Cladding Behavior during 

Postulated Loss-ofCoolant Accidents

Office of Nuclear Regulatory Research

NUREG/CR-7219

ANL-16/09

Cladding Behavior during 

Postulated Loss-ofCoolant Accidents

Manuscript Completed: March 2016

Date Published: July 2016

Prepared by: M.C. Billone, Y. Yan,* T.A. Burtseva, and R.O. Meyer

Nuclear Engineering Division

Argonne National Laboratory

9700 South Cass Avenue

Argonne, IL 60439

*Currently with Oak Ridge National Laboratory

H.H. Scott, NRC Project Manager

NRC Job Code Number V6199

Office of Nuclear Regulatory Research

NUREG/CR-7219

ANL-16/09

 

 

iii

ABSTRACT

A previous report by this laboratory provided results on cladding embrittlement, breakaway oxidation, 

and ballooning with rupture under conditions of loss-of-coolant accidents (LOCAs). This report updates 

those results, provides additional results in those areas, and adds results of mechanical testing of cladding 

after ballooning, rupture, oxidation at elevated temperature, and quench. Significant conclusions are as 

follows. Embrittlement of high-exposure cladding is accelerated by hydrogen that is absorbed during 

normal reactor operation (i.e., pre-transient) and the acceleration has been quantified as an oxidation limit 

vs. pre-transient hydrogen content in the cladding metal. Breakaway oxidation, which leads to excessive 

hydrogen pickup and embrittlement during a LOCA transient, can occur during times that are relevant for 

these accidents. Conditions that lead to breakaway are only partially understood. Factors that contribute to 

increased susceptibility to breakaway oxidation – especially surface finish and scratches – have been 

investigated. Ballooning strains increase as rupture temperature decreases within the temperature range 

explored in this work. Rupture temperature depends on alloy composition, pre-transient hydrogen level, 

internal gas pressure, and geometrical factors used to convert gas pressure to hoop stress. Four-point bend 

tests are effective for measuring post-quench failure limits for ballooned, ruptured, and oxidized cladding. 

The results indicate that strength and failure energy degrade significantly with increasing pre-transient 

hydrogen and transient oxygen levels, based on cross-section-average values for these parameters. 

Further, using the oxidation limit vs. hydrogen content derived for non-ballooned regions would ensure 

the preservation of strength for ballooned regions. 

 

v

FOREWORD 

Fuel rod cladding is the first barrier for retention of fission products, and the structural integrity of the 

cladding ensures coolable core geometry. In the early 1990s, new data from foreign research programs 

showed degraded cladding behavior for high-burnup fuel compared with low-burnup fuel in tests 

designed to simulate postulated accidents. Interim actions were taken, but it became clear that 

extrapolation from a low-burnup database needed to be reassessed more carefully for regulatory purposes.

One of NRC’s central regulations used in plant licensing deals with postulated loss-of-coolant accidents 

(LOCAs). A portion of that regulation in 10 CFR 50.46(b) specified criteria that were derived from tests 

with non-irradiated Zircaloy cladding, and these criteria limit the peak cladding temperature and the 

maximum cladding oxidation during the accident. These two limits are known as embrittlement criteria. 

Their purpose is to prevent cladding embrittlement during a LOCA, thus ensuring that the general core 

geometry will be maintained and be coolable.

Since 1997, NRC has sponsored a research program at Argonne National Laboratory to reassess these 

limits for the possible effects of fuel burnup. Two formal NRC reports have been published, NUREG/CR6967 and NUREG/CR-7139. The latter report discussed test methods to assess fuel-rod cladding behavior 

following a 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 purpose of this document is to (1) provide final results of tests performed and conclusions on 

hydrogen embrittlement of zirconium-based cladding alloys, and (2) support the technical bases for a 

revised regulation – now 10 CFR 50.46(c) – and Regulatory Guides 1.222, 1.223 and 1.224. It is 

considered appropriate to issue this document as a new report because it contains new data, as well as 

revisions to previously published results. 

 

vii

CONTENTS 

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

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

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

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

TABLES ................................................................................................................................... xxvii

EXECUTIVE SUMMARY ..................................................................................................... xxxi

ACKNOWLEDGMENTS .......................................................................................................xxxv

ACRONYMS AND ABBREVIATIONS ............................................................................. xxxvii

SYMBOLS .............................................................................................................................. xxxix

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

1.1 Cladding Materials .................................................................................................................... 1

1.2 Transient Cladding Temperatures ............................................................................................. 2

1.3 Transient Phase Transformations .............................................................................................. 3

1.4 Oxidation Equations .................................................................................................................. 4

1.5 Embrittlement Mechanisms ...................................................................................................... 5

1.5.1 Beta-layer embrittlement by oxygen ........................................................................... 5

1.5.2 Beta-layer thinning ..................................................................................................... 7

1.5.3 Localized hydrogen-induced embrittlement in the balloon region ............................. 7

1.5.4 Hydrogen-enhanced beta-layer embrittlement ............................................................ 7

1.5.5 General hydrogen-induced embrittlement from breakaway oxidation ....................... 8

1.5.6 Oxygen pickup from the cladding inner surface ......................................................... 9

1.6 Conventional Use of Equivalent Cladding Reacted (ECR) .................................................... 10

1.7 Pre-transient Hydrogen Content in the Cladding Metal (CHM) ............................................... 10

2 MATERIALS, TEST METHODS, AND PROCEDURES .............................................11

2.1 Description of Cladding Alloys and High-Burnup Fuel Segments ......................................... 11

2.2 General Description of Apparatus and Test Samples .............................................................. 16

2.3 Oxidation and Quench Tests with Short Cladding Samples ................................................... 16

2.4 Ductility Determination Using Ring-Compression Tests ....................................................... 22

2.4.1 Background information and general description ..................................................... 22

2.4.2 Ductility and ductility criteria ................................................................................... 24

2.4.3 Measured vs. calculated loading slopes .................................................................... 26

2.5 LOCA Integral Tests with Pressurized Cladding Samples ..................................................... 27

2.6 Four-point-bend Tests (4-PBTs) ............................................................................................. 28

3 EMBRITTLEMENT OF NON-BALLOONED CLADDING........................................31

3.1 Embrittlement of As-fabricated Cladding ............................................................................... 31

3.2 Embrittlement of Pre-hydrided Cladding ................................................................................ 33

3.2.1 Review of previous results ........................................................................................ 33

3.2.2 Current results ........................................................................................................... 34

3.3 Embrittlement of Irradiated Cladding ..................................................................................... 43

3.3.1 Review of previous results ........................................................................................ 43
 

viii

3.3.2 Current PQD results for high-burnup ZIRLO ........................................................... 49

3.4 Discussion of PQD Embrittlement Results ............................................................................. 57

4 BREAKAWAY OXIDATION ..........................................................................................71

4.1 Breakaway Oxidation for ZIRLO at Constant Oxidation Temperature .................................. 72

4.1.1. ZIRLO materials for breakaway oxidation ............................................................... 73

4.1.2. Breakaway oxidation test parameters and results ..................................................... 73

4.2 Breakaway Oxidation for M5 Cladding .................................................................................. 85

4.2.1 M5 materials for breakaway oxidation ..................................................................... 85

4.2.2 Test parameters and results for M5 breakaway oxidation ........................................ 86

4.3 Discussion on Breakaway Oxidation under Isothermal Conditions ........................................ 88

4.4 Breakaway Oxidation under Transient Temperature Conditions ............................................ 95

4.4.1 Transient temperature histories ................................................................................. 95

4.4.2 Outer-surface appearance of transient oxidation samples ........................................ 98

4.4.3 Weight gain and hydrogen pickup .......................................................................... 100

4.4.4 Discussion of transient oxidation results ................................................................ 100

5 BEHAVIOR OF BALLOONED AND RUPTURED CLADDING .............................107

5.1 Thermal and Metallurgical Benchmark Test Results ............................................................ 108

5.2 Scoping Test Results ............................................................................................................. 110

5.3 Determination of Rupture Strains and Post-Rupture Cladding Wall Thickness ................... 115

5.4 Bend-Test Results for AF ZIRLO ......................................................................................... 124

5.5 Bend-Test Results for PH ZIRLO ......................................................................................... 135

5.6 Discussion of Strength of Ballooned and Ruptured Cladding .............................................. 141

6 REFERENCES .................................................................................................................147

APPENDIX A: TEMPERATURE AND PRESSURE HISTORIES ............................... A-1

APPENDIX B: AXIAL PROFILES OF POST-RUPTURE CLADDING STRAINS ...B-1

APPENDIX C: PHOTOGRAPHS OF POST-OXIDATION LOCA INTEGRAL

 SAMPLES .................................................................................................. C-1

APPENDIX D: 4-PBT LOAD-DISPLACEMENT CURVES .......................................... D-1

APPENDIX E: PHOTOGRAPHS OF POST-BEND LOCA INTEGRAL 

SAMPLES ...................................................................................................E-1
 

ix

FIGURES

1. Schematic of cladding temperature history during a LOCA transient. .................................. 2

2. Qualitative diagram of oxygen concentration in Zircaloy cladding exposed at high 

temperature to steam on the outside surface and cooled to room temperature. OD and 

ID are outer and inner diameters, respectively. ...................................................................... 4

3. Average oxygen content (co) in the beta layer as a function of CP-ECR for isothermal 

oxidation at 1200oC. The calculation was performed for 17×17 Zry-4 with 0.12 wt. 

% initial oxygen content (coi) and 0.57 wt. % solubility limit (So). ....................................... 6

4. Three-layer diffusion “couple” character of oxygen sources and cladding metal. ................. 9

5. Well-developed fuel-cladding bond layer (gray) at the fuel mid-plane of a Zry-4-clad 

fuel rod irradiated to 67 GWd/t in the H.B. Robinson reactor. .............................................. 9

6. Test train for two-sided oxidation tests with holes drilled into the Inconel holder 

below (left) the sample for steam ingress and holes drilled into the holder above 

(right) the sample for steam egress. Also shown are the three thermocouples 

permanently welded to the Inconel holder just above the sample (silver in 

appearance). ......................................................................................................................... 17

7. Thermal benchmark results for bare ZIRLO cladding. At 100-s into the heating 

phase, the average sample temperature was 1200°C. After 150 s, the cladding 

temperature was 1202±7°C. Quench at 800°C is not shown in this figure. ......................... 19

8. Thermal histories for bare and pre-oxidized (40-μm) ZIRLO cladding tested in the 

same test train with the same control parameters. At 150-s into the heating phase, the 

average temperature for the pre-oxidized sample was 1200±3°C. ....................................... 21

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

controlled and the response force (P) is measured. .............................................................. 23

10. Load-displacement curve for as-fabricated (AF) M5 cladding compressed at RT to 

1.7-mm total displacement. .................................................................................................. 23

11. Load-displacement curve for as-fabricated ZIRLO oxidized (one-sided) at 1200°C to 

17% CP-ECR. The sample had one tight through-wall crack at the 6 o’clock position. ..... 24

12. Permanent strain vs. CP-ECR for 17×17 Zry-4 oxidized at 1200oC, cooled at ≈13oC/s 

to 800oC, quenched, and ring-compressed at RT and 135oC................................................ 25

13. Difference between offset and permanent strains as a function of oxidation level for 

as-fabricated, pre-hydrided, and high-exposure cladding alloys oxidized at 1200°C 

and compressed 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%. ......................................................................................................................... 26
 

x

14. Schematic of four-point-bend test (4-PBT) loading. The bending moment is constant 

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

measured. ............................................................................................................................. 28

15. PQD results for 15×15 Zry-4 (HBR-type) following oxidation at 1200°C and quench 

at 800°C: (a) permanent strain and (b) offset strain. ............................................................ 32

16. Sectioning diagrams for pre-hydrided ZIRLO segments: (a) Segment A13 and (b) 

Segment A15. Nominal lengths are shown for the oxidation test samples. ......................... 35

17. Thermal history determined from benchmark test for oxidation tests to be conducted 

with pre-hydrided ZIRLO. Quench at 800°C, which was used for PQD tests, is not 

shown in this figure because benchmark samples are cooled without quench..................... 37

18. RCT ductility (i.e., offset strain) results for pre-hydrided ZIRLO as a function of CPECR oxidation level. ............................................................................................................ 41

19. Micrographs taken from a sample sectioned near the middle of ZIRLO segment 

648A: (a) cladding cross section; (b) corrosion layer; and (c) fuel-cladding bond 

layer. ..................................................................................................................................... 50

20. Hydride morphology for high-burnup ZIRLO sample 648A5: (a) circumferential 

hydrides (typical); and (b) circumferential and radial hydrides. Based on measured 

values for an adjacent ring (648A4), the CH was 318±30 wppm. ........................................ 51

21. Thermal history for bare ZIRLO cladding determined from test ZLU#145. At 150 s 

into the heating phase, the average sample temperature was 1202±7oC. ............................. 53

22. Thermal history for bare ZIRLO (ZLU#145) and pre-oxidized (ZLU#145A) 

cladding. At 150 s into the heating phase, the average temperature for the preoxidized sample was 1200±3oC. .......................................................................................... 54

23. Thermal history and test times for high-burnup ZIRLO tests conducted to 8% (ZLI#9

& 11) and 10% (ZLI#7 & 8) CP-ECR. ................................................................................ 56

24. Embrittlement oxidation level (CP-ECR) as a function of pre-transient hydrogen 

content in cladding metal for peak oxidation temperatures of 1190–1200°C for <400 

wppm hydrogen content and 1130–1190°C for >400 wppm hydrogen content. ................. 58

25. Exponential fit to data for embrittlement CP-ECR as a function of pre-transient 

hydrogen content in cladding metal. .................................................................................... 59

26. Load-displacement curve for HBU ZIRLO cladding ring 648B2B oxidized to 10% 

CP-ECR and cooled without quench. ................................................................................... 61

27. Load-displacement curve for high-burnup ZIRLO cladding ring 648B2D oxidized to 

10% CP-ECR and cooled without quench. .......................................................................... 62
 

xi

28. Load-displacement curve for high-burnup ZIRLO cladding ring 648A3B oxidized to 

10% CP-ECR and quenched at 800oC. ................................................................................. 63

29. Load-displacement curve for high-burnup ZIRLO cladding ring 648A3C oxidized to 

10% CP-ECR and quenched at 800oC. ................................................................................. 64

30. Load-displacement curve for high-burnup ZIRLO cladding ring 648A6B oxidized to 

8% CP-ECR and quenched at 800oC. ................................................................................... 65

31. Load-displacement curve for high-burnup ZIRLO cladding ring 648A6C oxidized to 

8% CP-ECR and quenched at 800oC. ................................................................................... 66

32. Load-displacement curve for high-burnup ZIRLO cladding ring 648A6D oxidized to 

8% CP-ECR and quenched at 800oC. ................................................................................... 67

33. Load-displacement curve for high-burnup ZIRLO cladding ring 648B6B oxidized to 

8% CP-ECR and quenched at 800oC. ................................................................................... 68

34. Load-displacement curve for HBU ZIRLO cladding ring 648B6C oxidized to 8% 

CP-ECR and quenched at 800oC. ......................................................................................... 69

35. Thermal benchmark results with two TCs welded onto as-fabricated, 17×17 ZIRLO2008 cladding. The long-time sample hold temperature was 975±5oC. For datagenerating tests, the furnace power was increased to give 1000oC oxidation 

temperature. .......................................................................................................................... 74

36. Surface appearance of W- and ANL-cleaned ZIRLO samples after 4000 s at 1000oC. ....... 75

37. Metallographic images (at the same cross section) for W-cleaned sample oxidized for 

4000 s at 1000oC: (a) Area 8, which has outer-surface monoclinc oxide (≈0.4 mm 

long); (b) Area 7 with tetragonal oxide; and (c) Area 1 with tetragonal oxide. ................... 76

38. Hydrogen pickup for ANL-cleaned and W-cleaned ZIRLO samples oxidized at 

1000oC. ................................................................................................................................. 78

