for oxidized samples tested to failure (i.e., severed into 2 pieces). 

The permanent plastic strain can, in principle, be determined from the radius of curvature of the deformed 

sample. From beam theory, the relationship between the maximum plastic strain (εp) and the radius of 

curvature (ρp) of the neutral axis is given by: 

εp = (Do/2)/ρp (2) 

Based on an enlarged view of Fig. 8b, ρp = 1560 mm and εp = 0.35%. However, the measurement of 

radius of curvature is tedious and cannot be done accurately for an oxidized sample tested to failure. 

A third approach to determining plastic strain is to compare the offset displacement at several 

displacements beyond the elastic regime to FEA-calculated maximum plastic strains. The calculation was 

performed for both 15×15 and 17×17 Zry-4 geometries. The experimental results in Fig. 8a were used to 

determine offset displacement vs. total displacement for 15×15 Zry-4. Analytical load-displacement 

results were used for 17×17 Zry-4. The results are given in Table 1 for offset displacement, calculated 

maximum plastic strain, and the normalization length (Ln) needed to convert offset displacement to plastic 

strain. For small plastic strains in the range of 0.l6% to 0.30%, the approximate value for Ln is 250 mm, 

which is the distance between supports in Fig. 7. 

From a practical perspective, it is sufficient to use offset displacement as the metric for plastic 

deformation for the ANL 4-PBTs. Mathematically, δp > 0 implies plastic flow and ductility. Taking into 
 

13 

(a) Load-displacement curve 

(b) Post-bend-test sample 

Figure 8. Results for 4-PBT conducted at RT with an as-fabricated 15×15 Zry-4 sample: (a) total 

load (P) vs. displacement (δ) at loading points and (b) deformed shape of post-bend 

sample. Scale of 0.828:1 refers to original drawing size and is not relevant to Fig. 8b.
 

14 

account measurement errors, it would be prudent to set a lower limit on offset displacement in the 

determination of ductile vs. brittle behavior: δp ≥ 0.1 mm implies ductile behavior. In addition to plastic 

deformation, material strength (as measured by the maximum bending moment) and failure energy (area 

under load-displacement curve) can be determined. 

Table 1. Determination of normalization length (Ln) to convert offset displacement to maximum 

plastic strain for ANL 4-PBT samples with uniform geometry and material properties 

along the axial direction. Results are for as-fabricated Zry-4 cladding samples 

subjected to 4-PBTs (see Fig. 7) at RT. 

Cladding 

Geometry 

Offset Displacement, 

mm 

Maximum Tensile 

Plastic Strain, % 

Normalization 

Length, mm 

17×17 0.31 0.16 194 

15×15 0.50 0.20 250 

15×15 0.90 0.30 300 

17×17 1.56 0.38 410 

15×15 1.90 0.50 380 

Additional benchmark tests were conducted with 17×17 M5 cladding subjected to 4-PBTs at RT and 

135°C. The displacement rate was 2 mm/s for these tests. Measured stiffness values were 131±1 N/mm 

at RT and 121±1 N/mm at 135°C. The measured values were in excellent agreement with the calculated 

values of 128 N/mm and 121 N/mm, respectively. 

Four-point-bend tests were also performed with oxidized and quenched samples. One such test was 

performed with a specimen (OCZL#24) that had been oxidized on the outer surface at 1200°C to 17% 

CP-ECR and quenched at 800°C. Similar ZIRLO samples that were oxidized (two-sided) at 1200°C to 

17% CP-ECR had been subjected to RCTs and exhibited low ductility. Thus, the 4-PBT should also 

demonstrate low ductility if the two test methods yield consistent results with regard to the ductile-tobrittle transition oxidation level. 

Figure 9a shows the load-displacement curve for this 4-PBT conducted at 135°C and a displacement rate 

of 1 mm/s. The measured offset displacement was 0.4 mm. When compared with the 0.1 mm threshold 

offset displacement discussed above, this 4-PBT result indicates low ductility. The results also 

demonstrate that the 4-PBT is sensitive enough to distinguish ductile from brittle behavior even near the 

transition oxidation level. Figure 9b shows the post-test sample, which failed at two locations (±38 mm 

from axial mid-span location). 

The middle region in Fig. 9b was sectioned into one sample (4-mm-long) for metallographic examination 

and three samples (each 8-mm-long) for ring compression testing. As the sample had already experienced 

some axial plastic flow, the expectation was that compressed rings would exhibit near brittle behavior. 

Load-displacement curves for these three rings are shown in Figs. 10-12. The results of the ringcompression tests are summarized in Table 2. Two of the post-bend rings did exhibit brittle behavior and 

on average the rings were brittle: 1.5±0.4% offset and 0.9±0.3% permanent strains. 
 

15 

It is worth noting that the average of the measured loading stiffness values for the compressed rings was 

1000 N/mm. For non-oxidized, as-fabricated ZIRLO, compressed at 135°C, the measured stiffness values 

were 830 N/mm [4]. Thus, the presence of the outer-surface oxide layer increased the stiffness by about 

20%. This is consistent with the uniform oxidation of the rings in this region of the test sample. 

0

100

200

300

400

500

600

700

800

900

1000

1100

0123456789

Displacement, mm

Load, N

Mmax = 24.1 N•m

Emax = 4.15 J

δp = 0.4 mm

Km = 0.12 kN/mm

(a) Load-displacement curve 

(b) Post-test sample 

Figure 9. Results of 4-PBT conducted at 135°C and 1 mm/s with as-fabricated ZIRLO cladding 

oxidized (one-sided) at 1200°C to 17% CP-ECR and quenched at 800°C (Test OCZL#24): 

(a) load-displacement curve showing expected low ductility (0.4 mm offset 

displacement) and (b) post-bend-test appearance of central region of sample showing 

severing at two locations on either side of sample mid-span marked in ink on this 

specimen. The severed cross section to the right in Fig. 9b appears to be the primary 

failure location. The severing of the cross section to the left may have occurred during 

impact of the sample with the bottom of the test fixture. 
 

16 

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2

Displacement (mm)

Load (kN)

0.18 mm

Test Number OCZL#24

IPS 111B3B1B

Offset Strain = 1.9%

Permanent Strain = 1.2%

Km = 0.97 kN/mm

Figure 10. Load-displacement curve for Ring 1B sectioned from post-LOCA-bend sample OCZL#24 

oxidized (one-sided) at 1200°C to 17% CP-ECR. Sample had one tight crack at the 6 

o’clock position and 1.2% permanent strain (ductile). 

0

0.05

0.1

0.15

0.2

0.25

0.3

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2

Displacement (mm)

Load (kN)

0.11 mm

Test Number OCZL#24

IPS 111B3B1C

Offset Strain = 1.2%

Permanent Strain = 0.7%

Km = 1.01 kN/mm

Figure 11. Load-displacement curve for Ring 1C sectioned from post-LOCA-bend sample OCZL#24 

oxidized (one-sided) at 1200°C to 17% CP-ECR. Sample had one tight crack at the 12 

o’clock position and 0.7% permanent strain (brittle). 
 

17 

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2

Displacement (mm)

Load (kN)

0.13 mm

Test Number OCZL#24

IPS 111B3B3B

Offset Strain = 1.4%

Permanent Strain = 0.8%

Km = 1.03 kN/mm

Figure 12. Load-displacement curve for Ring 3B sectioned from post-LOCA-bend sample OCZL#24 

oxidized (one-sided) at 1200°C to 17% CP-ECR. Sample had one tight crack at the 6 

o’clock position and 0.8% permanent strain (brittle). 

The measured stiffness shown in Fig. 9a is 122 N/mm. This is the same stiffness as was calculated for 

non-oxidized 17×17 Zry-4 and measured for non-oxidized 17×17 M5. The oxidized ZIRLO described in 

Table 2 has a calculated bending stiffness of 156 N/mm assuming MATPRO [3] elastic constants for the 

oxide (14.8×104

 MPa) and the metal (8.65×104 MPa) at 135°C. The measured stiffness is about 20% 

lower than the calculated stiffness. Some of this difference may be due to decreasing oxide-layer 

thickness from about 50 mm from the mid-span to the support locations. 