39. RCT offset strains at 135oC for Zry-4 and ZIRLO rings following oxidation at 970-

1000oC, local-to-global breakaway oxidation, and cooling without quench. ....................... 78

40. Circumferential variation in hydrogen content measured after ring-compression 

testing for an 8-mm-long ring sectioned from the ANL-cleaned ZIRLO sample 

oxidized at 1000oC for 4000 s. ............................................................................................. 79

41. Thermal benchmark results for test ZLU#129. Peak temperature was ≈1065oC and 

long-time hold temperature was ≈980±5oC. ......................................................................... 80

42. Surface appearance of ANL-cleaned and W-cleaned samples after oxidation at 

980oC: (a) ANL-cleaned sample after 3400 s; (b) W-cleaned sample #1 after 3200 s; 

and (c) W-cleaned sample #2 after 3200 s. .......................................................................... 81
 

xii

43. Thermal benchmark results for test ZLU#132. Long-time hold temperature is 

975±3oC. The holder control temperature was increased by 10oC for data generating 

tests, which increased the long-time sample temperature to 985±5oC. ................................ 83

44. Hydrogen pickup for ANL- and W-cleaned samples oxidized at 970−985oC. ANLcleaned samples included as-fabricated, MS (20-μm depth), and PF (<1-μm oxide) 

ZIRLO. ................................................................................................................................. 85

45. Thermal benchmark results for 17×17 M5 (Validation Lot) oxidation tests at 1000oC. ...... 87

46. Metallographic images around the cross section of the MU#80 test sample, which 

show wavy interfaces between oxide and metal layers. However, no local breakaway 

oxidation was observed after 5000 s at 985oC. .................................................................... 89

47. Evolution of outer-surface oxide-layer/metal interface for ZIRLO oxidized at 1000°C 

for (a) 1500 s; (b) 2440 s; (c) 2480 s; and (d) 4000 s. .......................................................... 90

48. Evolution of outer-surface oxide-layer/metal interface for M5 cladding oxidized for: 

(a) 3440 s at 1000°C; (b) 5000 s at 985°C (low-magnification); and (c) 5000 s at 

985°C. .................................................................................................................................. 92

49. Evolution of oxide-layer/metal interface for polished Zry-4 oxidized for: (a) 3440 s 

at 1000°C and (b) 5000 s at 985°C....................................................................................... 93

50. Metallographic images at eight locations around the cross section of polished Zry-4 

sample oxidized at 1000°C for 5000 s, which show scalloped (wavy) oxidelayer/metal interfaces. Local breakaway oxidation was observed in Area BK. ................... 94

51. Schematics of transient temperature histories for breakaway oxidation tests with 

ZIRLO cladding samples. .................................................................................................... 96

52. Thermal benchmark results for Test ZLU#138 with as-fabricated ZIRLO-2008. ............... 97

53. Temperature history of the control TC for Test ZLU#139. ................................................. 97

54. Temperature history of the control TC for Test ZLU#140. ................................................. 98

55. Temperature history of control TC for Test ZLU#142. 1400-s hold time following 

cycling. ................................................................................................................................. 99

56. Temperature history of control TC for Test ZLU#143. 400-s hold time following 

cycling. ................................................................................................................................. 99

57. Appearance of as-received ZIRLO cladding outer-surface after steam oxidation tests 

(a) ZLU#139 for Transient #1, (b) ZLU#140 for Transient #2, (c) ZLU#142 for 

Transient #3, and (d) ZLU#143 for a modified Transient #3. Photos are in gray scale. .... 100
 

xiii

58. (a) ZIRLO sample after steam-oxidation at Transient #3 conditions for 2800 s total 

test time with cycling from 930°C to 1030°C (Test ZLU#143), and (b) highmagnification. .................................................................................................................... 103

59. High-magnification images of Area T in Fig. 58a showing different outer-surface 

areas of the ZLU#143 sample: (a) lustrous black area, (b) transition area from blackto-yellow, and (c) yellow area. ........................................................................................... 104

60. Axial distribution of hydrogen pickup for the ZLU#142 test sample. ............................... 104

61. Hydrogen-pickup vs. time data for ZIRLO oxidized in steam at 970−985°C: smooth 

samples (isothermal), pre-scratched (≈20-μm depth) samples (isothermal), preoxidized samples (isothermal) and smooth samples (transient) oxidized with five 

cycles from 930 to 1030°C for 400 s and 980°C isothermal temperature. ......................... 105

62. Schematic of ballooned and ruptured cladding, which includes the parameters 

routinely measured and/or calculated following a LOCA integral test. The sketch is 

not to scale. ........................................................................................................................ 108

63. Image of ramp-to-rupture test OCZL#11 sample cross section through rupture-node 

mid-span. Rupture width (δR) was 6.2-mm and ΔCmR/Cmi was 40%. ................................ 112

64. Low magnification image of the cross section through rupture-opening mid-span for 

the OCZL#7 sample with 22% rupture strain prior to 16% CP-ECR oxidation. ............... 113

65. Hydrogen content, ballooning strain, and oxidation level at rupture mid-span (CPECR) and failure locations (measured ECR) for the OCZL#12 sample after the 

reverse 4-PBT. ................................................................................................................... 114

66. Axial profile of hydrogen content for the OCZL#12 sample following oxidation to 

14.5% CP-ECR. ................................................................................................................. 115

67. Low magnification image of cross section through rupture mid-span of the OCZL#18 

test sample with 43% mid-wall circumferential strain oxidized to 12% CP-ECR. ........... 117

68. Rupture cross section for Studsvik Test 185 with AF ZIRLO sample pressurized to 

1200 psig at 300°C. ............................................................................................................ 117

69. Rupture cross section for Studsvik Test 192 with irradiated ZIRLO pressurized to 

1200 psig at 300°C. The central scale bar is 8 mm. ........................................................... 118

70. Variation of CmR/(πDavgR) with δR/DavgR for AF, PH, and irradiated (Irrad.) ZIRLO 

subjected to pressurized ramp-to-rupture tests and LOCA oxidation tests. ....................... 120

71. Measured and calculated (Eq. 23) mid-plane circumferences in the rupture region for 

AF and irradiated (Irad.) ZIRLO following ramp-to-rupture and oxidation tests. ............. 120

72. Rupture strain vs. rupture temperature following ramp-to-rupture tests (AF ZIRLO) 

and post-rupture oxidation at ≈1200°C (AF and PH ZIRLO). ........................................... 122
 

xiv

73. Maximum 4-PBT bending moment (Mmax) vs. oxidation level (CP-ECR) for AF 

ZIRLO following ballooning, rupture, oxidation, and quench. .......................................... 128

74. Metallographic images of the OCZL#18 severed cross section following oxidation to 

12% CP-ECR, quench, and 4-point bending: (a) rupture tip (0.14-mm metal wall) 

and (b) back side of cross section (0.44-mm metal wall). .................................................. 129

75. Axial profile of hydrogen content in OCZL#18 sample after oxidation to 12% CPECR. ................................................................................................................................... 129

76. 4-PBT failure locations for the OCZL#19 sample after oxidation to 17% CP-ECR. ........ 130

77. Axial profiles for hydrogen and oxygen content along the OCZL#19 sample after 

oxidation to 17% CP-ECR and 4-point bending. ............................................................... 130

78. Maximum (failure) 4-PBT energy (Emax) vs. oxidation level (CP-ECR) for AF 

ZIRLO following ballooning, rupture, oxidation, and quench. .......................................... 131

79. Post-bend appearance of the OCZL#15 sample oxidized to 18% CP-ECR, quenched, 

and subjected to a reverse 4-PBT with the rupture tips in compression. ........................... 133

80. Post-bend-test images of the OCZL#17 sample oxidized to 13% CP-ECR, quenched, 

and subjected to bending at 135°C: (a) side view and (b) rupture view. ........................... 133

81. Post-bend-test images of the OCZL#29 sample (17% CP-ECR) after 4-PBTs 

conducted under reference conditions: (a) side view after 1st loading-unloading cycle 

and (b) rupture view after 2nd loading-unloading cycle. .................................................... 134

82. Maximum 4-PBT bending moment (Mmax) vs. oxidation level (CP-ECR) for AF and 

PH ZIRLO following ballooning, rupture, oxidation, and quench. ................................... 137

83. Maximum (failure) 4-PBT energy (Emax) vs. oxidation level (CP-ECR) for AF and 

PH ZIRLO following ballooning, rupture, oxidation, and quench. ................................... 138

84. Sectioning diagram and characterization results for the OCZL#39 post-bend sample 

with 660-wppm pre-test hydrogen oxidized to 13.2% CP-ECR. ....................................... 139

85. Sectioning diagram and characterization results for the OCZL#40 post-bend sample 

with 390-wppm pre-test hydrogen oxidized to 11.9% CP-ECR. ....................................... 139

86. Sectioning diagram and characterization results for the OCZL#41 post-bend sample 

with 220-wppm pre-test hydrogen oxidized to 15.6% CP-ECR. ....................................... 140

87. Sectioning diagram and characterization results for the OCZL#44 post-bend sample 

with 700-wppm pre-test hydrogen oxidized to 13.5% CP-ECR. ....................................... 140

88. Sectioning diagram and characterization results for the OCZL#45 post-bend sample 

with 220-wppm pre-test hydrogen oxidized to 13.2% CP-ECR. ....................................... 141
 

xv

89. Images for the post-bend OCZL#19 sample oxidized to 17.3% CP-ECR and 

quenched: (a) failure locations and (b) severed cross section at 24-mm below the 

rupture mid-span. ............................................................................................................... 143

A.1: Temperature histories for the OCZL#5 thermal benchmark test with unpressurized 

ZIRLO cladding filled with zirconia pellets. Sample was subjected to outer-surface 

oxidation with 40 ± 2 μm predicted and measured oxide layer thicknesses at the TC 

locations. ............................................................................................................................ A-1

A.2: Temperature and pressure histories for thermal benchmark test OCZL#6 pressurized 

to 1200 psig, oxidized to 18% CP-ECR, and cooled without quench. The control TC 

orientation was about 180° from the rupture opening. The high temperature recorded 

on the rupture side was confirmed by oxide-thickness measurements. ............................. A-2

A.3: Temperature and pressure histories for thermal-benchmark/4-point-bend test 

OCZL#7 pressurized to 800 psig, oxidized to 16% CP-ECR, and cooled without 

quench. The control TC orientation was about 180° from the rupture opening. ................ A-3

A.4: Temperature and pressure histories for ramp-to-rupture scoping test OCZL#8 

pressurized to 600 psig and cooled without quench. .......................................................... A-4

A.5: Temperature and pressure histories for ramp-to-rupture scoping test OCZL#9 

pressurized to 400 psig and cooled without quench. .......................................................... A-5

A.6: Temperature and pressure histories for ramp-to-rupture scoping test OCZL#10 

pressurized to 1600 psig and cooled without quench. ........................................................ A-6

A.7: Temperature and pressure histories for ramp-to-rupture scoping test OCZL#11 

pressurized to 1400 psig and cooled without quench. ........................................................ A-7

A.8: Temperature and pressure histories for scoping test OCZL#12 pressurized to 1000 

psig, oxidized at 1196°C, and cooled without quench. ...................................................... A-8

A.9: Temperature and pressure histories for scoping test OCZL#13 pressurized to 1200 

psig, oxidized at ≈1200°C to 15% CP-ECR , and cooled without quench. ........................ A-9

A.10: Temperature and pressure histories for data-generation test OCZL#14 pressurized to 

1200 psig, oxidized at ≈1200°C to 18% CP-ECR, and cooled with quench. ................... A-10

A.11: Temperature and pressure histories for data-generation test OCZL#15 pressurized to 

1200 psig, oxidized at ≈1200°C to 18% CP-ECR, and cooled with quench. ................... A-11

A.12: Temperature and pressure histories for data-generation test OCZL#16 pressurized to 

1200 psig, oxidized at ≈1200°C to 14% CP-ECR, and cooled with quench. ................... A-12

A.13: Temperature and pressure histories for data-generation test OCZL#17 pressurized to 

1200 psig, oxidized at ≈1200°C to13% CP-ECR, and cooled with quench. .................... A-13
 

xvi

A.14: Temperature and pressure histories for data-generation test OCZL#18 pressurized to 

1200 psig, oxidized at ≈1200°C to 12% CP-ECR, and cooled with quench. ................... A-14

A.15: Temperature and pressure histories for data-generation test OCZL#19 pressurized to 

600 psig, oxidized at ≈1200°C to 17% CP-ECR, and cooled with quench. ..................... A-15

A.16: Temperature and pressure histories for data-generation test OCZL#21 pressurized to 

600 psig, oxidized at ≈1200°C to 11% CP-ECR, and cooled with quench. ..................... A-16

A.17: Temperature and pressure histories for data-generation test OCZL#22 pressurized to 

600 psig, oxidized at ≈1200°C to 12% CP-ECR, and cooled with quench. ..................... A-17

A.18: Temperature and pressure histories for data-generation test OCZL#24 nonpressurized, one-sided oxidized at ≈1200°C to 17% CP-ECR, and cooled with 

quench. ............................................................................................................................. A-18

A.19: Temperature and pressure histories for data-generation test OCZL#25 pressurized to 

1200 psig, oxidized at ≈1200°C to 16% CP-ECR, and cooled with quench. ................... A-19

A.20: Temperature and pressure histories for data-generation test OCZL#29 pressurized to 

1200 psig, oxidized at ≈1200°C to 17% CP-ECR, and cooled with quench. ................... A-20

A.21: Temperature and pressure histories for data-generation test OCZL#32 pressurized to 

1200 psig, oxidized at ≈1200°C to 16% CP-ECR, and cooled with quench. ................... A-21

A.22: Temperature and pressure histories for thermal-benchmark test OCZL#36 

pressurized to 1200 psig, oxidized at ≈1200°C to 18% CP-ECR, and cooled with 

quench. ............................................................................................................................. A-22

A.23: Temperature and pressure histories for thermal/4-point-bend benchmark test 

OCZL#37 pressurized to 1200 psig, oxidized at ≈1200°C to 23% CP-ECR, and 

cooled with quench. ......................................................................................................... A-23

A.24: Temperature and pressure histories for data-generation test OCZL#43 pressurized to 

1200 psig, oxidized at ≈1200°C to 18% CP-ECR, and cooled with quench. ................... A-24

A.25: Temperature and pressure histories for data-generation test OCZL#39 with prehydrided (660 wppm) ZIRLO pressurized to 600 psig, oxidized at ≈1200°C to 13% 

CP-ECR, and cooled with quench. ................................................................................... A-25

A.26: Temperature and pressure histories for data-generation test OCZL#40 with prehydrided (390 wppm) ZIRLO pressurized to 1200 psig, oxidized at ≈1200°C to 12% 

CP-ECR, and cooled with quench. ................................................................................... A-26

A.27: Temperature and pressure histories for data-generation test OCZL#41 with prehydrided (220 wppm) ZIRLO pressurized to 1200 psig, oxidized at ≈1200°C to 16% 

CP-ECR, and cooled with quench. ................................................................................... A-27
 

xvii

A.28: Temperature and pressure histories for data-generation test OCZL#42 with prehydrided (520 wppm) ZIRLO pressurized to 1200 psig, oxidized at ≈1200°C to 14% 

CP-ECR, and cooled with quench. ................................................................................... A-28

A.29: Temperature and pressure histories for data-generation test OCZL#44 with prehydrided (700 wppm) ZIRLO pressurized to 1200 psig, oxidized at ≈1200°C to 14% 

CP-ECR, and cooled with quench. ................................................................................... A-29

A.30: Temperature and pressure histories for data-generation test OCZL#45 with prehydrided (220 wppm) ZIRLO pressurized to 1200 psig, oxidized at ≈1200°C to 13% 

CP-ECR, and cooled with quench. ................................................................................... A-30

B.1: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

thermal benchmark test OCZL#6, which was pressurized to 1200 psig, oxidized to 

18% CP-ECR at ≈1200°C, and cooled without quench. .................................................... B-1