It is interesting that the compressed rings developed a single crack at the 6-o’clock (2 rings) or 12-o’clock 

(1 ring) position. The crack initiated at the prior-beta layer inner surface and propagated through the 

alpha and oxide layers. Although CEA has conducted many ring compression tests using one-sidedoxidized samples, it has never been clear at which location the crack initiates because CEA does not stop 

tests at the first significant load drop. It has been assumed that the first crack initiates at the outer surface 

of the 3 or 9 o’clock position. This is clearly not the case for rings approaching the embrittlement 

oxidation level. 

The 4-PBT requires longer samples than the 3-PBT, and this makes it more difficult to pre-hydride and 

oxidize the 4-PBT samples uniformly in a LOCA-type transient. In comparing the ANL 4-PBT sample to 

the CEA 3-PBT for non-ballooned specimens, the only advantage of the ANL 4-PBT is low sample 

stiffness, which allows for accurate determination of failure energy. 
 

18 

Table 2. Results of characterization and ring-compression testing (at 135°C and 2 mm/minute 

displacement rate) for the post-bend OCZL#24 sample that was oxidized (one-sided) at 

1200°C to 17% CP-ECR and quenched at 800°C. 

Parameters Value 

Oxidized Cladding OD, mm 9.73±0.02 

OD Oxide Layer Thickness, μm 103±9 

Average Metal Wall Thickness, mm 0.50 

Average Metal ID, mm 8.52 

RCT Offset Displacement, mm 0.18 

0.11 

0.13

0.14±0.04 

RCT Offset Strain, % 1.9 

1.2 

1.4

1.5±0.4 

RCT Permanent Displacement, mm 0.11 

0.07 

0.08

0.09±0.02 

RCT Permanent Strain, % 1.2 

0.7 

0.8

0.9±0.3 

2.4 Axial tensile Tests 

Axial tensile tests have not been used to determine ductile-to-brittle transitions. Special grips would need 

to be designed to avoid failure at the grip locations. Also, machining is required to define a gauge section 

over which stress and strain can be determined based on load-displacement data. Metal gauge sections 

can be machined precisely by an electro-discharge machine (EDM) after outer-surface-oxide removal. 

This would be very difficult to do for post-LOCA samples because the oxygen-stabilized alpha layer is 

more brittle than the oxide layer. Oxide layer removal by mechanical means would result in damage to 

this alpha layer. 

Axial tensile samples have a much higher stiffness (K = EA/L) than ring-compression and axial-bend 

samples. “A” is the gauge cross-sectional area and L is the gauge length. For 17×17 cladding tested at 

135°C, sample stiffness values range from 27 kN/mm (51-mm-long ASTM sample with no gauge) to 9.9 

kN/mm (ANL double gauge, 25.4-mm-long, 2.5-mm-width per gauge [7]). The long load train acts like a 

spring in series with the sample. High sample stiffness results in high load-train loads, small elastic 

strains along the load train, and load-train displacements larger than sample elastic displacements. The 

load cell accurately measures the sample load, but the measured displacement includes both load-train 

and sample displacements. Let Kg = EgAg/Lg be the stiffness of the sample gauge section and Km be the 
 

19 

effective machine stiffness (includes all loaded components outside of gauge region). The total 

displacement measured by the Instron actuator is the sum of the machine displacement (δm = P/Km) and 

the sample gauge displacement (δg = P/Kg). 

δ = δm + δg = P/Km + P/Kg = (1/Km + 1/Kg) P (3) 

Thus, the combined stiffness (K) can be found by solving Eq. 3 to get: 

K = (Kg/Km + 1)-1 Kg (4) 

Figure 13 shows (a) the load-displacement data and (b) the corresponding engineering stress-strain 

response for an as-fabricated HBR-type 15×15 Zry-4 ANL tensile sample tested at RT. The engineering 

strain is determined by the ratio of the measured total displacement (δ) and the gauge length (Lg = 25.4 

mm). The stress is simply the load (P) divided by the cross-sectional area (Ag = 3.8 mm2

). For this 

sample, Kg is calculated to be 13.8 kN/mm. As shown in Fig. 13a, the load-displacement curve has a 

linearized loading slope of K = 4.65 kN/mm. Solving Eq. 4 for Km gives, Km = 6.9 kN/mm. Thus, for 

this particular tensile testing, the machine stiffness (includes grips and non-gauge part of sample) is about 

one-half of the sample gauge stiffness. The effect of these two “springs” in series gives a response 

stiffness about one-third of the sample gauge stiffness. This is easily seen in the Fig. 13b stress-strain 

plot. The slope of the elastic stress-strain curve is expected to be Young’s modulus (E = 92,400 MPa), 

but the linearized slope of the stress-strain curve is only a third of that value. However, Fig. 13b also 

shows that the offset strain method can be used to determine the yield strength (≈600 MPa) at 0.2% offset 

strain. In principle, this methodology could be used to determine ductile-to-brittle transition (e.g., offset 

strains < 0.2% imply brittle behavior), but tensile tests with gauged samples are not practical for postLOCA embrittlement studies. 

2.5 Ring-stretch Tests 

Hoop tensile tests have been performed to determine plastic stress-strain relationships for as-fabricated, 

pre-hydrided, and irradiated cladding alloys. The hoop tensile test requires machined gauge sections and 

extensive finite-element analysis to determine the plastic stress-strain relationship in the hoop direction 

[7]. The length of the sample is small (≈3 mm with 1-mm gauge), but the stiffness is relatively high (1.6 

kN/mm for 17×17 Zry-4 at 135°C). Radial loading is induced by pulling on a pair of D-shaped mandrels 

or on a pair of modified D-shaped mandrels with a dog-bone insert to minimize bending. This test is not 

appropriate as a ductility screening test for post-LOCA samples because of pre-test machining 

requirements and difficulty in determining low-ductility values. 

The Penn State University (PSU) plane-strain test has been used as a ductility test conducted with 

samples subjected to a biaxial stress state [7]. Sample preparation requires oxide removal, machined 

notches at the axial ends of the sample, and micro-indents on the cladding metal outer surface. Samples 

are longer than hoop-tensile samples (≈13 mm with ≈7 mm between notches). The stiffness of these 

samples is high (>10 kN/mm). Radial loading is induced by pulling on a pair of D-shaped mandrels. 

Plastic displacement is determined from permanent change in length between micro-indents. This test is 

not appropriate as a ductility screening test for oxidized post-LOCA samples for the same reasons that the 

uniaxial ring-stretch samples are not appropriate. 
 

20 

0

0.5

1

1.5

2

2.5

3

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

Displacement (mm)

Load (kN)

K = 4.65 kN/mm

(a) Load-displacement curve 

0

100

200

300

400

500

600

700

800

0 0.01 0.02 0.03 0.04 0.05

Engineering Strain (mm/mm)

Engineering Stress (MPa)

Slope = 30,200 MPa

YS

(b) Stress-strain curve 

Figure 13. Load-displacement curve (a) and engineering stress-strain response (b) for 15×15 Zry-4 

sample subjected to axial tensile loading at room temperature. The slopes of both 

curves are about one-third of the expected slopes based on the sample gauge length, 

cross-sectional area, and Young’s modulus. 
 

21 

Another way of performing ring-stretch tests is the so-called expansion-due-to-compression (EDC) test 

method, for which pre-test machining of gauge sections is not needed [8,9]. Radial loading is induced by 

axial compression of a plug inserted into the sample and extending about 0.5 mm above the sample. Load 

is applied to the plug by the top ram or top-bottom rams. The hoop tensile stress induced in the sample is 

much more uniform than what can be achieved in the hoop-tension ring-stretch test. ORNL has used 

EDC to measure the RT hoop tensile properties of non-irradiated and irradiated Zry-4 [8]. Studsvik has 

used the same technique to measure the hoop tensile properties of irradiated Zry-2 in the temperature 

range of 25-340°C [9]. Typical cladding lengths for these tests are 7 mm (ORNL) and 20 mm (Studsvik). 

Stiffness values are >10 kN/mm for these samples. The expansion vs. time of the sample outer surface is 

measured directly by a pair of laser probes. The plug (polyurethane) is chosen to have a low hardness 

value. Expansion of the sample is limited by the plug displacement (≈1 mm for ORNL and 2.9 mm for 

Studsvik). Compression of the plug beyond the top edge of the sample may cause sample “barreling” and 

axial stresses (ORNL). Cladding diameter change is limited to ≈4% strain (ORNL) and ≈20% (Studsvik). 