B.2: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

scoping test OCZL#7, which was pressurized to 800 psig, oxidized to 16% CP-ECR 

at ≈1200°C, and cooled without quench. ........................................................................... B-2

B.3: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

ramp-to-rupture scoping test OCZL#8, which was pressurized to 600 psig and cooled 

without quench. .................................................................................................................. B-3

B.4: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

ramp-to-rupture scoping test OCZL#9, which was pressurized to 400 psig and cooled 

without quench. .................................................................................................................. B-4

B.5: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

ramp-to-rupture scoping test OCZL#10, which was pressurized to 1600 psig and 

cooled without quench. ...................................................................................................... B-5

B.6: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

ramp-to-rupture scoping test OCZL#11, which was pressurized to 1400 psig and 

cooled without quench. ...................................................................................................... B-6

B.7: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

test OCZL#12, which was pressurized to 1000 psig, oxidized to 15% CP-ECR, and 

cooled without quench. ...................................................................................................... B-7

B.8: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

test OCZL#13, which was pressurized to 1200 psig, oxidized to 14% CP-ECR, and 

cooled without quench. ...................................................................................................... B-8

B.9: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#14, which was pressurized to 1200 psig, oxidized to 18% 

CP-ECR, and cooled with quench. ..................................................................................... B-9
 

xviii

B.10: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#15, which was pressurized to 1200 psig, oxidized to 18% 

CP-ECR, and cooled with quench. ................................................................................... B-10

B.11: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#16, which was pressurized to 1200 psig, oxidized to 14% 

CP-ECR, and cooled with quench. ................................................................................... B-11

B.12: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#17, which was pressurized to 1200 psig, oxidized to 13% 

CP-ECR, and cooled with quench. ................................................................................... B-12

B.13: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#18, which was pressurized to 1200 psig, oxidized to 12% 

CP-ECR, and cooled with quench. ................................................................................... B-13

B.14: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#19, which was pressurized to 600 psig, oxidized to 17% 

CP-ECR, and cooled with quench. ................................................................................... B-14

B.15: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#21, which was pressurized to 600 psig, oxidized to 11% 

CP-ECR, and cooled with quench. ................................................................................... B-15

B.16: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#22, which was pressurized to 600 psig, oxidized to 12% 

CP-ECR, and cooled with quench. ................................................................................... B-16

B.17: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#25, which was pressurized to 1200 psig, oxidized to 16% 

CP-ECR, and cooled with quench. ................................................................................... B-17

B.18: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#29, which was pressurized to 1200 psig, oxidized to 17% 

CP-ECR, and cooled with quench. ................................................................................... B-18

B.19: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#32, which was pressurized to 1200 psig, oxidized to 16% 

CP-ECR, and cooled with quench. ................................................................................... B-19

B.20: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#37, which was pressurized to 1200 psig, oxidized to 23% 

CP-ECR, and cooled with quench. ................................................................................... B-20

B.21: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#43, which was pressurized to 1200 psig, oxidized to 18% 

CP-ECR, and cooled with quench. ................................................................................... B-21
 

xix

B.22: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#39 with 660 wppm pre-hydriding, which was pressurized 

to 600 psig, oxidized to 13% CP-ECR, and cooled with quench. .................................... B-22

B.23: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#40 with 390 wppm pre-hydriding, which was pressurized 

to 1200 psig, oxidized to 12% CP-ECR, and cooled with quench. .................................. B-23

B.24: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#41 with 220 wppm pre-hydriding, which was pressurized 

to 1200 psig, oxidized to 16% CP-ECR, and cooled with quench. .................................. B-24

B.25: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#42 with 520 wppm pre-hydriding, which was pressurized 

to 1200 psig, oxidized to 14% CP-ECR, and cooled with quench. .................................. B-25

B.26: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#44 with 700 wppm pre-hydriding, which was pressurized 

to 1200 psig, oxidized to 14% CP-ECR, and cooled with quench. .................................. B-26

B.27: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following 

data-generation test OCZL#45 with 220 wppm pre-hydriding, which was pressurized 

to 1200 psig, oxidized to 13% CP-ECR, and cooled with quench. .................................. B-27

C.1: Post-oxidation images of OCZL#6 sample, which was pressurized to 1200 psig, 

oxidized to 18% CP-ECR at ≈1200°C, and cooled without quench: (a) ruptureopening view and (b) side view. ........................................................................................ C-1

C.2: Post-oxidation images of OCZL#7 sample, which was pressurized to 800 psig, 

oxidized to 16% CP-ECR at ≈1200°C, and cooled without quench: (a) ruptureopening view and (b) side view. ........................................................................................ C-2

C.3: Post-oxidation images of OCZL#8 sample, which was pressurized to 600 psig, 

heated to rupture, and cooled without quench: (a) rupture-opening view and (b) side 

view. ................................................................................................................................... C-3

C.4: Post-oxidation images of OCZL#9 sample, which was pressurized to 400 psig, 

heated to rupture, and cooled without quench: (a) rupture-opening view and (b) side 

view. ................................................................................................................................... C-4

C.5: Post-oxidation images of OCZL#10 sample, which was pressurized to 1600 psig, 

heated to rupture, and cooled without quench: (a) low magnification and (b) higher 

magnification. .................................................................................................................... C-5

C.6: Post-oxidation images of OCZL#11 sample, which was pressurized to 1400 psig, 

heated to rupture, and cooled without quench: (a) low magnification and (b) higher 

magnification. .................................................................................................................... C-6
 

xx

C.7: Post-oxidation images of OCZL#12 sample, which was pressurized to 1000 psig, 

heated to rupture to 1200°C, oxidized to 15% CP-ECR, and cooled without quench: 

(a) low magnification and (b) higher magnification. ......................................................... C-7

C.8: Post-oxidation images of OCZL#13 sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 14% CP-ECR, and cooled without quench: (a) low 

magnification and (b) higher magnification. ...................................................................... C-8

C.9: Post-oxidation images of OCZL#14 sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 18% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. ...................................................................... C-9

C.10: Post-oxidation images of OCZL#15 sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 18% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-10

C.11: Post-oxidation images of OCZL#17 sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 13% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-11

C.12: Post-oxidation images of OCZL#18 sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 12% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-12

C.13: Post-oxidation images of OCZL#19 sample, which was pressurized to 600 psig, 

heated to 1200°C, oxidized to 17% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-13

C.14: Post-oxidation images of OCZL#21 sample, which was pressurized to 600 psig, 

heated to 1200°C, oxidized to 11% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-14

C.15: Post-oxidation images of OCZL#22 sample, which was pressurized to 600 psig, 

heated to 1200°C, oxidized to 12% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-15

C.16: Post-oxidation images of OCZL#24 sample, which was non-pressurized, heated to 

1200°C,one-sided oxidized to 17% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-16

C.17: Post-oxidation images of OCZL#25 sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 16% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-17

C.18: Post-oxidation images of OCZL#26 sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 16% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-18
 

xxi

C.19: Post-oxidation images of OCZL#27 sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 18% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-19

C.20: Post-oxidation images of OCZL#29 sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 17% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-20

C.21: Post-oxidation images of OCZL#32 sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 16% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-21

C.22: Post-oxidation images of OCZL#36 sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 18% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-22

C.23: Post-oxidation images of OCZL#37 sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 23% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-23

C.24: Post-oxidation images of OCZL#43 sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 18% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... C-24

C.25: Post-oxidation images of OCZL#39PH sample (660 wppm H), which was 

pressurized to 600 psig, heated to 1200°C, oxidized to 13% CP-ECR, and cooled 

with quench: (a) low magnification and (b) higher magnification. .................................. C-25

C.26: Post-oxidation images of OCZL#40 sample (390 wppm H), which was pressurized to 

1200 psig, heated to 1200°C, oxidized to 12% CP-ECR, and cooled with quench: (a) 

low magnification and (b) higher magnification. ............................................................. C-26

C.27: Post-oxidation images of OCZL#41 sample (220 wppm H), which was pressurized to 

1200 psig, heated to 1200°C, oxidized to 16% CP-ECR, and cooled with quench: (a) 

low magnification and (b) higher magnification. ............................................................. C-27

C.28: Post-oxidation images of OCZL#42 sample (520 wppm H), which was pressurized to 

1200 psig, heated to 1200°C, oxidized to 14% CP-ECR, and cooled with quench: (a) 

low magnification and (b) higher magnification. ............................................................. C-28

C.29: Post-oxidation images of OCZL#44 sample (700 wppm H), which was pressurized to 

1200 psig, heated to 1200°C, oxidized to 14% CP-ECR, and cooled with quench: (a) 

low magnification and (b) higher magnification. ............................................................. C-29

C.30: Post-oxidation images of OCZL#45 sample (220 wppm H), which was pressurized to 

1200 psig, heated to 1200°C, oxidized to 13% CP-ECR, and cooled with quench: (a) 

low magnification and (b) higher magnification. ............................................................. C-30
 

xxii

D.1: 4-PBT load-displacement curve for OCZL#8 test sample, which was pressurized to 

600 psig, heated to rupture, and cooled without quench.. .................................................. D-1

D.2: 4-PBT load-displacement curve for OCZL#9 test sample, which was pressurized to 

400 psig, heated to rupture, and cooled without quench. ................................................... D-2

D.3: 4-PBT load-displacement curve for OCZL#10 test sample, which was pressurized to 

400 psig, heated to rupture, and cooled without quench. ................................................... D-3

D.4: 4-PBT load-displacement curve for OCZL#12 test sample, which was pressurized to 

1000 psig, heated to ≈1200°C, oxidized to 15% CP-ECR, and cooled without 

quench. ............................................................................................................................... D-4

D.5: 4-PBT load-displacement curve for OCZL#13 test sample, which was pressurized to 

1200 psig, heated to ≈1200°C, oxidized to 14% CP-ECR, and cooled without 

quench. ............................................................................................................................... D-5

D.6: 4-PBT load-displacement curve for OCZL#14 test sample, which was pressurized to 

1200 psig, heated to 1200°C, oxidized to 18% CP-ECR, and cooled with quench. .......... D-6

D.7a: 4-PBT load-displacement curve for OCZL#15 test sample, which was pressurized to 

1200 psig, heated to 1200°C, oxidized to 18% CP-ECR, and cooled with quench. 

Sample was subjected to reverse bending with rupture region under maximum 

compression. Test was interrupted following non-linear displacement to examine 

sample. ............................................................................................................................... D-7

D.7b: Continuation of 4-PBT load-displacement curve for OCZL#15 test sample, which 

was pressurized to 1200 psig, heated to 1200°C, oxidized to 18% CP-ECR, and 

cooled with quench. Sample was subjected to reverse bending with rupture region 

under maximum compression. Results are for reloading sample. ...................................... D-8

D.8: 4-PBT load-displacement curve for OCZL#17 test sample, which was pressurized to 

1200 psig, heated to ≈1200°C, oxidized to 13% CP-ECR, and cooled with quench. ........ D-9

D.9: 4-PBT load-displacement curve for OCZL#18 test sample, which was pressurized to 

1200 psig, heated to ≈1200°C, oxidized to 12% CP-ECR, and cooled with quench. ...... D-10

D.10: 4-PBT load-displacement curve for OCZL#19 test sample, which was pressurized to 

600 psig, heated to ≈1200°C, oxidized to 17% CP-ECR, and cooled with quench. ........ D-11

D.11: 4-PBT load-displacement curve for OCZL#21 test sample, which was pressurized to 

600 psig, heated to ≈1200°C, oxidized to 11% CP-ECR, and cooled with quench. ........ D-12

D.12: 4-PBT load-displacement curve for OCZL#22 test sample, which was pressurized to 

600 psig, heated to ≈1200°C, oxidized to 12% CP-ECR, and cooled with quench. ........ D-13

D.13: 4-PBT load-displacement curve for OCZL#24 test sample, which was nonpressurized, heated to ≈1200°C, one-sided oxidized to 17% CP-ECR, and cooled 

with quench. ..................................................................................................................... D-14
 

xxiii

D.14: 4-PBT load-displacement curve for OCZL#25 test sample, which was pressurized to 

1200 psig, heated to ≈1200°C, oxidized to 16% CP-ECR, and cooled with quench. ...... D-15

D.15a: 4-PBT load-displacement curve for OCZL#29 test sample, which was pressurized to 

1200 psig, heated to ≈1200°C, oxidized to 17% CP-ECR, and cooled with quench. 

Test was interrupted to allow examination of sample. ..................................................... D-16

D.15b: 4-PBT load-displacement curve for OCZL#29 test sample, which was pressurized to 

1200 psig, heated to ≈1200°C, oxidized to 17% CP-ECR, and cooled with quench. 

Sample was reloaded after visual examination. ............................................................... D-17

D.16: 4-PBT load-displacement curve for OCZL#32 test sample, which was pressurized to 

1200 psig, heated to ≈1200°C, oxidized to 16% CP-ECR, and cooled with quench. ...... D-18

D.17: 4-PBT load-displacement curve for OCZL#36 test sample, which was pressurized to 

1200 psig, heated to ≈1200°C, oxidized to 18% CP-ECR, and cooled with quench. ...... D-19

D.18: 4-PBT load-displacement curve for OCZL#37 test sample, which was pressurized to 

1200 psig, heated to ≈1200°C, oxidized to 23% CP-ECR, and cooled with quench. ...... D-20

D.19: 4-PBT load-displacement curve for OCZL#43 test sample, which was pressurized to 

1200 psig, heated to ≈1200°C, oxidized to 18% CP-ECR, and cooled with quench. ...... D-21

D.20: 4-PBT load-displacement curve for OCZL#39 (660 wppm H) test sample, which was 

pressurized to 600 psig, heated to ≈1200°C, oxidized to 13% CP-ECR, and cooled 

with quench. ..................................................................................................................... D-22

D.21: 4-PBT load-displacement curve for OCZL#40 (390 wppm H) test sample, which was 

pressurized to 1200 psig, heated to ≈1200°C, oxidized to 12% CP-ECR, and cooled 

with quench. ..................................................................................................................... D-23

D.22: 4-PBT load-displacement curve for OCZL#41 (220 wppm H) test sample, which was 

pressurized to 1200 psig, heated to ≈1200°C, oxidized to 16% CP-ECR, and cooled 

with quench. ..................................................................................................................... D-24

D.23a: 4-PBT load-displacement curve for OCZL#44 (700 wppm H) test sample, which was 

pressurized to 1200 psig, heated to ≈1200°C, oxidized to 14% CP-ECR, and cooled 

with quench. Test was interrupted to allow for visual examination. ................................ D-25

D.23b: 4-PBT load-displacement curve for OCZL#44 (700 wppm H) test sample, which was 

pressurized to 1200 psig, heated to ≈1200°C, oxidized to 14% CP-ECR, and cooled 

with quench. Sample was reloaded. ................................................................................. D-26

D.24: 4-PBT load-displacement curve for OCZL#45 (220 wppm H) test sample, which was 

pressurized to ≈1200 psig, heated to ≈1200°C, oxidized to 13% CP-ECR, and cooled 

with quench. ..................................................................................................................... D-27
 

xxiv

E.1: Post-4-PBT images of ramp-to-rupture samples from OCZL#8, #9, and #10 test 

samples, which were pressurized at 300°C to 600, 400, and 1600 psig, respectively, 

heated to rupture, and cooled without quench. ................................................................... E-1

E.2: Post-4-PBT images of OCZL#12 test sample, which was pressurized to 1000 psig, 

heated to ≈1200°C, oxidized to 15% CP-ECR, and cooled without quench: (a) low 

magnification and (b) higher magnification. ...................................................................... E-2

E.3: Post-4-PBT image of OCZL#13 test sample, which was pressurized to 1200 psig, 

heated to ≈1200°C, oxidized to 14% CP-ECR, and cooled without quench. ..................... E-3

E.4: Post-4-PBT images of OCZL#14 test sample, which was pressurized to 1200 psig, 

heated to ≈1200°C, oxidized to 18% CP-ECR, and cooled with quench. .......................... E-3

E.5: Post-4-PBT images for the OCZL#15 test sample, which was pressurized to 1200 

psig, heated to ≈1200°C, oxidized to 18% CP-ECR, cooled with quench, and 

subjected to reverse bending: (a) low magnification and (b) higher magnification. .......... E-4

E.6: Post-4-PBT images of OCZL#17 test sample, which was pressurized to 1200 psig, 

heated to ≈1200°C, oxidized to 13% CP-ECR, and cooled with quench: (a) side view 

and (b) rupture view. .......................................................................................................... E-5

E.7: Post-4-PBT images of OCZL#18 test sample, which was pressurized to 1200 psig, 

heated to ≈1200°C, oxidized to 12% CP-ECR, and cooled with quench: (a) side view 

and (b) rupture view. .......................................................................................................... E-6

E.8: Post-4-PBT images of OCZL#19 test sample, which was pressurized to 600 psig, 

heated to ≈1200°C, oxidized to 17% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. ...................................................................... E-7

E.9: Post-4-PBT images of OCZL#21 test sample, which was pressurized to 600 psig, 

heated to ≈1200°C, oxidized to 11% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. ...................................................................... E-8

E.10: Post-4-PBT images of OCZL#22 test sample, which was pressurized to 600 psig, 

heated to ≈1200°C, oxidized to 12% CP-ECR, and cooled with quench: (a) low 

magnification and (b) high magnification. ......................................................................... E-9

E.11: Post-4-PBT images of OCZL#24 test sample, which was non-pressurized, heated to 

≈1200°C, one-sided oxidized to 17% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... E-10

E.12: Post-4-PBT images of OCZL#25 test sample, which was pressurized to 1200 psig, 

heated to ≈1200°C, oxidized to 16% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... E-11

E.13a: Post-4-PBT images of OCZL#29 test sample, which was pressurized to 1200 psig, 

heated to ≈1200°C, oxidized to 17% CP-ECR, and cooled with quench. Images were 

taken after first loading: (a) side A and (b) side B. .......................................................... E-12
 

xxv

E.13b: Post-4-PBT image of OCZL#29 test sample, which was pressurized to 1200 psig, 

heated to ≈1200°C, oxidized to 17% CP-ECR, and cooled with quench. Image was 

taken after second loading. ............................................................................................... E-13

E.14: Post-4-PBT Images of OCZL#32 test sample, which was pressurized to 1200 psig, 

heated to ≈1200°C, oxidized to 16% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... E-14

E.15: Post-4-PBT images of OCZL#36 test sample, which was pressurized to 1200 psig, 

heated to ≈1200°C, oxidized to 18% CP-ECR, and cooled with quench: (a) low 

magnification and (b) high magnification. ....................................................................... E-15

E.16: Post-4-PBT images of OCZL#37 test sample, which was pressurized to 1200 psig, 

heated to ≈1200°C, oxidized to 23% CP-ECR, and cooled with quench: (a) low 

magnification and (b) high magnification. ....................................................................... E-16

E.17: Post-4-PBT images of OCZL#43 test sample, which was pressurized to 1200 psig, 

heated to 1200°C, oxidized to 18% CP-ECR, and cooled with quench: (a) low 

magnification and (b) higher magnification. .................................................................... E-17

E.18: Post-4-PBT images of OCZL#39 (660 wppm H) test sample, which was pressurized 

to 600 psig, heated to ≈1200°C, oxidized to 13% CP-ECR, and cooled with quench: 

(a) low magnification and (b) higher magnification. ....................................................... E-18

E.19: Post-4-PBT images of OCZL#40 (390 wppm H) test sample, which was pressurized 

to 1200 psig, heated to ≈1200°C, oxidized to 12% CP-ECR, and cooled with quench: 

(a) low magnification and (b) higher magnification. ....................................................... E-19

E.20: Post-4-PBT images of OCZL#41 (220 wppm H) test sample, which was pressurized 

to 1200 psig, heated to ≈1200°C, oxidized to 16% CP-ECR, and cooled with quench: 

(a) low magnification and (b) higher magnification. ....................................................... E-20

E.21: Post-4-PBT images of OCZL#44 (700 wppm H) test sample, which was pressurized 

to 1200 psig, heated to ≈1200°C, oxidized to 14% CP-ECR, and cooled with quench: 

(a) low magnification and (b) higher magnification. ....................................................... E-21

E.22: Post-4-PBT images of OCZL#45 (220 wppm H) test sample, which was pressurized 

to 1200 psig, heated to ≈1200°C, oxidized to 13% CP-ECR, and cooled with quench: 

(a) low magnification and (b) higher magnification. ....................................................... E-22

 

xxvii

TABLES 

1. Dimensions and compositions of PWR cladding alloys used in test program; OD is outer 

diameter and h is wall thickness. ......................................................................................................... 2

2. Dimensions and Chemistry of Zr-lined Zry-2 Used in the Argonne Test Program (the “<” 

sign means below the detection limit and “---” means not measured). ............................................. 13

3. Dimensions and Chemistry of Zry-4 Used in the Argonne Test Program (the "<" sign means 

below the detection limit and “---” means not measured). ................................................................ 13

4. Dimensions and chemistry of ZIRLO and M5 cladding materials used in the Argonne test 

program (the "<" sign means below the detection limit “---” means not measured)......................... 14

5. Characterization of the four lots of ZIRLO used in the Argonne test program. ............................... 14

6. Characteristics of high-burnup fuel rod segments and defueled high-burnup cladding for the 

Argonne LOCA test program. Burnup values are rod averaged. ...................................................... 15

7. Ductile-to-brittle transition oxidation levels (embrittlement CP-ECR values) for as-fabricated 

cladding alloys oxidized at 1200°C and ring-compressed at 135°C. Most data points used to 

determine ductility were for samples quenched at 800°C. ................................................................ 32

8. Embrittlement pre-test hydrogen content (CHM) as a function of oxidation level (CP-ECR) for 

pre-hydrided Zry-4 cladding oxidized at ≈1200oC and quenched at 800oC. ..................................... 33

9. Pre-test estimates of CHM levels, measured dimensions, and measured weights for prehydrided ZIRLO oxidation test samples. .......................................................................................... 36

10. Oxidation conditions and results for pre-hydrided ZIRLO samples. Wg is the measured 

change in sample weight (in %) and wg is weight gain normalized to surface area. ......................... 39

11. Oxidation levels, RCT results, and CHM values for pre-hydrided ZIRLO samples. .......................... 39

12. Analysis of PQD data to determine ductile-to-brittle transition CP-ECR and CHM values. .............. 40

13. Comparison between pre-test estimated CHM values for oxidation samples and CHM values for 

RCT samples sectioned from oxidation test samples and determined from measured postoxidation (CH)POX values. .................................................................................................................. 42

14. RCT results for HBU Zry-4 samples oxidized at ≤1200°C and cooled without quench. CHM

was ≈550±100 wppm for these samples............................................................................................ 45
 

xxviii

15. Determination of CHM based on post-oxidation hydrogen measurements [(CH)POX] for highburnup M5 samples used for oxidation tests; fWg is post-test-weight/pre-test-weight and fMM

is pre-test sample-mass/metal-mass. Sample masses are listed in parentheses. ................................ 46

16. RCT results for HBU M5 samples oxidized at 1200°C and cooled without and with quench 

(see Table 74 in Ref. 1). .................................................................................................................... 47

17. Determination of CHM based on measured post-oxidation hydrogen content for high-burnup 

ZIRLO samples used for oxidation tests and PQD determination; fWg = post-test-weight/pretest-weight, fMM = pre-test sample-mass/metal-mass. Sample masses are given in parentheses. ..... 47

18. RCT results for high-burnup ZIRLO samples oxidized at <1200°C and cooled with quench at 

800°C. ............................................................................................................................................... 48

19. Characterization of high-burnup ZIRLO segments used for current oxidation tests. ....................... 52

20. Post-test ductility results for high-burnup ZIRLO cladding oxidized at ≤1200oC, cooled with 

or without (SC) quench at 800oC, and ring-compressed at 135oC. ................................................... 57

21. Minimum breakaway oxidation times and corresponding oxidation temperatures reported in 

Ref. 1 for as-fabricated cladding alloys. ........................................................................................... 71

22. Results of ZIRLO breakaway oxidation tests at 1000oC; new results are in bold font. .................... 77

23. Results of ZIRLO breakaway oxidation tests at 970-985oC; new results are in bold font. ............... 84

24. Characterization of 17×17 M5 cladding used in the ANL testing program. ..................................... 86

25. Results of previous M5 oxidation tests at 1000oC in long test train. ................................................ 87

26. M5-2007 breakaway oxidation test results for long-time hold temperatures of 980oC and 

985oC in the short test train. .............................................................................................................. 88

27. Results of ZIRLO-2008 transient breakaway-oxidation tests. ........................................................ 101

28. Summary of results for LOCA integral and post-LOCA-bend scoping tests with AF ZIRLO 

cladding. Fill pressures were at 300°C. Oxidation tests were conducted with a target hold 

temperature of 1200°C and without quench. 4-PBTs were conducted at 135°C and 2 mm/s 

displacement rate to 14-mm maximum displacement. Failure locations are referenced to the 

mid-span of the rupture opening. .................................................................................................... 111

29. Summary of data for cross sections at rupture mid-spans for ZIRLO samples subjected to 

ramp-to-rupture tests and samples subjected to LOCA oxidation tests. ......................................... 119

30. Internal pressure (Pi[300°C]), rupture strain (ΔCmR/Cmi, measured values in bold font), 

rupture temperature (TR), wall thickness (hR), and CP-ECR for AF and PH ZIRLO tests. ............ 121

31. Summary of LOCA integral post-quench parameters that may affect AF-ZIRLO cladding 

performance during 4-PBTs. ........................................................................................................... 125
 

xxix

32. Summary of results for LOCA integral tests terminated with quench at 800°C and postquench 4-PBTs with AF ZIRLO cladding. The 4-PBT displacement rate was decreased from 

2 mm/s to 1 mm/s after OCZL#21 sample test. .............................................................................. 126

33. Summary of LOCA integral post-quench parameters that may affect PH-ZIRLO cladding 

performance during 4-PBTs. ........................................................................................................... 136

34. Summary of results for LOCA integral tests terminated with quench at 800°C and postquench 4-PBTs with PH ZIRLO cladding. 4-PBT displacement rate was 1 mm/s. ........................ 137

35. Dimensions, properties, and failure bending moment at severed location for the OCZL#19 

test sample with 17% CP-ECR in the rupture node and 12% measured ECR at the severed 

location 24-mm below the rupture mid-span. ................................................................................. 144

 

xxxi

EXECUTIVE SUMMARY

Previous loss-of-coolant-accident (LOCA) acceptance criteria in 10 CFR 50.46(b) limited the peak 

cladding temperature to 2200°F (1204°C) and the maximum oxidation level to 17% of the cross-sectionaveraged cladding wall thickness. It had been assumed that ductility would be maintained in Zircaloys 

(Zr-Sn alloys) and ZIRLO (Zr-Sn-Nb alloy) if these criteria were satisfied. Data presented previously 

(NUREG/CR-6967) showed that the 17% oxidation level is not conservative for irradiated Zircaloy-4, 

ZIRLO, and M5 (Zr-Nb alloy) because the embrittlement oxidation level decreases with increasing 

hydrogen content picked up during normal operation. These data also showed that embrittlement can 

occur at even lower oxidation levels and lower temperatures due to excessive hydrogen pickup if 

breakaway oxidation occurs. 

In work performed since publication of NUREG/CR-6967, the following were accomplished: (a) 

methodologies were improved for determining pre-transient hydrogen content and post-quench ductility, 

(b) high-temperature oxidation limits were established for high-exposure and pre-hydrided cladding with 

intermediate levels of pre-transient hydrogen (200-340 wppm), (c) minimum breakaway oxidation times 

were determined for as-fabricated ZIRLO and M5 cladding under near-isothermal conditions, (d) 

minimum transient breakaway oxidation times were determined for as-fabricated ZIRLO, (e) LOCA 

integral tests were conducted with as-fabricated and pre-hydrided ZIRLO to determine balloon size and 

rupture temperature, and (f) post-test integral samples were subjected to four-point bend tests (4-PBTs) to 

determine failure limits as functions of pre-transient hydrogen content and maximum oxidation level. 

High-temperature embrittlement tests were conducted with pre-hydrided (180 to 280 wppm) and highexposure (300 to 340 wppm) ZIRLO cladding oxidized to 8 – 11% CP-ECR at peak temperatures of 

1200°C, quenched at 800°C, cooled to room temperature, and subjected to ring-compression tests at 

135°C. CP-ECR is Equivalent Cladding Reacted (ECR) calculated using the Cathcart-Pawel (CP) weight 

gain correlation. It is used throughout this work to represent the oxidation level of cladding alloys. For 

pre-hydrided ZIRLO, ductile-to-brittle transition oxidation levels were 11% and 10% for pre-transient 

hydrogen contents of 200 wppm and 240 wppm, respectively. For high-exposure ZIRLO, the transition 

oxidation level was 9% for 300-wppm hydrogen and <8% for 340-wppm hydrogen. These transition 

oxidation levels correspond to borderline ductility as defined by the permanent-strain criterion (≥1%) 

and/or the corresponding offset-strain criterion. Past and current results are summarized in the following 

figure. The consistency in results for different cladding alloys and for pre-hydrided and high-exposure 

cladding materials indicates that the decrease in embrittlement threshold is due primarily to the pretransient hydrogen content. No effects of alloy differences or irradiation were observed. The results also 

indicate that pre-hydrided cladding is a good surrogate for high-burnup cladding. 

Based on previous results for as-fabricated ZIRLO cladding, the minimum breakaway oxidation time 

(3100±300 s) occurred for average hold temperatures of 970°C to 985°C. The criterion used by Argonne 

to determine breakaway-oxidation initiation is 200-wppm hydrogen pickup. The breakaway time was 

longer (4000±200 s) for 1000°C-oxidation tests. Using a different test apparatus, test methods, and 

breakaway criterion, Westinghouse (WEC) observed post-breakaway surface discoloration on ZIRLO 

samples after ≈4400 s at 1000°C. However, WEC data indicated no surface-oxide discoloration after 

≈4400 s at <1000°C. Additional tests were conducted by Argonne and WEC using samples from the same 

cladding lots. The same differences were observed at <1000°C. These apparent differences may be due to 

the high temperature sensitivity of ZIRLO breakaway oxidation. Circumferential temperature variations 

in Argonne’s radiant-heating furnace resulted in oxidation of samples at all temperatures from 970°C to 

1000°C. WEC tests were conducted in a more homogeneous resistance furnace at discrete temperatures of 

960°C, 980°C, and 990°C. 
 

xxxii

Embrittlement oxidation level as a function of pre-transient hydrogen content for peak oxidation 

temperatures of 1200°C for <400 wppm and 1130−1180°C for >400 wppm hydrogen. 

Argonne also conducted four tests with as-fabricated ZIRLO oxidized under transient conditions, which 

included single step changes up and down in temperature and temperature cycling. The minimum 

breakaway oxidation time was determined to be 2800 s for these transient temperature tests. This time is 

within the data scatter for isothermal tests and the same as the breakaway oxidation time for a mildly 

scratched (20-microns deep) ZIRLO sample subjected to isothermal oxidation. These results suggest that 

temperature transients do not reduce the minimum breakaway oxidation time measured in isothermal 

tests. Two breakaway tests were conducted with as-fabricated M5 cladding oxidized at 985°C for 5000 s. 

Breakaway oxidation was not observed at 5000 s, nor was it observed in previous tests with M5 cladding 

oxidized at 1000°C for 4100 s. These results are consistent with CEA/Saclay results (≈6400 s breakaway 

time) for 1000°C-oxidized samples. 