The measured ram load and ram displacement represent the response of the ram-plug-sample system. 

Load-displacement curves for the ram cannot be used directly to determine sample offset displacements 

and strains. Determination of the hoop stress in the cladding as a function of the measured diameter or 

change in diameter is non-trivial. In addition to uncertainties in the elastic-plastic behavior of the plug, 

friction forces must be accounted for. Data plots are usually in the form of measured cladding radial 

displacement vs. ram displacement. A factor is derived by independent analyses or experiments to relate 

the measured ram load to the cladding hoop stress. ORNL determined this factor by using data from axial 

tensile tests for non-irradiated Zry-4 and assuming isotropy in the hoop and axial directions: σθ = 0.53 

P/(h L). The same 0.53 factor was used for irradiated cladding. Although the EDC test has certain 

advantages relative to the other ring-stretch tests, it has not been used to measure the ductility of hightemperature, steam-oxidized cladding samples. This would require considerable development work and 

careful error analysis. It is anticipated that the relative error would increase as the cladding ductility 

decreased. 

2.6 Comparison of Test Methods for Non-deformed Cladding 

Table 3 shows a summary comparison of test methods for ductility and embrittlement determination of 

non-deformed oxidized cladding samples. Tests that require samples with end grips (e.g., axial tension) 

and/or machined gauge sections (e.g., axial and hoop tension) are not practical for LOCA-oxidized 

samples. Samples with high stiffness values relative to machine stiffness values are not desirable if one 

wants to compute failure energy, as well as ductility, from load-displacement curves. Tests that give 

load-displacement curves with a loading stiffness equal to the unloading stiffness are better if the offset 

displacement method is used to determine ductility. Also, tests with smaller samples are better for 

irradiated cladding. Taking all of these factors into account, the 3-PBT is the best choice for post-LOCA 

embrittlement determination of as-fabricated and pre-hydrided samples whereas the ring-compression test 

is the best choice for irradiated samples. However, it is sometimes desirable to have a single test method 

for as-fabricated, pre-hydrided, and irradiated cladding to eliminate systematic errors when comparing the 

behavior of irradiated and non-irradiated cladding. For such a comprehensive study, the RCT is 

preferred. 

For the 3-PBT, the CEA span length of 80 mm has been chosen for calculations and comparison 

purposes. CEA 3-PBT results suggest that this span is adequate for determining offset displacement (δp) 

and embrittlement threshold, along with the criterion that δp ≥ 0.1 mm implies ductility. However, if the 

CEA 3-PBT were to be used to determine failure energy, the span length could be increased from 80 mm 

to 110 mm (reduces the stiffness from 1300 N/mm to 500 N/mm, as stiffness varies inversely with the 

cube of the span length) and a relatively stiff machine (Km > 5000 N/mm) could be used. This would 
 

22 

result in better agreement between the calculated and measured (within 10%) loading stiffness and failure 

energy values. Although not currently used in ductility screening tests, the failure energy (measure of 

toughness) is a useful parameter to determine. The same machine could be used for RCT samples, which 

would result in better agreement between the measured and calculated slopes of the loading curve. At 

135°C, the predicted loading slope for as-fabricated 17×17 Zry-4 is 1000 N/mm. The ANL-measured 

RCT loading slope for CEA samples (600-wppm-H 17×17 Zry-4) was 820 N/mm following oxidation 

and quench. CEA could conduct benchmark RCTs and 3-PBTs with as-fabricated 17×17 Zry-4 to 

determine if there is a significant difference between calculated and measured loading slopes. The results 

of such simple tests would provide useful information regarding the stiffness of the machine and fixtures 

used to conduct these tests. 
23 

Table 3. Comparison of test methods for determination of ductile-to-brittle-transition oxidation level of post-LOCA samples at 135°C. 

Sample stiffness values are given qualitatively relative to machine stiffness values. 3-PBT and 4-PBT represent 3- and 4-

point bend tests, respectively. PSU is Penn State University. EDC is expansion due to compression. R&D = research and development. 

Test Type 

Sample 

Length, 

mm 

Grips 

Required 

Gauge 

Sections 

Required 

Sample 

Stiffness 

Determination of Offset 

Displacement or Strain 

From Load-Displacement 

Curve 

Ranking 

for 

Non-irradiated

Cladding 

Ranking 

for 

Irradiated 

Cladding 

ANL-CEA 

Ring 

Comp. Test 8-10 No No Low Good High Very High 

CEA 3-PBT 90 No No Low Very Good Very High High 

ANL 4-PBT 300 No No Very Low Very Good Low Low 

ANL Axial 

Tension Test 76 Yes Yes High OK for ≥0.2% strain Low Low 

ANL Hoop 

Tension Test 3 No Yes Moderate Difficult Low Low 

PSU Plane 

Strain Test 13 No Yes High Difficult Low Low 

EDC Test 

ORNL 

Studsvik 

7 

20 

No No 

No No 

High High 

Difficult Difficult 

Needs R&D Needs R&D 

Needs R&D 

Needs R&D 

 

25 

3 TEST METHODS FOR BALLOONED AND RUPTURED CLADDING 

Long, pressurized cladding samples subjected to a LOCA transient can balloon and rupture. Such 

samples will contain local areas that exhibit considerable axial variation in diameter, wall thickness, 

hydrogen content, and oxidation level. A substantial circumferential variation in wall thickness and 

oxidation level will also occur in cross sections containing a rupture opening (see Refs. 1, 10-12). Figure 

14 shows the variation in wall thickness at the mid-span rupture cross section for the ANL Test OCZL#18 

sample which had 43% maximum circumferential strain. The maximum oxidation level (12% CP-ECR 

based on average wall thickness prior to oxidation) occurred in this cross section for which the average 

metal wall thickness prior to oxidation was 0.40 mm. Consistent with current licensing guidelines in 10 

CFR 50.46(b), all CP-ECR values reported in this section for the rupture-region cross section are based 

on average wall thickness prior to oxidation. 

Figure 14. Low magnification image of cross section through rupture mid-span for Test OCZL#18 

sample with 43% mid-wall circumferential strain oxidized to 12% CP-ECR based on a 

pre-oxidation average wall thickness of 0.40 mm. 

Figure 15 shows the axial variation for the OCZL#18 sample in (a) cladding diametral strain at two 

orientations and (b) cladding hydrogen content. Figure 16 shows axial profiles for (a) diametral strains 

and (b) hydrogen and oxygen content for the OCZL#19 test sample with 24% maximum circumferential 

strain and 17% oxidation level. While the maximum oxidation level occurred at the rupture-opening midspan, the maximum hydrogen pickup occurred outside the rupture opening close to the necks of the 

balloon. These tests were conducted using as-fabricated 17×17 ZIRLO cladding filled with zirconia 

pellets, pressurized to 1200 psig (OCZL#18) or 600 psig (OCZL#19) at 300°C, ramped at 5°C/s to 

1200°C in steam, held at 1200°C, cooled at 3°C/s to 800°C, and rapidly cooled via water quench from 

800°C to 100°C. 
 

26 

0

10

20

30

40

50

60

70

80

0 1 2 3 4 5 6 7 8 9 10 11 12 13

Distance from the bottom (in.)

Strain (

ΔD/Do in %)

90° from Burst

0° from Burst

Top

41%

66%

(a) Axial profile of diametral strains 

(b) Axial profile of hydrogen content 

Figure 15. Axial profiles of (a) cladding diametral strains and (b) hydrogen content for the 

OCZL#18 test sample with 43% maximum circumferential strain oxidized to 12% CPECR at T ≤ 1200°C, and quenched at 800°C. Do was 9.50 mm.
 

27 

0

10

20

30

40

50

60

0 1 2 3 4 5 6 7 8 9 10 11 12 13

Distance from the bottom (in.)

Strain (

ΔD/Do in %)

90° from Burst

0° from Burst

Top

17%

38%

(a) Axial profile of diametral strains 

(b) Axial profile of hydrogen and oxygen content 

Figure 16. Axial profiles of (a) cladding diametral strains and (b) hydrogen and oxygen content for 

the OCZL#19 test sample with 24% maximum circumferential strain oxidized to 17% CPECR at T ≤ 1200°C and quenched at 800°C. Do was 9.50 mm. 
 