Long (≈300 mm), pressurized, as-fabricated and pre-hydrided ZIRLO samples were subjected to 

controlled heating at 5°C/s from 300°C to 1200°C, oxidation at 1200°C for various times, controlled 

cooling at 3°C/s to 800°C, and quench at 800°C. All samples ballooned and ruptured during the heating 

ramp. Rupture temperatures and hence ballooning strains for as-fabricated ZIRLO varied with the internal 

gas pressure, which remained relatively constant from 300°C up to the rupture temperature: (a) 843±6°C 

and 25±3% strain for 4.14 MPa and (b) 749±7°C and 46±4% strain for 8.28 MPa. Ballooning, rupture, 

and subsequent oxidation processes resulted 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. The effects of prehydriding (220 to 700 wppm) ZIRLO cladding prior to the heating transient were opposite to what was 

expected based on the effect of hydrogen on phase changes. At 8.28 MPa fill pressure, rupture 

temperatures were lower (670°C to 740°C) than anticipated and rupture strains were consequently higher 

(47% to 71%) than anticipated. These results suggest that hydrogen has a more significant effect on 

ballooning strains due to weakening of the cladding (resulting in lower rupture temperatures) than on 

shifting the phase-change temperature.

0

2

4

6

8

10

12

14

16

18

20

0 100 200 300 400 500 600 700 800

Embrittlement Oxidation Level (%)

Hydrogen Content (wppm)

15x15 Vintage Zry-4

15x15 Low-Sn Zry-4

17x17 Low-Sn Zry-4

10x10 Zr-lined Zry-2

17x17 ZIRLO

17x17 M5

 

 

63 GWd/t

70 GWd/t

70 GWd/t

64 GWd/t

Ductile

Brittle

Previous 10CFR50.46 Limit
 

xxxiii

Four-point bend tests (4-PBTs) were used to determine post-quench failure location, maximum bending 

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

of plastic deformation). For 4-PBT loading, the bending moment is constant between the load application 

points, which were spaced far enough apart (150 mm) to include the ballooned region. With analysis, the 

4-PBT bending moment can be related to an equivalent axial tensile load. Non-oxidized, ballooned and 

ruptured samples were strong and highly ductile, but oxidized samples exhibited decreasing strength with 

increasing oxidation level and pre-transient hydrogen content. However, no clear pattern was exhibited 

with regard to plastic deformation; most samples with high ballooning strain severed through the rupture 

cross section prior to accumulating offset strain. Thus, even though the rupture cross section contained 

highly ductile regions, cracks initiated in the brittle region (near rupture tips) and propagated rapidly 

through the locally ductile section of the cladding (i.e., near the thick back wall). For low ballooning 

strains, failure (i.e., severing of the cross section) occurred in the regions between the edge of the rupture 

opening and the balloon neck. 

Strength results are shown in the following figure. Data include samples that failed within the rupture 

region and outside the rupture region. At 17% CP-ECR, the linearized fit to the data for as-fabricated 

ZIRLO (round blue dots) gives 6.4 N•m maximum bending moment, which corresponds to an axialrestraint failure load of ≈2600 N. This value is higher than maximum loads measured by the Japanese 

Atomic Energy Agency at Tokai (1200-2400 N) for ballooned, ruptured, oxidized and quenched Zircaloy4 samples that survived conditions of full restraint during quench from 700°C to ≈100°C. Thus, asfabricated ZIRLO cladding restricted to ≤17% CP-ECR would survive this severe loading. The strength 

of ZIRLO with 220-wppm hydrogen following oxidation to ≈13% CP-ECR (round red dot) is comparable 

to the strength of as-fabricated ZIRLO oxidized to 17% CP-ECR. Based on the oxidation embrittlement 

limit figure, the oxidation level would be limited to <12% for cladding with 220-wppm pre-transient 

hydrogen. Samples with more pre-transient hydrogen or higher oxidation levels had correspondingly 

lower strength. Thus, the hydrogen-dependent oxidation limit based on ring-compression tests (see 

previous figure) ensures preservation of cladding strength for the balloon region for as-fabricated 

cladding and cladding with pre-transient hydrogen, including irradiated cladding. 

Maximum bending moment for ZIRLO LOCA integral samples with and without (AF) the addition 

of pre-transient hydrogen. Results are from 4-PBTs conducted at 1-2 mm/s and 135°C.

0

2

4

6

8

10

12

14

16

9 11 13 15 17 19 21 23 25

Max. Bending Moment (N•m) 

Oxidation Level (%)

AF ZIRLO

220-wppm ZIRLO

390-wppm ZIRLO

660-wppm ZIRLO

700-wppm ZIRLO

 

xxxv

ACKNOWLEDGMENTS 

The authors would like to express their 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. Michelle Bales (nee Flanagan) provided significant technical guidance with 

regard to data needs relevant to licensing issues.

Argonne technicians Jakub Dobrzynski, David McGann, and Kevin Byrne are acknowledged for their 

careful efforts in test setup, test conduct, and post-test measurements. The authors would like to express 

their appreciation to international colleagues Jean-Christophe Brachet and Valerie Vandenberghe of the 

Commissariat a l’Energie Atomique (CEA) at Saclay, Claude Grandjean and Georges Hache of the 

Institut de Radioprotection et de Surete Nucleaire (IRSN), and Fumihisa Nagase of the Japanese Atomic 

Energy Agency (JAEA) for sharing details, data, and insights regarding test methods and embrittlement 

mechanisms.

Many organizations, and individuals within those organizations, have contributed test materials, reviewed 

test plans, participated in program review meetings, shared experiences, and provided both guidance and 

technical expertise. The Electric Power Research Institute (EPRI), Global Nuclear Fuels (GNF), AREVA, 

and Westinghouse Electric Company (WEC) provided high-burnup commercial reactor fuel rods, as well 

as defueled M5 (Studsvik and AREVA) and ZIRLO (Studsvik and WEC) cladding from high-burnup fuel 

rods, to the program. These organizations, along with others, also supplied us with as-fabricated cladding 

and tubing. In particular, WEC is acknowledged for providing the ZIRLO cladding tubes used in the 

current study. 

This report updates as necessary and includes all major results from the earlier work published in 

NUREG/CR-6967. Therefore, we also acknowledge those who contributed to that earlier report. 

 

 

xxxvii

ACRONYMS AND ABBREVIATIONS 

3-PBT 3-point bend test

4-PBT 4-point bend test 

AF As-fabricated

AGHCF Alpha-Gamma Hot-Cell Facility

ANL Argonne National Laboratory 

BJ Baker-Just 

BP Belt polished

Bu Burnup 

BWR Boiling water reactor

CEA Commissariat a l’Energie Atomique

CFR Code of Federal Regulations 

CP Cathcart-Pawel

CP-ECR ECR calculated using the CP weight-gain correlation 

ECR Equivalent cladding reacted

EdF Electricitie de France

EPRI Electric Power Research Institute

FEA Finite Element Analysis

GWd/t Gigawatt-days per metric tonne of uranium 

HBR H.B. Robinson reactors 

HBU High burnup 

ID Inner diameter

IML Irradiated Materials Laboratory

IRSN Institute for Radiological Protection and Nuclear Safety 

JAEA Japanese Atomic Energy Agency

LOCA Loss-of-coolant accident

NA North Anna reactors 

NIST National Institute of Standards and Testing 

NRC Nuclear Regulatory Commission

OCZL Out-of-cell oxidation test for ZIRLO

OD Outer diameter

ORNL Oak Ridge National Laboratory 

PCT Peak cladding temperature

PH Pre-hydrided 

PQD Post-quench ductility 

PWR Pressurized water reactor

R Universal gas constant

RCT Ring compression test

RES Office of Nuclear Regulatory Research

RT Room temperature 

TC Thermocouple 

TMI-1 Three Mile Island Unit 1 Reactor

W Westinghouse 

WC Westinghouse cleaned

WEC Westinghouse Electric Company

Zry-2 Zircaloy-2 

Zry-4 Zircaloy-4 

 

xxxix

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

CH Hydrogen content (in wppm) in sample measured prior to conducting oxidation tests 

CHi Hydrogen content (in wppm) in as-fabricated cladding

CHM Hydrogen content (in wppm) in cladding metal prior to conducting oxidation tests 

Cm Cladding mid-wall circumference

Cmf Post-rupture value of cladding mid-wall circumference

Cmi Pre-transient value of cladding mid-wall circumference

CoR Maximum post-rupture value of cladding metal outer circumference

CmR Maximum post-rupture value of cladding metal mid-wall circumference

(CH)POX Hydrogen content (in wppm) in sample measured after oxidation tests 

Coi Pre-transient value of cladding metal outer-wall circumference

co Average oxygen concentration in the beta (β) layer

coi Initial oxygen concentration in as-fabricated cladding 

dap Permanent displacement at loading positions in 4-PBTs

Davg Average diameter of ballooned cladding 

DavgR Average diameter of ballooned and rupture cladding at about mid-span of rupture opening

Di Inner diameter of cladding

Dmax Maximum diameter at a cross section of ballooned cladding

Dmi Inner diameter of cladding metal

Dmin Minimum diameter at a cross section of ballooned cladding

Dmo Outer diameter of cladding metal

Do Outer diameter of cladding

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 of cladding 

HB Total mass of hydrogen in fuel-cladding bond layer 

HC Total mass of hydrogen in corrosion layer 

hi Wall thickness of as-fabricated cladding

hR Minimum wall thickness (averaged over cross section) of ballooned and ruptured cladding

HM Total mass of hydrogen in metal layer 

hm Wall thickness of cladding metal

HS Total hydrogen mass in cladding sample 

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 
 

xl

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

k Temperature dependent coefficient 

K Stiffness; linearized loading slope of load-displacement curve

Kg Stiffness of gauge section in response to axial load

KLC Calculated loading stiffness

KLM Measured loading stiffness

Kmach Instron machine stiffness

L Length

LR Axial length of rupture opening for ballooned and ruptured cladding

Ls Distance between applied loads in 4-PBT

M Bending moment 

Ms Total sample mass

MB Mass of fuel-cladding bond layer 

MC Mass of corrosion layer

MM Mass of metal layer

Mmax Maximum bending moment 

P Load measured by load cell

Pmax Maximum load for 4-PBTs and JAEA tests

R Ideal gas constant

Rmid Mid-wall radius of cladding metal

Rmo Outer radius of cladding metal

So Solubility of oxygen in the beta layer

T Temperature

TR Rupture temperature 

W Weight (mass) of sample

WG Weight gain due to oxygen pickup in mass units 

Wg Weight gain due to oxygen pickup in % 

wg Weight gain due to oxygen pickup normalized to surface area (in g/cm2 or mg/cm2

) 

δ Applied displacement 

ΔCm/Cmi Circumferential strain at the mid-wall location ([Cmf – Cmi]/Cmi) 

ΔCHM Hydrogen pickup (in wppm) in cladding during oxidation normalized to pre-test metal mass 

δb Fuel-cladding bond layer thickness 

δc Corrosion layer thickness

δe Elastic displacement

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

(δox)i Thickness of inner-surface oxide

(δox)o Thickness of outer surface oxide

δR Maximum width of the rupture opening 

δp Offset displacement based on load-displacement curve

εp Plastic hoop strain

εθ Elastic hoop strain

εz Elastic axial strain

ν Poisson’s ratio

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

σmax Maximum axial stress in cladding metal for samples subjected to 4-PBTs and JAEA tests

σθ Elastic hoop stress

σz Elastic axial stress
 

1

1 INTRODUCTION

By the mid 1990s, it was well known that significant cladding oxidation (corrosion) can take place during 

extended periods of normal operation, and that some of the liberated hydrogen from the interaction with 

water is absorbed in the zirconium alloy cladding material. It thus seemed likely that an oxide layer on the 

surface and hydrogen in the metal would affect the behavior of cladding under conditions of a loss-ofcoolant accident (LOCA).

Several aspects of cladding behavior under LOCA conditions are involved in safety analyses: ballooning 

strains, flow area reduction, rupture conditions, axial distribution of heat sources, and cladding 

embrittlement. A previous report by this laboratory provided results on oxidation-embrittlement limits for

non-ballooned cladding as a function of pre-transient hydrogen content, on minimum times and 

corresponding temperatures for breakaway oxidation, and on ballooning with rupture under LOCA 

conditions [1]. The present report updates those results, provides additional results in those areas, and 

adds results of mechanical testing of cladding after ballooning, rupture, oxidation, and quench. 

Improved testing procedures and data-analysis methods have been developed and utilized in our work 

since the publication of Ref. 1. Step-by-step procedures for conducting breakaway-oxidation tests and 

post-quench-ductility tests, along with appropriate data-analysis methods, were provided to NRC and 

were then incorporated in regulatory guides [2-4]. In addition, an assessment of post-LOCA mechanical 

tests methods was performed to confirm that we were using the best available testing procedures [5].

Some of the material in the Introduction of Ref. 1 is repeated here in an effort to make the present report 

relatively self-sufficient. Although all the important results from Ref. 1 are summarized in the present 

report, the earlier report would have to be consulted for details. The background and regulatory 

significance of cladding behavior during postulated LOCAs can be found in Ref. 6. 

1.1 Cladding Materials

Table 1 lists the dimensions (outer diameter, OD, and wall thickness, h) and major alloying elements of 

pressurized-water-reactor (PWR) cladding materials that have been tested in this program. In addition to 

the PWR alloys included in Table 1, two sizes of boiling-water-reactor (BWR) cladding were used in the 

previous study: 9×9 (11.18-mm OD and 0.71-mm h) and 10×10 (10.29-mm OD and 0.66-mm h) Zr-lined 

Zircaloy-2 (Zry-2) cladding, where the inner lining of Zr constitutes ≈10% of the total cladding thickness. 

The chemical composition of Zry-2 is similar to standard Zircaloy-4 (Zry-4) except for the addition of 

≈0.06 wt. % Ni.

Wall thickness values of these cladding materials are important in the calculation of oxidation level as a 

percentage of wall thickness. The alloying elements have a significant impact on coolant-side corrosion 

and hydrogen pickup during normal operation and on phase-change temperatures during LOCA 

transients. In general, a reduction of Sn results in decreased corrosion levels and hydrogen pickup during 

normal reactor operation. Reactor coolant temperature is also an important factor in determining corrosion 

level and hydrogen pickup, as both decrease with decreasing coolant temperatures. For BWRs, the inlet 

and outlet coolant temperatures are about 278°C and 288°C, respectively. For PWRs, the inlet and outlet 

coolant temperatures are about 296°C and 328°C, respectively. Thus, corrosion and hydrogen pickup tend 

to be greater in PWR cladding than in BWR cladding. Additional details are provided in Chapter 2 for the 

as-fabricated and high-exposure (sometimes referred to as high-burnup) cladding materials used in the 

test program. 
 

2

Table 1. Dimensions and compositions of PWR cladding alloys used in test program; OD is outer 

diameter and h is wall thickness.

Parameter

15×15 Zry-4

Vintage 

Cladding

15×15 Zry-4

Low-Sn

Cladding

17×17 Zry-4

Low-Sn

Cladding

17×17 ZIRLO

Standard

Cladding

17×17 M5

Standard

Cladding

OD, mm 10.76 10.91 9.50 9.50 9.50

h, mm 0.77 0.67 0.57 0.57 0.57–0.61

Sn, wt. % 1.29–1.42 1.29 1.29 0.99 0.02

Nb, wt. % --- --- --- 0.98 1.02

O, wt. % 0.137 0.124 0.12 0.120 0.145

Fe, wt. % 0.21 ≈0.21 ≈0.21 0.11 0.05

Cr, wt. % 0.10 ≈0.10 ≈0.10 <0.01 <0.01

1.2 Transient Cladding Temperatures

In a LOCA transient, much of the coolant inventory could be lost, and the fuel cladding would begin to 

heat up due to stored heat and decay heat – coupled with the absence of sufficient cooling. Around 750oC 

to 850oC during the temperature rise, the pressurized cladding would reach its ultimate tensile strength, 

and a local plastic instability would produce a ballooning strain and a rupture. A modest temperature 

reduction would accompany this deformation in the vicinity of the balloon because of enhanced turbulent

cooling, but the cladding temperature would then continue to rise. Eventually, the emergency cooling 

water would stop the temperature rise and begin to cool the cladding. The cooling rate would be gradual 

at first because of steam-blanketing around the hot cladding, but then rapid cooling, or quench, would 

take place, cooling the cladding to about 135oC (saturation temperature at the expected reactor 

containment pressure). Such a temperature transient is shown schematically in Fig. 1. 