28 

Ring-compression tests with samples sectioned from the balloon region are not appropriate tests because 

of the steep variation in outer diameter. Loading would be highly localized at the point of maximum 

outer diameter, and it would spread axially as the cladding deformed or cracked. This would result in an 

increase in the load-displacement slope prior to plastic deformation and/or cracking. Also, results would 

be highly dependent on orientation of the loading location relative to the rupture opening. The 

determination of “ductility” prior to cracking along the length of the sample would be highly uncertain. 

Therefore, axial tensile and bend tests are more appropriate for studying ballooned specimens, and such 

testing with ballooned specimens is relatively recent. 

To produce ballooned specimens for axial tensile and bend tests, ANL and the Japanese Atomic Energy 

Agency (JAEA) [13-14] at their Tokai research center have performed LOCA integral experiments with 

lengths of as-fabricated and pre-hydrided cladding alloys. In addition, JAEA has conducted experiments 

with defueled cladding sectioned from irradiated rods [15-16]. 

JAEA integral experiments differ from ANL’s in terms of sample length, internal pressure and volume, 

heating rate, hold temperature, cooling rate and quench temperature. Experimental parameters are 

compared in Table 4. (Not included in Table 4 are JAEA high-burnup Zry-2, ZIRLO, MDA and NDA 

samples [16].) However, the most significant difference between ANL and JAEA testing has to do with 

mechanical test methods. ANL has performed post-quench 4-PBTs at 135°C, while JAEA has performed 

full and partial axial-restraint tests during cooling from the hold temperature to <100°C. These 

mechanical test methods are reviewed in Sections 3.2 and 3.3. 

3.1 Axial Tensile Tests 

Traditional axial tensile tests could be performed on post-LOCA integral samples, and pulling the sample 

to failure would give an accurate measure of failure load, as every cross section is subjected to the same 

tensile load. The samples are long, have welded end caps for gripping, and are highly unlikely to fail at 

the grip locations because these are well outside the uniform temperature zone. The oxidation level and 

hydrogen content at these end locations are both low relative to the middle of the sample. However, 

during the pressurized phase of the transient, the sample would experience some bending (see Fig. 17 for 

ANL 17×17 ZIRLO sample). Subjecting this sample to a displacement-controlled axial tension test 

would result in a low initial loading slope to straighten out the sample followed by a steeper slope due to 

combined machine stiffness and sample stiffness (see Eqs. 3 and 4). However, if the load-displacement 

curve exhibited any offset displacement, it would not be clear if plastic displacement had occurred within 

the ruptured region of the balloon, just above and/or below this region where the oxidation temperature 

decreases, or outside the middle region where the cladding is annealed and lightly oxidized with 

negligible hydrogen pickup. Therefore, the only meaningful data that could be derived from traditional 

axial tensile tests are failure loads. 

3.2 Axial Restraint Tests 

The JAEA LOCA integral samples were not tested in the traditional way, but were restrained from axial 

contraction during quench in the integral experiment. The purpose of these tests was to determine: the 

fracture/no-fracture boundary as a function of hydrogen content and oxidation level and the fracture load 

for tests in which fracture occurred. Both full- and partial-axial-restraint tests have been conducted 

during cooling from the oxidation temperature to <100°C. Figure 18a shows the gripping mechanism 

used to restrain the cladding. Figure 18b shows typical axial load vs. time results for fully restrained and 
 

29 

partially restrained cases that were limited to maximum axial loads of 735 N, 540 N (reference JAEA 

case), and 390 N. 

Table 4. Comparison of ANL and JAEA [13-16] LOCA integral sample and experimental 

parameters. AF is as-fabricated, PH is pre-hydrided, and Irr is irradiated. 

Parameters ANL JAEA 

AF PH AF PH Irr 

17×17 Cladding ZIRLO ZIRLO Zry-4 Zry-4 Zry-4 

Cladding OD, mm 9.50 9.50 9.42 9.42 --- 

Wall Thick., mm 0.57 0.57 0.51 0.51 --- 

Sample Length, mm 

(minus end caps) 

295 295 570 570 180 

Hydrogen Content, wppm ≈10 200-600 ≈10 100-1400 150±40 

Pellets zirconia zirconia alumina alumina alumina 

Pellet Stack 

Length, mm 

280 280 550 550 170 

Gas Volume, cm3 10 10 4.4 4.4 2.4 

Internal Pressure, MPa 

(Gauge) 

4.14 

8.28 

@ 300°C 

4.14 

8.28 

@ 300°C 

5 @ RT 5 @ RT 5 @ RT 

Heating Rate, °C/s 5 5 10 10 10 

Rupture T, °C 843±6 

750±7 

742 

700±30 

--- 656-880 786±25 

Rupture Strain, % 22±3 

46±4 

58 

60±10 

35 (avg.) 11- 40 16±8 

Hold Temp., °C 1200 1200 947-1257 947-1257 1030-1178 

Cooling Rate to 

Quench Temp., °C/s 

3 3 20 to 900°C 

5 to 700°C 

20 to 900°C 

5 to 700°C 

20 to 900°C 

5 to 700°C 

Quench Temp., °C 800 800 700 700 700 

Figure 17. Shape of the ANL OCZL#29 post-quench sample with maximum circumferential strain 

of 49% and maximum CP-ECR of 17%. 
 

30 

(a) JAEA LOCA apparatus 

(b) Axial tensile loads vs. time for several constraint conditions 

Figure 18. JAEA LOCA apparatus (a) showing the gripping device used to restrain sample during 

cooling and (b) load vs. time curves for fully restrained samples and partially restrained 

(maximum loads of 735 N, 540 N, and 390 N) samples [17].
31 

The measured load is relatively low in these tests during cooling from the oxidation hold temperature to 

the 700°C quench temperature due to the transition from a bent sample to a straight sample and due to the 

low yield strength and high plasticity of the cladding metal at elevated temperatures. However, the load 

builds up rapidly during quench cooling because the metal yield strength increases with decreasing 

temperature and the metal does not have time to flow plastically to relax the thermal stresses. 

For the fully restrained cases [13-14], the maximum load was 1200 to 2400 N for samples that survived 

quench without failure. For samples that failed during quench, the failure loads were 200 to 1700 N. 

Most of the failures occurred in the rupture node, which is not expected to pick up much (<30 wppm) 

hydrogen during the oxidation phase. These are important results, however, because regardless of the 

degree of restraint, the maximum axial load due to differential contraction was found to be ≤2400 N. 

JAEA conducted 118 tests with pre-hydrided cladding, which included variation of the hydrogen content, 

oxidation hold temperature, oxidation hold time, and degree of axial constraint. Forty-seven samples 

failed during quench with 42 of the failures occurring in the rupture node. 

JAEA also conducted 6 tests using irradiated Zry-4 cladding from PWR fuel rods with ≤44 GWd/MTU 

burnup, ≤25-μm corrosion layer, and estimated hydrogen contents in the range of 150±40 wppm. Based 

on several analyses of the strength of grid spacers, JAEA argued that axial restraint loads had to be < 

1000 N. They adopted 540 N for their partially restrained tests using irradiated cladding samples. 

For irradiated cladding, the sample length was reduced from 580 mm to 190 mm. Notice from Table 4 

that the RT gas volume was reduced from 4.4 cm3

 to 2.4 cm3. Rupture strains were smaller (16±8%) 

than measured for longer, pre-hydrided test samples of comparable hydrogen content. The smaller gas 

volume and the shorter (≈40 mm) uniform temperature zone may have contributed to the lower rupture 

strains. Oxidation temperatures were 1171±9°C for five of the tests and 1030°C for one of the tests. 

Partially restrained tensile tests were conducted during cooling with the maximum axial load set at 540 N. 

Two of the samples oxidized at about 1170°C failed at axial loads of 498 N (170-wppm H and 23% CPECR) and 385 N (120 wppm hydrogen and 20% CP-ECR). 

For the sample that fractured at 498 N, post-test imaging (see Fig. 19) indicated that failure occurred near 

the edge of the rupture opening. JAEA interpreted this as a rupture-node failure. However, the results are 

open to interpretation. In particular, the high-hydrogen concentrations near the fracture are indicative of 

regions outside the rupture opening, which have high hydrogen pickup from inner-surface oxidation. 

Consistent with JAEA interpretation, the crack may have initiated at the edge of the rupture opening, but 

it clearly propagated into the high hydrogen zone outside the rupture opening. 