Figure 1. Schematic of cladding temperature history during a LOCA transient.
 

3

The calculated peak cladding temperature (PCT) for rods in a fuel assembly depends on fuel burnup, core 

location, and degree of conservatism used in the thermal analysis. Not all fuel rods will reach the cladding 

rupture temperature, nor will all rods that rupture reach a high PCT. The PCT and quench temperature can 

be lower for some LOCA transients, while the time at elevated temperature and pressure can be much 

longer. For those longer times at relatively lower temperatures, the normal oxidation process can become 

enhanced (so-called “breakaway oxidation”) and accompanied by rapid hydrogen absorption, which leads 

to cladding embrittlement. 

In the current work, no distinction is made between large- and small-break LOCAs because the 

experimental results and derived limits apply to both. PCT values in the range of 1000°C to 1204°C 

(licensing limit) were used to derive oxidation limits based on maintaining post-quench ductility. Tests 

were also conducted at 800°C to 1000°C PCT for times up to 5000 s to determine the minimum 

breakaway-oxidation time.

1.3 Transient Phase Transformations

During the heating phase of a LOCA transient, the zirconium-based cladding alloys of interest have a 

hexagonal crystal structure, the alpha phase, from room temperature up to about 650–800oC. Above this 

temperature, there is a mixed-phase regime in which the crystal structure begins to change to a facecentered cubic geometry, the beta phase. For standard Zry-4 with 0.11 wt. % oxygen, the alpha to the 

mixed alpha-beta phase-change temperature is 810oC, and the alpha-beta to beta phase-change 

temperature is 980oC [7]. These phase-change temperatures increase with increasing oxygen content [7] 

and temperature ramp rate [8], and they decrease with increasing hydrogen content [8-13]. They also 

decrease with the substitution of Nb for Sn because Nb stabilizes the beta phase while Sn stabilizes the 

alpha phase [12]. Alloying elements Fe and Cr also stabilize the beta phase [12]. Thus, oxygen and Sn are 

alpha-stabilizers, while hydrogen, Nb, Fe and Cr are beta-stabilizers. The alpha to alpha-beta phasechange temperature is particularly sensitive to concentrations of these alloying elements, to hydrogen 

picked up during normal operation, and to the heating rate during the LOCA transient. 

During the oxidation portion of a LOCA transient, alpha-phase Zr-based alloys have both high affinity 

for, and high solubility of, oxygen at high temperatures [13]. At LOCA-relevant temperatures, the 

oxygen-stabilized alpha phase is more thermodynamically stable than the oxide itself. Oxygen diffusion 

in the ZrO2 oxide layer and in the Zr metal becomes rapid enough above 800oC that measurable changes

in metal oxygen concentration occur during the time period of a LOCA transient. Therefore, after some of 

the Zr alloy has transformed from the oxygen-rich alpha phase to the beta phase, the temperature is high 

enough that oxygen diffuses into the beta-phase metal, exceeding its solubility limit at the interface with 

oxygen-stabilized alpha layer. This oxygen-rich metal transforms back to the alpha phase, where the 

oxygen solubility limit is much higher than in the beta phase. Figure 2 is a qualitative representation of 

the oxygen content within the oxide, alpha phase, and beta phase for cladding heated to high temperature 

in flowing steam (outer surface only). Within a ruptured balloon, steam will also enter through the rupture 

and produce similar oxygen penetration from the inside surface of the cladding. Oxygen penetration can 

also occur from oxide and fuel bonded to the inside surface of high-exposure cladding. Upon cooling, the 

beta phase also converts back to the alpha phase, but this prior-beta region retains its low oxygen 

concentration and is ductile at 135oC for average oxygen concentrations <0.6 wt. %. The prior oxygenstabilized alpha region retains its relatively high oxygen concentration and is brittle, as is the oxide layer. 
 

4

Figure 2. Qualitative diagram of oxygen concentration in Zircaloy cladding exposed at high 

temperature to steam on the outside surface and cooled to room temperature. OD 

and ID are outer and inner diameters, respectively.

1.4 Oxidation Equations

Although Baker-Just equations [14] have been used in many LOCA analyses, Cathcart-Pawel (CP) 

equations are more accurate for Zircaloys and will be used throughout this report. In their report [15], 

Cathcart et al. gave four related rate equations, one for each of the following: (a) the oxide layer 

thickness, (b) the alpha layer thickness, (c) the sum of the oxide-plus-alpha layer thickness, and (d) the 

total oxygen consumed, which was given as weight gain in grams per square centimeter of surface area. 

All these obeyed parabolic kinetics. The rate equation for weight gain, wg (Cathcart et al. called it τ), is:

d(wg)/dt = (k2

/2)/ wg, (1) 

where k is a temperature-dependent coefficient (Cathcart et al. labeled it δτ). Under isothermal oxidation 

conditions, the integration of Eq. 1 is simply

(wg)

2

/2 = (k2/2) t, (2) 

or 

wg = k t1/2, (2a) 

which plots as a parabola. 

The coefficient, k, is an Arrhenius-type function of temperature: 

k = a exp(-Q/[R×T]), (3) 

where R is the universal gas constant, and the parameters "a" and "Q" are given by Cathcart et al. for each 

of the four rate equations reported. Using the Cathcart et al. values for "a" and "Q" for weight gain 

determined from metallographic results and the assumption of stoichiometric oxide, the isothermal CP 

equation becomes:

wg = 0.602 exp(-1.005×104

/T) t1/2, (4) 

26

6.7

0.7

OD ID

Oxide

 Oxygen

Stabilized

 Alpha Prior Beta

Cladding Thickness

Oxygen Concentration (wt.%, Not to Scale)
 

5

where wg is in g/cm2, T is temperature in K, and t is time in s. Thus, the weight gain increases as the 

square root of time at a given temperature, and the rate increases exponentially with temperature. The 

results of Eq. 4 should be multiplied by a factor of 10 to convert to SI units of kg/m2

. For convenience in 

the current work, the measured weight gain is expressed as mg/cm2

. In comparing the measured weight 

gain per unit surface area to the CP-predicted weight gain, results of Eq. 4 are multiplied by 103 to give 

units of mg/cm2

. 

For steam oxidation at 1100oC and 1200oC, the agreement between calculated and measured weight gains 

was found to be excellent for the cladding materials (Zry-2, Zry-4, ZIRLO, M5, and E110) tested earlier 

in this program [1]. Differences among the alloys were observed at 1000oC prior to breakaway oxidation: 

(a) measured weight gains for Zry-2 and Zry-4 remained in excellent agreement with calculated weight 

gains, (b) measured weight gains for Nb-bearing alloys (ZIRLO, M5, and E110) were considerably less 

than calculated weight gains due to slowing down of oxide-layer growth with increasing time, and (c) 

measured weight gains (based on oxide layer thickness) were considerably less for the Zr liner within 

Zry-2 [1]. However, as will be discussed later in this section, the post-quench ductility decrease and 

embrittlement with increasing oxidation time were similar for all alloys tested (even at 1000oC) and 

correlated quite well with calculated weight gain normalized to cladding wall thickness.

A related parameter that is often used is ECR, or equivalent cladding reacted. ECR is defined as the 

percentage of the cladding thickness that would be oxidized if all the oxygen pickup stayed in the oxide 

layer as ZrO2. This is an artificial parameter because some of the oxygen diffuses into the metal, but it is 

useful and is directly related to weight gain by simple geometric factors, along with the density of Zr 

(6500 kg/m3

 = 6.5 g/cm3), the atomic mass of Zr (91.2 kg/kg-mole), and the atomic mass of diatomic 

oxygen (32.0 kg/kg-mole). The conversion is given in Eq. 5 for one-sided oxidation and Eq. 6 for twosided oxidation. 

One-sided oxidation ECR = 43.9 [(wg /h)/(1 ˗ h/Do)], (5) 

Two-sided oxidation ECR = 87.8 wg /h, (6) 

where ECR is in %, wg is in g/cm2

, h is cladding thickness in cm, and Do is cladding OD in cm.

1.5 Embrittlement Mechanisms

Six embrittlement mechanisms are described below. Data relating to these mechanisms are presented in 

subsequent sections of this report.

1.5.1 Beta-layer embrittlement by oxygen 

At 1200oC the oxygen solubility limit of beta-phase Zircaloy is about 0.6 wt% (0.57 wt. % [7]), and the 

solubility increases with increasing temperature. Above 1200oC, oxygen solubility and diffusion rates in 

the beta phase become high enough in Zr-based cladding alloys that, when cooled, the prior-beta region is 

brittle after relatively short oxidation times. This mechanism was understood in 1973, and the 1204oC 

(2200oF) temperature limit in NRC’s regulation precluded such rapid embrittlement in non-irradiated 

Zircaloy [6].

Oxygen solubility in the beta-layer is lower at 1100oC (0.38 wt. %) and 1000oC (0.24 wt. %) [7], and the 

prior-beta layer remains ductile at temperatures as low as room temperature (RT) even after the layer is 

saturated with oxygen. These solubility limits apply to as-fabricated cladding.
 

6

On the basis of the previous discussion, it would be more logical to correlate post-quench ductility with 

the average oxygen content in the prior-beta layer rather than with an ECR value. However, it is easier to 

calculate weight gain and CP-ECR, and there is a strong correlation between oxygen content in the priorbeta layer and CP-ECR. The layer-thickness correlations, oxygen concentrations at the metal-phase 

interfaces, and diffusion equations in Ref. 15 were used to calculate wg, CP-ECR, and average oxygen 

content (co) in the beta layer as a function of oxidation time at 1200oC for as-fabricated 17×17 Zry-4 

cladding with an initial oxygen content (coi) of 0.12 wt. % and a solubility limit (So) of 0.57 wt. %. Figure 

3 shows the near-linear relationship between the normalized oxygen pickup in the beta layer and CP-ECR 

in the range of interest (i.e., up to 18%, beyond which the beta layer is essentially saturated with oxygen). 

Therefore, using the CP-ECR value to characterize post-quench ductility of as-fabricated cladding is just 

as effective as using the prior-beta oxygen content. 

Figure 3. Average oxygen content (co) in the beta layer as a function of CP-ECR for 

isothermal oxidation at 1200oC. The calculation was performed for 17×17 Zry-4 

with 0.12 wt. % initial oxygen content (coi) and 0.57 wt. % solubility limit (So).

Notwithstanding this good correlation with calculated CP-ECR values, measured ECR values for Nbbearing alloys (e.g., M5) did not correlate very well with ductility decrease for oxidation at 1000oC [1]. 

The slow growth rate of the oxide layer resulted in low values of measured ECR for M5 with increasing 

test time. At oxidation times corresponding to about 20% CP-ECR, the measured ECR for M5 was only 

13% (the measured ECR for Zry-4 [22%] was close to the CP-ECR value). However, the RT post-quench 

ductility value (3.2%) was essentially the same for M5 and Zry-4 [1]. The decreased weight-gain rate for 

M5 was predominantly due to the decreased rate of oxide-layer growth. This reduction would not affect 

the diffusion of oxygen into the beta-phase metal and hence the oxygen concentration in the beta phase. 

The average oxygen concentration in the beta phase increases at a diffusion-dependent rate, which 

depends on time at temperature. These are reflected in the CP-ECR correlation. 

Calculated CP-ECR values and measured ECR are similarly independent of pre-transient hydrogen 

content, which increases the oxygen content in the beta layer and thus decreases post-quench ductility for 

a given oxidation temperature and time. Therefore, hydrogen content must be considered as an 

independent parameter in correlating ductility decrease with increasing CP-ECR. 

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0 5 10 15 20

CP-ECR (%)

(Co-Coi)/(So-Coi)
 

7

1.5.2 Beta-layer thinning

With increasing time, diffusion of oxygen into the metal will convert more and more of the beta phase to 

the oxygen-stabilized alpha phase – the alpha layer grows and the beta region shrinks. For long times at 

temperature, the ductile prior-beta region becomes so thin that the macroscopic specimen exhibits brittle 

behavior. This mechanism was also understood in 1973 and was accommodated by the 17% ECR limit in 

NRC’s regulation [6]. 

However, on the basis of previous test results [1], beta-layer thinning did not contribute to the 

embrittlement of as-fabricated cladding alloys within the range of oxidation levels (5% to 20%) included 

in the test program. For 17×17 Zry-4 (0.57-mm h) oxidized at 1100oC, the RT post-quench ductility 

leveled off to ≈5% strain between 15% and 20% CP-ECR even though the beta-layer thickness continued 

to decrease to a minimum of 0.33 mm at 20% CP-ECR.

Nevertheless, beta-layer thinning does contribute to hydrogen-enhanced beta-layer embrittlement by 

oxygen. This subject is discussed in Section 1.5.4. 

1.5.3 Localized hydrogen-induced embrittlement in the balloon region

Steam that enters through a rupture in a balloon causes oxidation inside the cladding. Hydrogen that is 

released during this reaction is not swept away as it is on the outside of the cladding, but is absorbed in 

the metal above and below the rupture opening. Beta-phase Zr-alloys have very high hydrogen solubility 

(e.g., ≈104

 wppm at 1200oC), and hydrogen absorption results in enhanced embrittlement in the balloon 

region from the rupture edges to the balloon necks. This effect, which is neither burnup nor alloy 

dependent, was discovered earlier (see Ref. 6) and has been confirmed in the present research program. 

For ballooned, ruptured, and oxidized Zry-2 cladding (as-fabricated and high-burnup), hydrogen peaks as 

high as 3000-4000 wppm were measured in the oxidized cladding [1]. In this previous work, only two 

post-quench four-point bend tests (4-PBTs) were conducted at RT with ballooned and ruptured asfabricated Zry-2 samples oxidized to 9% and 15% CP-ECR, respectively. Neither test sample failed in the 

high-hydrogen region because the balloons were large and the weakest cross section was in the rupture 

region. However, in the current work, a large number of such LOCA oxidation and post-quench 4-PBTs 

were performed with as-fabricated and pre-hydrided ZIRLO cladding, and some did fail in the highhydrogen region. These test results allowed us to better characterize degradation of mechanical properties 

as a function of pre-transient hydrogen content and oxidation level. 

1.5.4 Hydrogen-enhanced beta-layer embrittlement

During normal operation, some hydrogen from the corrosion process is absorbed in the cladding metal. 

When that cladding is exposed to high-temperature LOCA conditions, the elevated hydrogen levels 

increase the solubility of oxygen in the beta phase and the rate of diffusion of oxygen into the beta phase. 

Thus, even for LOCA temperatures below 1204oC, embrittlement can occur at times corresponding to less 

than 17% oxidation in corroded cladding with hydrogen pickup. Significant degradation can occur at 

hydrogen concentrations as low as 100 wppm. For high-burnup M5 with 140-wppm hydrogen, previous 

results showed that the ductile-to-brittle transition oxidation level for oxidation at 1200°C decreased from 

20% (for as-fabricated M5) to about 14%. For high-burnup ZIRLO with 600-wppm hydrogen, the 

transition oxidation level decreased from 19% (for as-fabricated ZIRLO oxidized at 1200oC) to 4% with a 

peak oxidation temperature of 1130oC. In the current work, pre-hydrided and high-burnup ZIRLO 
 

8

samples with intermediate pre-transient hydrogen contents of 200 to 340 wppm were used to determine 

oxidation embrittlement thresholds. These levels refer to hydrogen contents in the cladding metal. 

During the oxidation portion of a LOCA transient, pre-transient hydrogen concentrates in the beta-phase 

metal, as this phase has a much higher hydrogen solubility and affinity for hydrogen than the oxygenstabilized alpha phase. As the beta layer decreases in thickness, the hydrogen content in this layer 

increases correspondingly, along with the oxygen solubility [13]. Beta-layer thinning thus contributes to 

hydrogen-enhanced post-quench embrittlement of cladding at lower oxidation levels. 

In addition to elevating oxygen solubility, there appears to be an intrinsic embrittlement effect of 

hydrogen dissolved in the beta phase [9-11]. During very rapid cooling from the oxidation temperature, 

hydrogen is essentially frozen in the low-oxygen prior-beta phase. For a fixed oxidation time (e.g., 50 s at 

1200oC and one-sided oxidation), Zry-4 samples with 600-wppm hydrogen were very brittle at CP-ECR 

values of ≈3%. For much slower cooling rates and lower quench temperatures (e.g., 600oC and 700oC), 

significant ductility was retained in these samples even though the oxidation level increased to ≈6% CPECR. The enhancement in ductility is attributed to small-scale diffusion of hydrogen out of the loweroxygen regions of the prior beta phase to the interfaces between these lower-oxygen regions and higheroxygen regions formed during pre-quench cooling. Although the cooling rates at which ductility was 

enhanced [9-11] were too slow to be LOCA-relevant, the results do support the hypothesis that there is an 

intrinsic embrittlement effect of hydrogen beyond the dominant effect of increasing the oxygen solubility. 

1.5.5 General hydrogen-induced embrittlement from breakaway oxidation 

Zirconium dioxide (ZrO2) can exist in several crystallographic forms (allotropes) and has a monoclinic 

structure at normal operating temperatures. The tetragonal oxide that generally develops under LOCA 

conditions is dense, adherent, and protective with respect to hydrogen pickup. Although the monoclinicto-tetragonal phase transformation temperature is nominally ≈1150°C, the tetragonal oxide phase can be 

stabilized at lower oxidation temperatures by compressive stress, hypostoichiometry, and to some extent 

grain size. There are, however, mechanical (e.g., local regions of tensile stress) and chemical (e.g., 

impurities at metal surface) conditions that promote a transformation to the lower-temperature monoclinic 

phase that is neither fully dense nor protective. When these conditions are present, the tetragonal-tomonoclinic transformation can begin in an unstable way at local regions of the metal-oxide interface and 

grow rapidly throughout the oxide layer due in part to the local volume increase (≈5%) associated with 

the phase transformation. Because this transformation results in an increase in oxidation rate, it is referred 

to as breakaway oxidation. Along with this increase in oxidation rate due to cracks in the monoclinic 

oxide, there is significant hydrogen pickup. Hydrogen that enters in this manner during a LOCA transient 

has a similar effect on embrittlement as hydrogen from the normal burnup process. Although breakaway 

oxidation was known in 1973, a connection to embrittlement was not made at that time.

In the previous work, a large test matrix was used to determine minimum breakaway times and 

corresponding isothermal oxidation temperatures for two variants of 15×15 Zry-4, for 10×10 Zry-2, for 

17×17 ZIRLO, and for an older lot of Russian E110 (Zr-1Nb alloy). In the current work, isothermal tests 

were conducted with M5 cladding (up to 5000 s), the isothermal database for ZIRLO was expanded to 

include numerous tests within the narrow temperature range of 970oC to 1000oC, and tests were 

conducted with ZIRLO subjected to transient oxidation temperature histories.
 

9

1.5.6 Oxygen pickup from the cladding inner surface

For non-irradiated cladding, the oxygen source for diffusion into the metal is the ZrO2 layer that forms 

during high temperature oxidation in steam. As burnup increases, normal corrosion will contribute to the 

outer-surface oxygen source. There will also be an oxygen source on the inner surface of irradiated 

cladding due to gas-phase UO3 transport prior to gap closure, fuel-cladding-bond formation (UO2 in solid 

solution with ZrO2), and the oxide fuel bonded to this layer. Thus in high-burnup fuel under LOCA 

conditions, oxygen can enter the cladding from the inner surface even away from the balloon location. 

The situation is like a three-layer diffusion “couple” as illustrated in Fig. 4. The actual thicknesses of the 

OD and ID oxygen sources may be relatively unimportant at high burnup because they contain much 

more oxygen than will diffuse into the metal (see related tests by Hofmann and Politis [16]). However, if 

the ID oxide layer is the only source of inner-surface oxygen, it may be completely reduced by the 

cladding metal prior to high-temperature embrittlement and the effect would be limited. Nevertheless, if 

the inner-surface oxide layer or the oxide-bond-fuel layer is well developed (see Fig. 5), then two-sided 

oxidation would have to be accounted for in LOCA behavior away from the balloon region. 

Figure 4. Three-layer diffusion “couple” character of oxygen sources and cladding metal. 

Figure 5. Well-developed fuel-cladding bond layer (gray) at the fuel mid-plane of a Zry-4-clad 

fuel rod irradiated to 67 GWd/t in the H.B. Robinson reactor.

Oxide Metal Oxide

Steam

Oxide 

and

Corrosion

Bonding

Layer 

and

Fuel

Cladding

OD ID
 

10

1.6 Conventional Use of Equivalent Cladding Reacted (ECR)

The embrittlement mechanisms identified in Section 1.5 depend on oxygen diffusion into the cladding 

metal. It is clear from the work of Cathcart & Pawel and others that these diffusion-related processes have 

a form similar to the weight-gain relation (Eq. 4) for isothermal oxidation. Although it is not clear how to 

characterize the amount of oxygen diffusion that results in certain mechanical behavior of interest, it is 

convenient to empirically correlate this mechanical behavior with CP-ECR – a measure of time at 

temperature. In other words, mechanical behavior is correlated with the time it takes to produce a certain 

amount of oxygen pickup at a given temperature, as normalized to cladding dimensions. This practice has 

been used for 40 years and we continue to find it useful. For consistency, we will always use the CP 

weight-gain correlation to calculate wg in Eqs. 5 and 6 and refer to that parameter as CP-ECR. As CPECR is independent of hydrogen content while ductility decrease is highly dependent on hydrogen 

content, the pre-transient hydrogen content will be treated empirically as an independent parameter such 

that the embrittlement oxidation level (CP-ECR) will be expressed as a function of this hydrogen content. 

1.7 Pre-transient Hydrogen Content in the Cladding Metal (CHM) 

The pre-transient hydrogen content in the cladding metal is easier to define than it is to measure for highexposure, corroded cladding. Experimentally, CHM is referred to as the pre-test (i.e., pre-LOCA-oxidation 

test) hydrogen content (in wppm) in the cladding metal. The methodology developed and applied in the 

current work (see Sections 3 and 5) used a combination of pre-test and post-test measurements to 

determine CHM. Post-test hydrogen-content measurements were corrected to account for weight gain 

during the oxidation phase of the testing, the mass of the pre-test corrosion layer, and the mass of the pretest fuel-cladding bond layer. As hydrogen in the corrosion and fuel-cladding-bond layers is driven off 

during the oxidation test, no additional correction factors were needed to account for this extra source of 

hydrogen. 

The criterion used for initiation of breakaway oxidation (i.e., breakaway oxidation time) is 200 wppm 

hydrogen pickup (ΔCHM). The parameter ΔCHM refers to the increase in hydrogen content during the 

oxidation test referenced to the pre-test metal mass of the cladding sample. As pre-test cladding samples 

were non-irradiated and non-oxidized (see Section 4), correction factors were not needed for the masses 

of the corrosion and fuel-cladding bond layers. Post-test measurements of hydrogen content were 

corrected for weight gain, and the initial hydrogen content in as-fabricated cladding was subtracted to 

determine ΔCHM. 
 

11

2 MATERIALS, TEST METHODS, AND PROCEDURES

Testing in this work includes LOCA integral tests (heat-up, ballooning, rupture, oxidation, and quench) 

and separate tests for oxidation and embrittlement. Interpretation and application of LOCA-relevant data 

require an understanding of cladding materials used for test samples, test methods, test procedures, and 

data-analysis methods. These were described in detail in Ref. 1 for work performed through June 2008. 

For the convenience of the reader, some of these details are repeated in this section. 

Also included in this section are refinements and improvements made during the course of the current 

work in sample preparation, data collection, and data analysis. During this time, draft procedures for 

conducting breakaway-oxidation tests [17] and post-quench ductility tests [18], along with data analysis 

methods, were documented and submitted to NRC in March 2009. These procedures were initially 

followed in the current work and later improved based on the practical experience of applying these 

procedures and data-analysis methods. Argonne also worked with NRC to improve, generalize, and 

formalize these documents into three regulatory guides needed for the planned performance-based LOCA 

acceptance criteria: RG 1.222 [2], RG 1.223 [3], and RG 1.224 [4]. These regulatory guides are readily 

accessible, and they contain additional improvements that were made since Refs. 17 and 18 were 

submitted to NRC.

2.1 Description of Cladding Alloys and High-Burnup Fuel Segments

As-fabricated cladding alloys were provided by Global Nuclear Fuels (8×8, 9×9, and 10×10 Zry-2), 

Westinghouse (17×17 low-Sn Zry-4 and ZIRLO), AREVA (1970s 15×15 Zry-4 tubing, 1980s 15×15 lowtin Zry-4, 15×15 low-tin Zry-4, 15×15 M5, and 17×17 M5), and Fortum (standard E110 tubing and 

cladding). Alloys, except for E110, listed above without a year are modern, belt- or wheel-polished (OD) 

and sand-blasted (ID) cladding alloys with reduced surface roughness and impurities. The 8×8 Zry-2 was 

used for early thermal and metallurgical benchmark testing of the LOCA integral apparatus. The 9×9 Zry2 was fabricated by the same methods used to fabricate high-burnup Limerick fuel-rod cladding and had 

the same nominal composition and dimensions as the Limerick cladding. It was used in the previous test 

program to generate baseline LOCA integral data and one-sided, high-temperature steam oxidation data. 

The 1970s 15×15 Zry-4 tubing was close to being archival material for the high-burnup H. B. Robinson 

(HBR) cladding tested in this program. However, this tubing had ≈0.05-mm larger outer diameter and 

≈0.025-mm thicker wall than the HBR cladding. The final finishing step after fueling the HBR tubing to 

convert it to cladding was pickling to remove ≈0.05 mm from the outer diameter. Because this tubing was 

not in final cladding form, it was not used to generate baseline data. Rather, the 1980s 15×15 low-Sn Zry4 cladding, which had similar dimensions, oxygen content, and mechanical properties as the HBR 

cladding, was used in previous work to determine baseline HBR oxidation and post-quench ductility 

properties. This cladding is referred to as "HBR-type Zry-4" to distinguish it from more modern Zry-4 

materials.

Some 15×15 and 17×17 M5 cladding was provided in 2002 by AREVA for benchmarking purposes but 

not for data generation. This 17×17 M5 (0.57-mm wall) was characterized in terms of chemical 

composition and used to develop and benchmark the two-sided oxidation test train. At a later date, 

AREVA provided 17×17 M5 with 0.61-mm wall thickness for data-generation purposes. The post-quench 

ductility of the 15×15 M5 and the 0.61-mm-wall 17×17 M5 was compared to confirm that post-quench 

ductility correlated with the calculated extent of oxidation (CP-ECR) at a given temperature independent 

of differences in cladding wall thickness and outer diameter. 
 

12

For the post-quench ductility program, the primary alloys of interest were 10×10 Zry-2, 17×17 low-Sn 

Zry-4, 17×17 ZIRLO, and 17×17 M5 (0.61-mm wall). E110 was also added to the test program to 

determine why its steam-oxidation behavior was so poor as compared to M5, even though the 

compositions of both alloys were very similar. E110 tubing and final-form cladding were provided by 

Fortum (Finland). Results of previous tests performed with E110 material are well documented [1,19]. As 

such, these test results are not included in this report. 

For breakaway-oxidation studies, the test plan called for the same alloys as used for the post-quenchductility studies. However, due to the limited supply of 17×17 low-Sn Zry-4, wheel-polished 15×15 and 

HBR-type (rough surface) 15×15 low-Sn Zry-4 were used instead. This allowed the effects of surface 

roughness to be included in these studies for Zry-4 samples with similar geometry, Sn content, and 

oxygen content. The other three alloys (10×10 Zry-2, 17×17 ZIRLO, and 17×17 M5) used in the postquench-ductility tests were also used in the breakaway oxidation tests. However, test results for the 

ZIRLO received in 2003 indicated early breakaway oxidation at the cladding inner surface for oxidation 

at 1000°C. Additional lots were received in 2006, 2008, and 2010. As none of these lots exhibited early 

breakaway oxidation on the cladding inner surface, it is hypothesized that there was a change in final 

treatment of the cladding inner surface to reduce surface roughness and impurities. Breakaway-oxidation 

test results are provided in Ref. 1 for ZIRLO–2006. ZIRLO–2008 was used in the current work for 

breakaway oxidation tests and post-quench ductility (PQD) tests with pre-hydrided cladding. LOCA 

integral tests were conducted in the current work with as-fabricated and pre-hydrided ZIRLO–2010. 

Measurements were performed by Argonne to determine cladding OD, wall thickness, surface roughness, 

oxygen content, and hydrogen content. Nominal Sn contents for the cladding alloys were provided to 

Argonne by the fuel vendors. Some of the alloys were sent to an outside organization to determine 

chemical composition. For ZIRLO–2003 cladding, Westinghouse also performed chemical analysis on a 

sample from the cladding sent to Argonne. Some Argonne wall-thickness measurements were merely spot 

checks of this dimension at several axial locations. More detailed dimensional characterization was 

performed for HBR archive, for HBR-type Zry-4, for 17×17 M5 (0.61-mm h), and for lots of ZIRLO 

cladding received after 2003. 

Table 2 summarizes characterization results for the Zr-lined Zry-2 used in the previous test program. The 

Zr liner was on the cladding inner surface (≈10% of the wall thickness). Chemical composition data listed 

in Table 2 refer to the final cladding product (≈90% Zry-2 and ≈10% Zr liner). They should not be 

interpreted as the composition of standard Zry-2. 

Dimensions and chemistry of the Zry-4 used in the previous test program are listed in Table 3. Pickling 

was used for the inner-surface treatment of near-archive HBR tubing and HBR-type cladding. Based on 

24 measurements for adjacent 25-mm long samples, wall thicknesses were determined to be 0.790±0.006 

mm and 0.761±0.006 mm, respectively, for tubing and cladding. 

Dimensions and chemistry for 17×17 ZIRLO (2003), 17×17 M5 (validation lot), 17×17 M5 (data lot), and 

15×15 M5 (validation lot) used in the previous test program are listed in Table 4. Cladding OD and wall 

thickness values listed in Table 4 are nominal values provided by the vendor and spot checked by 

Argonne. Additional measurements were made for a 1-m-long segment of 17×17 M5 cladding used for 

data generation in both previous and current test campaigns. The OD was 9.48±0.00 mm (based on 24 

measurements at 12 axial locations) and h was 0.062±0.01 mm (based on 12 measurements). 
 

13

Table 2. Dimensions and chemistry of Zr-lined Zry-2 used in the Argonne test program (the “<” 

sign means below the detection limit and “---” means not measured).

Parameter 8×8 Zry-2 9×9 Zry-2 10×10 Zry-2

OD, mm 12.27 11.18 10.29

Wall Thickness, mm 0.82 0.71 0.66

ID Liner Thickness, mm ≈0.08 ≈0.07 ≈0.07

OD Surface Roughness, μm --- 0.14 0.11

Sn, wt.% --- 1.18 ---

Nb, wt.% --- <0.01 ---

O, wt.% --- 0.11 ---

Fe, wt.% --- 0.20 ---

Cr, wt.% --- 0.12 ---

Ni, wppm --- 550 ---

S, wppm --- 30 ---

C, wppm --- 250 ---

Hf, wppm --- <100 ---

Si, wppm --- ≤100 ---

N, wppm --- 45 ---

H, wppm --- 6 13

Table 3. Dimensions and chemistry of Zry-4 used in the Argonne test program (the "<" sign 

means below the detection limit and “---” means not measured). 