For non-irradiated cladding, the hydrogen peaks were 30 to 50 mm from the rupture mid-span. Based on 

the results shown in Fig. 19, one hydrogen peak appears to occur at <10 mm from the rupture mid-span 

for irradiated Zry-4. It is not clear whether this was due to a difference between irradiated and prehydrided cladding or differences in sample length (190 mm vs. 580 mm), uniform temperature zone (40 

mm vs. 100 mm), and/or rupture strain (14% vs. 20-40%). 
32 

Figure 19. Appearance [17] of JAEA A 3-1 sample with about 170-wppm pre-test hydrogen, 

oxidized at 1176°C to 29.3% BJ-ECR (23% CP-ECR), and partially constrained during 

cooling to a maximum axial tensile load of 540 N. Measured failure load was 498 N. 

3.3 Axial Bend Tests 

There are three locations (see Fig. 16b) in as-fabricated LOCA integral samples that are vulnerable to 

failure (i.e., severing of the cross-section) in response to axial tensile stresses induced by axial bending: 

the rupture location with the thinnest average wall thickness and therefore the highest oxidation level; and 

two locations on either side of the rupture location with higher hydrogen content and intermediate 

oxidation levels. These two locations are between the edges of the rupture opening and the hydrogen 

peaks. Hydrogen enhancement of oxygen embrittlement is responsible for failure at the intermediate 

locations. For as-fabricated cladding, the hydrogen pickup is negligible (<30 wppm) in the cross sections 

containing the rupture opening. 

As discussed in Section 2.2, the 3-PBT with the load applied to the thick back of the rupture-opening 

cross section would bias the failure to occur in the cross section under the load that contained the rupture 

opening. The 4-PBT does not bias the failure location because the bending moment is uniform over the 

span Ls (see Fig. 7). As such, ANL has used 4-PBTs to determine post-LOCA sample failure location, 

maximum bending moment (measure of strength), failure energy (measure of toughness), and offset 

displacement (measure of plastic deformation). 

The bending moment M is determined from the applied lateral force as shown in Fig. 7. With analysis, 

the bending moment can also be related to an equivalent axial tensile load (see Section 3.4). The 

maximum energy Emax is determined by calculating the area under the load-displacement curve and 

dividing by 1000 mm/m to express Emax in joule energy units. For samples that fail during 4-PBTs, the 

failure energy is the maximum energy. The offset displacement is determined at the loading points (see 
33 

Figs. 7 and 8a). ANL 4-PBTs were conducted in the standard displacement-controlled mode at 2 mm/s 

(later reduced to 1 mm/s) and a maximum displacement at the loading points of 14 mm. In accordance 

with ASTM standards for fracture toughness and Charpy impact tests, the load was applied to cladding at 

180° relative to the rupture opening in order to subject the flawed rupture tips to maximum tensile 

stresses. Two tests were conducted with the rupture tips subjected to maximum compressive stresses to 

determine the effects of sample orientation relative to the bending moment. 

Table 4 summarizes the ANL test conditions for ballooning, rupture, oxidation, and quench of 

pressurized, as-fabricated 17×17 ZIRLO LOCA integral samples. Table 5 summarizes the 4-PBT results. 

With the exception of the OCZL#32 test sample, which was subjected to 4-PBT at RT, bending tests were 

conducted at 135°C. Table 5 contains three metrics for assessing cladding performance: maximum 

bending moment, failure or maximum energy, and offset displacement. Also included in Table 5 are the 

failure locations relative to the center of the rupture opening. LOCA integral samples with 

circumferential strains ≤32% all failed in a region where the cladding was fully brittle between the rupture 

tips and the hydrogen peaks. Six out of nine samples with ≥40% rupture strain failed in the rupture node 

location where some of the cladding had significant local ductility (e.g., thick back side of the balloon). 

Results for the three metrics are discussed in the following. 

Maximum Bending Moment 

Three ramp-to-rupture samples with rupture strains ranging from 21% to 69% were subjected to 4-PBTs 

to determine reference values for maximum bending moment and energy at ≈0% oxidation level (CPECR). After 14-mm displacement, the load-displacement curves were relatively flat, indicating that the 

end-of-test bending moments were close to their maximum values. 

The maximum bending moment values for ZIRLO samples are plotted in Fig. 20 where it is seen that 

bending moment is a strong function of cladding oxidation that occurs after rupture. 

Figure 20a shows results from 0% to 23% (maximum tested) CP-ECR. The data points for low rupturestrain samples that failed during 4-PBTs are within the scatter band of high rupture-strain samples. 

Figure 20b shows data for failed samples tested at 135°C to emphasize the data trend for ≥10% CP-ECR. 

For these samples, the best-fit linear correlation to the data is given by: 

Mmax = 13.92 – 1.073 (CP-ECR – 10%), N•m (5) 

Equation 5 represents a reasonable fit to the data for CP-ECR values ≥10%. The upper oxidation limit for 

Eq. 5 is 19.2% at which the failure bending moment (4.1 N•m) is equal to the failure bending moment at 

about 23% CP-ECR. Thus, the oxidation limits for Eq. 5 are 10% to 19% CP-ECR. At the current 

licensing limit of 17%, Eq. 5 gives 6.4 N•m as the failure bending moment. Measured values in the range 

of 16-18% CP-ECR and 135°C test temperature were 4.7 to 8.3 N•m. 

In order to achieve oxidation levels as low as about 11% CP-ECR (Tests OCZL#21 and #22) with a 

maximum oxidation temperature of 1200°C, it was necessary to increase the average pre-oxidation 

cladding wall thickness by lowering the rupture strain to about 20%. This change made the rupture 

opening cross section stronger than the higher hydrogen regions, in which severing then occurred. 

Similar results were obtained for all samples with ≤32% rupture strain. 
34 

Table 5. Summary of results for LOCA integral and post-LOCA bend tests with as-fabricated ZIRLO cladding. Reference LOCA test 

conditions were: 600- or 1200-psig fill pressure at 300°C, 5°C/s heating rate to 1200°C, 1200°C hold temperature, 3°C/s cooling rate to 800°C, and quench at 800°C. Reference conditions for 4-PBTs were 135°C test temperature and 2 mm/s displacement rate to 14-mm maximum displacement. The displacement rate was lowered to 1 mm/s after the OCZL#21 bend test. 

Test 

ID 

OCZL# 

Fill 

Pressure, 

psig 

Rupture 

Strain, %

(TR, °C) 

CP-ECR

% 

Quench 

at 

800°C 

Stress in 

Rupture 

Node 

Failure 

Location 

Maximum 

Bending 

Moment 

N•m 

Maximum 

Energy 

J 

Offset 

Displace. 

mm 

8 600 21 

(845±25) 

0 No Maximum 

tension 

No 

cracking 

20.9 >8.4 >7.7 

9 400 33 

(875±15) 

0 No Maximum 

tension 

No 

cracking 

20.6 >8.3 >7.7 

10 1600 69 

(715±10) 

0 No Maximum 

tension 

No 

cracking 

19.5 >7.7 >7.1 

12 1000 32 

(805±20) 

14.3 No Maximum 

compression 

-40 mm 

 +33 mm 

10.5 0.78 0 

13 1200 41 

(741±15) 

14.4 No Maximum 

tension 

Rupture 

opening 

8.8 0.58 0 

14 1200 47 

(735±6) 

18.2 Yes Maximum 

tension 

Rupture 

opening 

5.7 0.24 0 

15 1200 51 

(755±23) 

18.3 Yes Maximum 

compression 

Cracking; 

no failure 

8.9 >2.3 >13 

17 1200 49 

(750±17) 

13.1 Yes Maximum 

tension 

Rupture 

opening 

8.4 0.71 >0.5 

18 1200 43 

(748±4) 

12.3 Yes Maximum 

tension 

Rupture 

opening 

13.5 1.29 0 
35 

Table 5. Summary of results for LOCA integral and post-LOCA bend tests with as-fabricated ZIRLO cladding. Reference LOCA test 

conditions were: 600- or 1200-psig fill pressure at 300°C, 5°C/s heating rate to 1200°C, 1200°C hold temperature, 3°C/s cooling rate to 800°C, and quench at 800°C. Reference conditions for 4-PBTs were 135°C test temperature and 2 mm/s displacement rate to 14-mm maximum displacement. The displacement rate was lowered to 1 mm/s after the OCZL#21 bend test. (Cont’d) 

Test 

ID 

OCZL# 

Fill 

Pressure, 

psig 

Rupture 

Strain, % 

(TR, °C) 

CP-ECR

% 

Quench 

at 

800°C 

Stress in 

Rupture 

Node 

Failure 

Location 

Maximum 

Bending 

Moment 

N•m 

Maximum 

Energy 

J 

Offset 

Displace. 

mm 

19 600 24 

(840±12) 

17.2 Yes Maximum 

tension 

+23 mm 

-23 mm 

5.7 0.23 0 

21 600 27 

(850±10) 

10.5 Yes Maximum 

tension 

+33 mm 

-29 mm 

13.8 1.17 0 

22a 600 22 

(837±12) 

11.5 Yes Maximum 

tension 

+25 mm 

-27 mm 

11.1 0.83 0 

25a 1200 42 

(757±21) 

16.1 Yes Maximum 

tension 

-26 mm 

+26 mm 

8.3 0.50 0 

29a 1200 49 

(746±19) 

17.4 Yes Maximum 

tension 

Rupture 

opening 

4.7 0.40 >8.5 

32a,b 1200 49 

(748±8) 

17.2 Yes Maximum 

tension 

Rupture 

opening 

6.7 0.26 0 

37a 1200 46 

(755±30) 

23.1 Yes Maximum 

tension 

-40 mm 

+35 mm 

4.1 0.13 0 

43a 1200 50 

(738±16) 

17.6 Yes Maximum 

tension 

+35 mm 6.1 0.24 0 

aDisplacement rate lowered to 1 mm/s to better match maximum elastic strain rate for ring-compression tests. b4-PBT conducted at 30°C. 
 

36 

Based on one RT 4-PBT (OCZL#32 at 17% CP-ECR), the maximum bending moment appears to be 

relatively insensitive to test temperature. This result is quite different from RCT ductility results, which 

indicated that permanent and offset strains were strong functions of test temperature (135°C vs. RT). 

The 4-PBT strength results appear to be a meaningful measure of cladding performance in that the 

strength exhibits an expected decrease with increasing oxidation level. However, application of these 

results to LOCA acceptance criteria is not straightforward. Bending moments and axial bending stresses 

during quench are not anticipated to be significantly large, especially compared to potential axial stresses 

from partial contraction restraint. However, JAEA results for intact, fully restrained cladding gave an 

upper bound on the maximum axial force (1200 to 2400 N) that could be generated during quench 

following oxidation at 1200°C. It is also shown in Section 3.4 that a failure bending moment of ≈6 N•m 

corresponds to an axial failure load of ≈2500 N. Thus, it appears that the current 17% oxidation limit 

may be sufficient to ensure that the balloon region of low-burnup cladding would remain intact following 

quench. 

Failure Energy 

The failure or maximum (when no failure occurs) energy is plotted vs. oxidation level in Fig. 21 where it 

is seen that the failure energy is also a strong function of cladding oxidation that occurs after rupture. 

Load-displacement results from ramp-to-rupture tests (0% CP-ECR) were used to determine upper-bound 

energies of ≈8 J for unfailed samples through 14-mm displacement. Although this value does not 

represent the maximum energy that such samples could accumulate prior to failure, it does represent the 

maximum that can be accumulated through 14-mm displacement. As can be seen in Fig. 21a, there is a 

sharp decrease in failure energy with an increase in oxidation level from 0 to 10%. The decrease in 

failure energy is more gradual for increasing oxidation levels from 10 to 19%. It appears to level off to 

about 0.13 J in the oxidation range of 19 to 23% CP-ECR. This can be better seen in Fig. 21b for which 

the data are limited to samples that failed during 4-PBT loading at 135°C. 

Similar to the maximum bending moment results, there was no significant difference between failure 

energies for small (≤33% rupture strain) and large (≥40% rupture strain) balloons. At about 12% CPECR, the failure energies for 22% (OCZL#22) and 43% (OCZL#18) rupture-strain samples were 

1.17 J and 1.29 J, respectively. 

The best-fit linear correlation for failure energy vs. CP-ECR in the range of 10 to 19% oxidation level is 

given by: 

Emax = 1.225 – 0.1236 (CP-ECR – 10%), J (6) 

The failure energy represented by Eq. 6 gives 0.36 J and 0.11 J at 17% CP-ECR and 19% CP-ECR, 

respectively. The measured value at 23% CP-ECR was 0.13 J. More data would be needed in the range 

of 18 to 23% CP-ECR to confirm this leveling off of failure energy. However, the results clearly show 

that the oxidation level should be limited to maintain a high level of toughness. 
 

37 

0

2

4

6

8

10

12

14

16

18

20

22

0 5 10 15 20 25

CP-ECR (%)

Maximum Bending Moment (N•m) 

AF; ΔC/Ci ≥ 40%; 135°C

AF; ΔC/Ci ≤ 33%; 135°C

AF; ΔC/Ci ≥ 40%, 30°C

(a) Data for all samples subjected to 4-PBTs 

0

2

4

6

8

10

12

14

16

9 11 13 15 17 19 21 23 25

CP-ECR (%)

Maximum Bending Moment (N•m) 

(b) Data for failed samples subjected to 4-PBTs at 135°C 

Figure 20. Maximum bending moment as a function of maximum oxidation level (CP-ECR) for 

post-LOCA samples subjected to 4-PBTs with the rupture region in tension for all but 

one test. 4-PBTs were performed at 135°C and 2 or 1 mm/s to 14-mm maximum 

displacement. One bend test was conducted at 30°C. 
 

38 

0

1

2

3

4

5

6

7

8

9

0 5 10 15 20 25

CP-ECR (%)

Maximum Energy (J) 

AF; ΔC/Ci ≥ 40%; 135°C

AF; ΔC/Ci ≤ 33%; 135°C

AF; ΔC/Ci ≥ 40%, 30°C

(a) Data for all samples subjected to 4-PBTs 

0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4

1.6

9 11 13 15 17 19 21 23 25

CP-ECR (%)

Failure Energy (J) 

(b) Data for failed samples subjected to 4-PBTs at 135°C 

Figure 21. Maximum (for 0% CP-ECR) and failure (for ≥10% CP-ECR) energy as a function of 

oxidation level (CP-ECR) for post-LOCA samples subjected to 4-PBTs with the rupture 

region in tension for all tests but one. 4-PBTs were performed at 135°C and 2 or 1 mm/s 

to 14-mm maximum displacement. One bend test was conducted at 30°C. 
 

39 

Offset Displacement 

Offset displacement is a measure of plastic deformation. For three ramp-to-rupture samples (0% CPECR), there was a smooth transition between the constant load-displacement slope in the elastic bending 

regime and the decreasing slope after initiation of plastic flow in the axial direction (see Fig. 22). 

Measured offset displacements were in the range of 7 to 8 mm for these ductile samples. The loading 

stiffness values were reasonably close, but they appeared to increase somewhat with decreasing rupture 

strain: 120 N/mm for 69% rupture strain, 130 N/mm for 33% rupture strain, and 135 N/mm for 21% 

rupture strain. 

The load-displacement curves for oxidized samples were quite different from the ones for ramp-to-rupture 

samples. Most samples (9 out of 11 quenched samples) severed with an abrupt load drop during the 

linear portion of the loading ramp. These samples exhibited no offset displacement prior to crack 

initiation and propagation through the cross section. Six of the nine brittle samples failed outside the 

rupture node and the remaining three failed in the rupture node. 

The six quenched samples that severed outside the rupture node with zero offset strain all had rupture 

strains ≤32% and hydrogen contents and oxidation levels at the failure locations that would qualify them 

as brittle based on RCT results for non-deformed LOCA samples. Metallographic imaging and analysis 

were performed at one of two severed cross sections for the post-bend OCZL#19 sample, and the 

hydrogen content was determined from rings sectioned close to the severed cross section. Figure 23 

shows results of the post-bend characterization. The primary failure occurred at 24 mm below the rupture 

mid-span with 530-wppm hydrogen and 12% measured ECR. Based on the RCT results (see Fig. 4), 

cladding samples with this combination of hydrogen content and oxidation level are expected to fail in a 

brittle manner with no offset or plastic displacement. 

The three quenched samples that severed in the rupture node with zero offset strain had rupture strains 

≥40% and oxidation levels of 12, 17, and 18% CP-ECR. Figure 24 shows (a) the hydrogen distribution, 

(b) the severed sample, and (c) the cross section at the severed location for the OCZL#18 sample. The 

CP-ECR was only 12% and the hydrogen pickup at the failure location was negligible. For a ring with no 

rupture-opening flaw and uniform oxidation around the circumference, the sample would have exhibited 

high ductility (30 to 50%) at this oxidation level. However, the local oxidation for this sample varied 

from much greater than 35% at the rupture tips (see Fig. 25a) to 10% at the thick back region 180° from 

the center of the rupture opening (see Fig. 25b). Under tensile bending stress, the crack initiated at the 

brittle rupture tips and propagated rapidly through the cross section with no indication of offset 

displacement. Although it required a relatively high bending moment and energy to sever this sample, the 

“ductile” regions of the cladding did not have enough fracture toughness to blunt the growth of the crack. 

Figure 26 shows results of the post-bend characterization of a sample that was tested with the rupture 

opening in compression (OCZL#12). This sample had an average oxidation of 14% CP-ECR in the 

rupture node, and the primary failure occurred at 40 mm below the rupture mid-span with 1700-wppm 

hydrogen and 8% ECR. However, severing of the sample at 33 mm above the rupture mid-span appears 

to be an equally probable failure location based on hydrogen content and oxidation level. Figure 27 

shows the OCZL#15 sample, which was another sample subjected to reverse bending (rupture region 

under compression), after oxidation to 18% CP-ECR and cooling without quench. The offset strain for 

this intact sample was 13 mm. 
40 

0

900800700600500400300200100

1000

0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

Displacement (mm)

Load (N)

7.8 mm

K = 130 N/mm

Ramp to Rupture OCZL#9As-fabricated ZIRLO Sample135°C 4-Point-Bend TestMmax = 20.6 N•mEmax = 8.3 J

Offset Displacement = 7.8 mm

Figure 22. Four-point-bend test (4-PBT) load-displacement curve for ramp-to-rupture sample OCZL#9. The as-fabricated ZIRLO sample 

ruptured at 875±15°C with a mid-wall rupture strain of 33%. The 4-PBT was conducted at 135°C and 2 mm/s displacement rate 

to 14-mm maximum displacement. 
 

41 

(a) Hydrogen-content profile 

b) Measured values at failure locations 

(c) Low-magnification image of severed cross section at -24 mm 

Figure 23. Post-bend characterization of OCZL#19 sample that was subjected to bending at 135°C 

with the rupture region under maximum tensile stress: (a) hydrogen-content profile; (b) 

measured values at failure locations; and (c) low-magnification image of severed cross 

section at 24 mm below rupture mid-span. 
 

42 

(a) Hydrogen-content profile 

(b) Failure location 

(c) Low-magnification image of severed cross section 

Figure 24. Post-bend characterization of the OCZL#18 sample oxidized to 12% CP-ECR, quenched, 

and subjected to bending at 135°C with the rupture region under maximum tensile 

stress: (a) hydrogen-content profile; (b) failure location; and (c) low-magnification 

image of the severed cross section. 
 

43 

(a) Rupture tip 

(b) Back side of rupture node cross section 

Figure 25. Metallographic images for the OCZL#18 severed cross section following oxidation to 

12% CP-ECR, quench, and bending at 135°C with the rupture region under maximum 

tension: (a) rupture tip with 0.14-mm average metal wall and (b) back side of cross 

section with 0.44-mm metal wall thickness.
 

44 

(a) Hydrogen-content profile 

(b) Measured values at failure locations 

(c) Low-magnification image of severed cross section at -40 mm 

Figure 26. Post-bend characterization of OCZL#12 sample that was subjected to bending at 135°C 

with the rupture region under maximum compressive stress: (a) hydrogen-content 

profile; (b) measured values at failure locations; and (c) low-magnification image of 

severed cross section at 40 mm below rupture mid-span. 
 

45 

(a) 1st loading 

(b) 2nd loading 

(c) Sample appearance after 2nd loading 

Figure 27. Post-bend results for OCZL#15 sample (18% CP-ECR and quench) with rupture node in 

compression (reverse bending), 135°C test temperature and 2 mm/s displacement rate: 

(a) load-displacement curve for 1st loading; (b) load-displacement curve for 2nd loading; 

and (c) sample appearance after 2nd loading. 

Figure 28 shows the OCZL#17 sample, which was subjected to standard bending (rupture region under 

tension). The back region of the cladding remained intact after 10-mm displacement with an offset 

displacement of 6.3 mm. In Fig. 28a, the load-displacement curve is truncated at 4-mm displacement. 
 

46 

From 4- to 10-mm displacement, the load remained quite low (3 to 13 N) as the ductile ligament at the 

back of the cladding deformed plastically. This ductile deformation region contributed about 10% to the 

maximum energy. The OCZL#29 sample was also subjected to the standard ANL bend test. This test 

was terminated after ≈3-mm displacement (Fig. 29) and reloaded (Fig. 30). Cracking began at a relatively 

low load (188 N) and bending moment (4.7 N•m). The load dropped rather abruptly from 188 N to 107 N 

and again from 124 N to 85 N. At the end of these two load drops, the crack had propagated across more 

than half the cross section. After reloading, the crack continued to propagate in a ductile manner leaving 

only the back region intact. 

In summary, the post-LOCA 4-PBT load-displacement curves give three types of results: severing with 

zero offset displacement in brittle regions outside the rupture node; severing with zero offset displacement 

in brittle-to-ductile regions of the rupture cross section; and partial-wall, brittle cracking followed by 

ductile crack growth during which offset strain is observed. It is clear from these tests at ≤18% CP-ECR 

that a significant fraction of the cladding cross section in the rupture node has plasticity; yet most of these 

tests show no offset displacement. Offset displacement appears to be affected by balloon size, loading 

history, circumferential temperature gradient, and other factors not yet quantified. Further, well behaved 

trends are not observed as they are for bending moment and failure energy. Although offset displacement 

is related to ductility, it is not a useful metric for behavior of the balloon region. 

3.4 Comparison of ANL 4-PBT and JAEA Axial-restraint Tests 

ANL 4-PBT results can be compared to JAEA axial-restraint results as both tests induce axial stresses. 

The stress distributions are different in that the stress acting on a cross section is tensile everywhere in the 

axial tensile test, while the stress in the axial bend tests varies from tensile to compressive across the cross 

section. However, failure of a brittle material is governed by the maximum tensile stress. Given the 

failure bending moment for a 4-PBT, the equivalent axial force that gives the maximum tensile stress can 

be determined. This determination can be made more easily for 4-PBT samples that fail outside the 

rupture region. Figure 23c shows the severed cross section for the 17% CP-ECR sample from Test 

OCZL#19. The cross section is nearly circular. Table 6 gives the dimensions determined from 

profilometry, low-magnification metallography, and high-magnification metallography. 

For the 4-PBT sample, the maximum tensile stress (σmax) in the metal is related to the maximum bending 

moment (Mmax) according to: 

σmax = Mmax Rmo/Ieq (7) 

where Rmo is the outer radius (Rmo = Dmo/2) of the metal wall (5.285 mm in this case). 

The maximum tensile stress in the metal is σmax = 5.7 N•m (1000-mm/m) (5.285 mm)/248 mm4

 = 121.5 

MPa. 
 

47 

(a) Load-displacement curve for 1st 4-mm displacement 

(b) Post-test image of side view 

(c) Post-test image of rupture view 

Figure 28. 4-PBT results for Test Sample OCZL#17 oxidized to 13% CP-ECR, quenched, and tested 

at 135°C and 2 mm/s displacement rate: (a) load-displacement curve for 1st 4-mm 

displacement (out of 10-mm total displacement); (b) post-test image of side view; and 

(c) post-test image of rupture view. 
 

48 

0

20

40

60

80

100

120

140

160

180

200

0 0.5 1 1.5 2 2.5 3 3.5

Displacement (mm)

Load (N)

K = 118 N/mm

(a) 1st loading-unloading sequence 

(b) Side view after 1st loading-unloading 

Figure 29. Test Sample OCZL#29 oxidized to 17% CP-ECR, quenched, and subjected to 4-PBTs at 

135°C and 1 mm/s displacement rate: (a) load-displacement curve for 1st loadingunloading sequence and (b) sample appearance after 1st loading-unloading. 
 

49 

0

20

40

60

80

100

120

140

160

180

200

0 1 2 3 4 5 6 7 8 9 10 11

Displacement (mm)

Load (N)

(a) Second loading-unloading 

(b) Rupture view after 2nd loading-unloading 

Figure 30. Test Sample OCZL#29 oxidized to 17% CP-ECR, quenched, and subjected to 4-PBTs at 

135°C and 1 mm/s displacement rate: (a) load-displacement curve for 2nd loadingunloading sequence and (b) sample appearance after 2nd loading-unloading. 
 

50 

Table 6. Dimensions, properties, and failure bending moment at severed location for test 

sample OCZL#19 with 17% maximum CP-ECR in the rupture node and 12% CP-ECR at 

the severed location. 

Parameter Value Comment 

Outer Diameter (OD, Do), mm 10.67 Measured 

OD Oxide Thickness (δox)o, μm 50±6 Measured 

Metal OD (Dmo), mm 10.57 Dmo = Do – 2 (δox)o

Metal Wall Thickness (hm), mm 0.46±0.04 Measured 

Metal ID (Dmi), mm 9.65 Dmi = Dmo – 2 hm

ID Oxide Thickness (δox)i, μm 35±6 Measured 

Inner Diameter (ID, Di), mm 9.58 Di = Dmi – 2 (δox)i

Cross-sectional Area (A), mm2

OD Oxide (Aox)o

Metal (Am) 

ID Oxide (Aox)i

1.67 

14.61 

1.06 

(Aox)o = (π/4) [(Do)

2

 - (Dmo)

2

] 

Am = (π/4) [(Dmo)

2

 - (Dmi)

2

] 

(Aox)i = (π/4) [(Dmi)

2

 - (Di)

2

] 

Area Moment of Inertia (I), mm4

OD Oxide (Iox)o

Metal (Im) 

ID Oxide (Iox)i

23.5 

187.1 

12.2 

(Iox)o = (π/64) [(Do)

4

 - (Dmo)

4

] 

Im = (π/64) [(Dmo)

4

 - (Dmi)

4

] 

(Iox)i = (π/64) [(Dmi)

4

 - (Di)

4

] 

Young’s Modulus (E) at 135°C, MPa 

Oxide (Eox) 

Metal (Em) 

14.8×104

8.65×104

MATPRO [3] 

MATPRO [3] 

Equivalent A (Aeq) Relative to Metal, mm2 19.3 Aeq = Am + (Eox/Em) Aox

Equivalent I (Ieq) Relative to Metal, mm4 248 Ieq = Im + (Eox/Em) Iox

Maximum Bending Moment (Mmax), N•m 5.7 Measured 

For the axial tensile test, the maximum load (Pmax) is related to the maximum metal stress (σmax) according 

to: 

Pmax = σmax Aeq (8) 

Setting σmax = 121.5 MPa and Aeq = 19.3 mm2

 gives Pmax = 2345 N. If the correlation value (see Eq. 5) of 

6.4 N•m at 17% CP-ECR had been used, then Pmax would be equal to 2632 N. 

Based on the results for the JAEA fully restrained samples [13,14], the maximum load measured for 

samples that survived quench from 700°C to <100°C was in the range of 1200 N to 2400 N. For samples 

oxidized at 1200°C to >12% BJ-ECR, the maximum was 1200 to 2000 N. These results are for cladding 

with pre-hydride levels ≥ 100 wppm. Thus, even under these conditions, it appears that the ANL sample 

from Test OCZL#19 would have survived JAEA fully restrained quench without failure. As the ANL 

bend tests were performed at 135°C, it would be more proper to say that the 17% CP-ECR OCZL#19 

sample would have survived the JAEA full-constraint test with quench from 700°C to 135°C. 

A similar analysis was performed for the OCZL#12 sample which had 14% maximum CP-ECR in the 

rupture node and 8% measured ECR at the severed location 40 mm below the rupture mid-span. The 

failure bending moment was 10.5 N•m and the equivalent cross-sectional area and area moment of inertia 
 

51 

were 17.9 mm2

 and 249 mm4, respectively. Using these values, along with a metal outer radius of 5.467 

mm, gives an equivalent axial failure load of 4127 N, which is much higher than any load measured in the 

JAEA fully constrained tests. 

Equations 7 and 8 can be combined to give the JAEA tensile failure load (Pmax) as a function of the ANL 

failure bending moment (Mmax): 

Pmax = Mmax Rmo (Aeq/Ieq) (9) 

In principle, Eq. 9 can be used for samples that severed in the rupture node (e.g., OCZL#29 at 17% CPECR). However, the severed cross section for the OCZL#29 sample (49% rupture strain) would be 

similar to the one shown in Fig. 24c for the OCZL#18 sample (43% rupture strain). The equivalent area 

and area moment of inertia would have to be calculated numerically for this case. Although the 

maximum bending moment for the OCZL#29 sample (4.7 N•m) was less than for the OCZL#19 sample 

(5.7 N•m), the sample did exhibit plastic flow following the first significant load drop and did not sever 

into two pieces. This adds additional complexity to the determination of the axial load needed to sever 

the sample at this location. 

 

53 

4 CONCLUSIONS 

The LOCA acceptance criteria that limit peak oxidation temperature and maximum oxidation level are 

based on retention of ductility. For non-deformed (i.e., non-ballooned) regions of a fuel rod during a 

LOCA, the ductility concept is unambiguous and can be determined with straightforward testing. Among 

the testing methods evaluated, 3-PTB tests and RCTs are quite good for LOCA evaluation because they 

can measure loads and displacements readily – and they give similar results with regard to the ductile-tobrittle transition oxidation level as a function of pre-oxidation hydrogen content. The 3-PBT has 

advantages for as-fabricated and pre-hydrided material when long specimens are available. However, 

RCTs use much shorter specimens, and this becomes advantageous when testing irradiated material. 

Therefore, for comprehensive studies of as-fabricated, pre-hydrided, and irradiated cladding, the RCT is 

preferred. 

Although the LOCA temperature and oxidation criteria may also protect ballooned and ruptured regions 

from severing and fragmenting during and following quench, ductility will not be retained everywhere in 

this region and interpretation of test results is not straightforward. Higher hydrogen-content regions from 

the rupture edge to the hydrogen peak will contain brittle material. Also, within the rupture region cross 

sections, the cladding transitions from brittle (at the thin, heavily oxidized rupture tips) to ductile (at the 

thicker back region which is at lower oxidation level) at the averaged 17% oxidation level. 

Of the two methods being used for testing ballooned sections, the ANL 4-PBT is preferred over the JAEA 

full- or partial-restraint quench test because load-versus-displacement curves can be measured accurately 

in the 4-PBT. This permits the determination of offset displacement (measure of plastic deformation), 

maximum bending moment (measure of strength), and failure energy (measure of toughness). Offset 

displacement of a ballooned segment is not well behaved, however, and it is therefore not a useful metric 

for structural behavior. No offset displacement was observed for bending failures in the region bounded 

from the rupture tip to the hydrogen peak near the neck of the balloon. The offset displacement was also 

zero prior to >50% severing of the cross section within the rupture region. 

On the other hand, the 4-PBT maximum bending moments and failure energies exhibited smooth trends 

toward very low values with increasing oxidation level up to 19% CP-ECR, but remained significantly 

above zero for oxidation levels ≤17%. Our calculation based on bending moment showed that a sample 

oxidized to 17% CP-ECR would have sufficient strength to survive the fully restrained tests performed by 

JAEA. Further, none of our unrestrained samples failed upon quench, and none of the samples tested to 

failure in the ANL program fragmented or failed in a “low-toughness” mode. By contrast, glass and 

ceramic rods with much lower toughness than the post-LOCA samples sever into as many as 5 to 10 

pieces when subjected to 4-PBT loading. Therefore, the 4-PBT can discriminate between good and poor 

structural performance of ballooned segments. The ANL data used for this evaluation show that the 

current limits of 17% CP-ECR and 1204°C protect fresh cladding during quench not only from 

fragmentation but also from severing into two pieces under a wide range of loading conditions. 

 

55 

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NUREG/CR-7139 Assessment of Current Test Methods for Post-LOCA Cladding Behavior