Parameter

15×15 Zry-4 

HBR Tubing

15×15 Zry-4 

HBR Cladding

15×15 Zry-4 

Cladding

17×17 Zry-4 

Cladding

OD, mm 10.84 10.77 10.91 9.50

Wall Thickness, mm 0.80 0.76 0.67 0.57

OD Surface Roughness, μm 0.36 0.32 0.10 0.14

Sn, wt.% 1.42 1.29 1.29 1.29

Nb, wt.% --- --- --- ---

O, wt.% 0.137 0.136 0.124 0.120

Fe, wt.% 0.21 --- --- ---

Cr, wt.% 0.10 --- --- ---

Ni, wppm <35 --- --- ---

S, wppm --- --- --- ---

C, wppm 140 --- --- ---

Hf, wppm <50 --- --- ---

Si, wppm 92 --- --- ---

N, wppm 54 --- --- ---

H, wppm 12 22 26 5
 

14

Table 4. Dimensions and chemistry of ZIRLO and M5 cladding materials used in the Argonne 

test program (the "<" sign means below the detection limit “---” means not measured). 

Parameter

17×17 

ZIRLO–2003 

17×17 M5

(Validation) 

17×17 M5

(Data)

15×15 M5

(Validation) 

OD, mm 9.50 9.50 9.50 10.91 

Wall Thickness, mm 0.57 0.57 0.61 0.64 

OD Surface Roughness, μm 0.11 0.12 0.12 0.11 

Sn, wt.% 0.99 0.02 --- ---

Nb, wt.% 0.98 1.02 --- ---

O, wt.% 0.12 0.145 0.145 ---

Fe, wt.% 0.11 0.05 --- ---

Cr, wt.% <0.01 <0.01 --- ---

Ni, wppm <100 <100 --- ---

S, wppm --- 25±5 --- ---

C, wppm 135 110 --- ---

Hf, wppm 40 <100 --- ---

Si, wppm 53 <100 --- ---

N, wppm 46 55 --- ---

H, wppm 5 6 5 ---

Additional lots of ZIRLO were used in the previous and current work, and a comparison of the four lots 

of ZIRLO that have now been tested is shown in Table 5. Detailed chemistry values were not provided for 

the later quantities of ZIRLO, but more detailed dimensional characterization was performed by Argonne. 

Where one-sigma uncertainty ranges are shown in Table 5, multiple measurements were made (e.g., 90 

data points over 11 m for the OD of ZIRLO–2008). Dimensions for each PQD and breakaway-oxidation 

sample were recorded and used to determine weight gain per unit surface area and CP-ECR. 

Table 5. Characterization of the four lots of ZIRLO used in the Argonne test program. 

Parameter

17×17 

ZIRLO-2003 

17×17

ZIRLO-2006 

17×17

ZIRLO-2008 

17×17

ZIRLO-2010 

OD, mm 9.50 9.49 9.48±0.03 9.50±0.01 

Wall Thickness, mm 0.57 0.57 0.59±0.01 0.58±0.01 

OD Surface Roughness, μm 0.11±0.01 0.17±0.03 0.14±0.03 0.16±0.01 

O, wt.% 0.119±0.003 0.116±0.003 ≈0.12 ≈0.12

H, wppm 5±0 11±3 17±2 22±2 
 

15

Irradiated materials used in the Argonne test program are summarized in Table 6. The Electric Power 

Research Institute (EPRI) provided two fueled TMI-1 PWR rods at 48-50 GWd/t for validation of 

characterization and testing methodologies, seven fueled Limerick BWR rods at 54-57 GWd/t for LOCA 

data generation, and seven fueled H. B. Robinson PWR rods at 64-67 GWd/t. EPRI and AREVA 

provided two high-burnup fueled M5 rods at 63 and 70 GWd/t for LOCA integral testing of fueledcladding samples. In addition, under agreements among NRC-Studsvik-AREVA and NRC-StudsvikWestinghouse, high-burnup defueled M5 cladding from the European Ringhals PWR reactor and ZIRLO 

cladding from the North Anna (NA) PWR reactor, respectively, were supplied for post-quench ductility 

testing.

Table 6. Characteristics of high-burnup fuel rod segments and defueled high-burnup cladding for 

the Argonne LOCA test program. Burnup values are rod averaged.

Parameter TMI-1 Limerick H. B.

Robinson

North Anna

M5a North Anna

ZIRLO

Ringhals

M5

Reactor PWR BWR PWR PWR PWR PWR

Enrichment,

wt.%

4.00 3.95

(3.40-3.95) 

2.90 4.20 ≈4 ≈4

Burnup, 

GWd/t 

48-50 54-57 63-67 63-70 70 63 

Discharge

Date

1997 1998 1995 2004 2001 2003 

Fast 

Fluence, 

1025 n/m2

9 11 14 Not

Provided 

Not

Provided 

Not

Provided 

Cladding 15×15 

Low-Sn

Zry-4 

9×9 

Zr-lined

Zry-2 

15×15 

Standard 

Zry-4 

17×17 

M5

17×17 

ZIRLO

17×17 

M5

Initial Wall

Thickness,

mm

0.69 0.71 0.76 0.57 0.57 0.57 

OD

Oxide, μm

≤30 ≈10 +

≈10 crud

≤100 <10 ≈20-50 <15 

Hydrogen

Pickup, 

wppm 

≤300 70 ≤800 <100 ≤650 ≤140

Fueled Yes Yes Yes Yes No No a

Not tested because of closure of Alpha Gamma Hot Cell Facility to programmatic work.

For both the previous and current work, oxidation and PQD tests were conducted on defueled samples of 

ZIRLO cladding from NA Rod AM2-L17: (a) 40−50 μm corrosion layer thickness and 450−600 wppm 

hydrogen in cladding metal for previous work and (b) 25−30 μm corrosion layer thickness and 300−340 

wppm hydrogen in cladding metal for current work. 
16 

2.2 General Description of Apparatus and Test Samples

The LOCA high-temperature apparatus was designed to perform both LOCA integral tests using long 

(≈300 mm) pressurized fueled cladding samples and oxidation-quench tests using short (25 mm) defueled 

cladding samples. An out-of-cell unit was used for thermal-benchmarking purposes and for generating 

data for as-fabricated and pre-hydrided alloys. Two in-cell units have been used in the Alpha-Gamma 

Hot-Cell Facility (AGHCF) and the Irradiated Materials Laboratory (IML) for testing high-burnup fueled 

and defueled cladding, respectively. When in operation, the units shared the same control and data 

acquisition systems used for the out-of-cell apparatus. The main difference in the apparatus between 

LOCA-integral and oxidation-quench tests is the test train holding the sample in position within the quadelliptical, radiant-heating furnace and the connections required. Both use a quartz-tube steam chamber 

(48-mm ID and 610-686 mm long). 

The radiant-heating furnace has four vertical bulbs, a reflecting inner surface, and a 250-mm-high 

uniform heating zone. LOCA integral tests were conducted with 300-mm-long samples initially under 

high internal pressure (≈8 MPa), which required connection to the high-pressure argon line, along with 

top and bottom pressure transducers. In the current work, the LOCA integral sample length was extended 

to 318 mm to minimize secondary ballooning in the top plenum region. Oxidation-quench-embrittlement 

tests with non-deformed cladding did not require the high-pressure connection or pressure transducers. 

For both types of testing, steam, at near-atmospheric pressure, flowed up through the quartz-tube test 

chamber at 5.3 g/minute and exited the chamber into a condenser. The quartz tube length was initially 686 

mm and later reduced to 610 mm for tests conducted in the IML. Following oxidation and slow-cooling 

phases, steam flow was turned off, and quench water was introduced (bottom flooding) to give very rapid 

sample cooling from 800oC to about 100oC at the selected test time. 

During the course of the current work, two changes were made to pre-test sample characterization and 

cleaning: (a) the measured cladding thickness was used for as-fabricated and pre-hydrided cladding 

samples, rather than the nominal cladding thickness, in the CP-ECR calculation and (b) etching (i.e., 

pickling) with an HF-containing acid was used prior to gaseous charging of samples with hydrogen (i.e., 

pre-hydriding). Following pre-hydriding, the etched surfaces were polished (removal of ≈1 μm from 

etched surfaces) to remove surface impurities, especially fluorine. Argonne experience with pickling 

indicated that even a very light etch (e.g., 15 to 60 s in an acid bath containing 1% HF) can induce early 

breakaway oxidation and hydrogen pickup at 1000°C to 1200°C oxidation temperatures (e.g., 100 s ramp 

and 180 s hold time at 1200°C resulted in ≈500-wppm hydrogen pickup). Higher HF concentrations (2 to 

4%) and longer etching times (120 to 240 s) did not result in early breakaway oxidation. As pickling has 

been used in some earlier experimental work [20,21] for sample cleaning, it is important to verify, in 

advance of data-generating tests, that this surface cleaning technique does not contribute to early 

breakaway oxidation due to the disruptive effects of fluorine on surface oxidation. In the Argonne 

program, the final cleaning methods consisted of sample submersion in an organic solvent (ethanol) 

within an ultrasonic water bath followed by a distilled-water rinse.

2.3 Oxidation and Quench Tests with Short Cladding Samples

Figure 6 shows the test train used in past and recent work for oxidizing and quenching samples prior to 

conducting post-quench ductility tests. Holes were drilled into the hollow Inconel holder below and above 

the sample to allow adequate steam flow inside the cladding sample. During the course of test-train 

development, the Inconel holder wall thickness and outer-diameter were reduced to decrease the thermal 

mass and the corresponding transient temperature lag of the holder. 
17 

Figure 6. Test train for two-sided oxidation tests with holes drilled into the Inconel holder 

below (left) the sample for steam ingress and holes drilled into the holder above

(right) the sample for steam egress. Also shown are the three thermocouples

permanently welded to the Inconel holder just above the sample (silver in 

appearance). 

Temperature control and monitoring are extremely important in conducting oxidation-quench tests, 

followed by post-quench ductility tests. As discussed in Section 1, the time-at-temperature for the 

transition between ductile and brittle behavior is a strong function of temperature. The furnace power and 

sample temperature are controlled by feedback from the designated control thermocouple (TC) output 

through a proportional-integral-differential controller to the furnace power. Because the sample has such 

low thermal mass per unit length, it is important to avoid temperature overshoot (due to the initially rapid 

heat generation from cladding oxidation) during the rapid temperature ramp to the hold temperature. In 

setting controller parameters, the requirements were that the temperature overshoot during the ramp be 

<20oC relative to the target hold temperature for a short period of time (less than a few seconds), and that 

the average hold temperature be within 10oC of the target temperature. The 10oC deviation is reasonable 

given the uncertainty of a few degrees in the Type S thermocouple readings at the maximum oxidation 

temperature (1204oC). Temperature overshoot is not much of an issue for long-time oxidation 

temperatures ≤1100oC, but it can have a significant embrittlement effect for higher oxidation 

temperatures. For tests conducted at 1200oC, temperature overshoot was minimized to <10oC by slowing 

down the furnace heating rate (to 2–3oC/s) at ramp temperatures within about 100oC to 200oC of the target 

temperature.

The following benchmarking procedure was established for each test train: 

a) Calibrate purchased TCs to ANL-owned, NIST-calibrated TC wire (introduced in January 2006). 

b) Weld two calibrated Type S TCs onto the sample outer surface; align one with the same orientation 

as the holder’s control TC and align the other at ±120o from the control TC. 

c) Use repeated testing with fresh and oxidized samples and same orientation for TCs welded to 

sample to determine optimum controller parameters for desired temperature ramp and hold 

temperature for test sample oxidized in steam from deionized water with ≈4 ppm dissolved oxygen. 

d) Repeat test with new sample and TCs welded to sample. (See Refs. 1–3 for additional steps for 

larger-diameter cladding [e.g., BWR 9×9 and PWR 15×15].) 

e) Perform metallographic examination to determine oxide-layer thicknesses at eight circumferential 

locations; compare measured and predicted average values to Cathcart-Pawel (CP) predicted values 

to ensure <10% difference; compare circumferential variation to ANL criterion (≤±3 μm).

Exit Hole

Inlet Hole
18 

f) Conduct oxidation test without TCs welded on the sample and compare measured sample weight 

gain at a time corresponding to about 10–13% CP-ECR to weight gain calculated with CP weightgain correlation; also perform metallographic examination to determine average inner- and outersurface oxide layer thicknesses, as well as variation. (Note: Zry-4 should be used for ≤1050°C 

oxidation temperatures.) 

g) Begin data-generating oxidation tests if the following are satisfied: overshoot is <20°C for <3 s 

during temperature ramp, average hold temperature is within 10°C of target temperature, 

circumferential temperature variation is ≤20°C, measured weight gain is within 10% of predicted 

weight gain, average oxide layer thickness on inner and outer surfaces is within 10% of predicted 

thickness, and circumferential variation in oxide layer thickness is consistent with circumferential 

temperature variation.

h) Continue to run tests with benchmarked test train and controller parameters for specific hold 

temperature until measured weight gain and CP–predicted weight gain differ by ≥10%; repeated 

quench will cause test-train warping, which moves the sample away from the furnace focal point; 

generally, a test train should be good for ≈15 runs with quench and more without quench.

Figure 7 shows the thermal benchmark result for the as-fabricated 17×17 ZIRLO cladding used in the 

current program. The cladding sample temperature (TC4 and TC5) rise is very rapid to ≈1100°C. To 

prevent overshoot at 1200°C, controller parameters were set to slow down the heating rate. After 100 s 

from initiation of heating from 300°C, the average cladding temperature was 1200oC. After 150 s, the 

cladding had essentially reached the 1202±7°C hold temperature. These temperatures and temperature 

variations were acceptable for proceeding to the weight gain benchmark test. The test was repeated with a 

fresh sample that did not have TCs welded to it. 

As shown in Figure 7, the oxidation test is not isothermal. The significance of the weight gain during the 

initial heating ramp depends on the hold temperature and the hold time. For the temperature history

shown in Figure 7, the CP-predicted weight gain was determined numerically by integrating Eq. 1 over 

the time-dependent temperature profile T(t), which was recorded every 0.1 s (i.e., 10 Hz): 

(wg)

2

 = 0.362 ∫

0

t

exp (-20100/T) dt, (7) 

where wg is in g/cm2

, T is in K, and t is in s. The integral in Eq. 7 was converted to an integral with 

respect to temperature, and the numerical integration for high-temperature oxidation tests (e.g., Fig. 7) 

was performed for T≥800oC, where the predicted weight gain between 800oC and 1000oC was very small 

during the temperature ramp and negligible during the cooling ramp. Because of the parabolic nature of 

the weight gain correlation, the weight gain accumulated during the cooling ramp is usually insignificant. 

However, the calculation does include cooling to 800°C prior to quench. For the test conducted without 

TCs welded to the sample, the measured weight gain (7.91 mg/cm2

) was within 5% of the CP-calculated 

weight gain (8.27 mg/cm2

 with corresponding 12.7% CP-ECR), which was acceptable for conducting 

data-generating tests with this test train and set of control parameters. Also, the measured OD and ID 

oxide layers (40 μm) were in good agreement with CP-predicted values. 
19

Figure 7. Thermal benchmark results for bare ZIRLO cladding. At 100-s into the heating phase, the average s

was 1200°C. After 150 s, the cladding temperature was 1202±7°C. Quench at 800°C is not shown in t

900

800

700

600

500

400

300

1300

1200

1100

1000

500 550 600 650 700 750 

C) ° Temperature (

Time (s)

Control TC1 on holder

TC2 on holder

TC3 on holder

TC4 on sample

TC5 on Sample
20

For oxidation and quench of as-fabricated and pre-hydrided samples used for post-quench ductility tests, 

the procedure is the same as the one used in the thermal and metallurgical benchmark tests:

a) Stabilize the system at 300oC in flowing saturated steam for 500 s. 

b) Initiate temperature ramp with feedback between the holder TC and the furnace power.

c) Hold at temperature for a predetermined time to achieve the desired CP weight gain and ECR. 

d) Turn furnace power off while maintaining steam flow at the end of the hold time. 

e) Stop steam flow and initiate bottom-flooding quench at a test time that will give rapid cooling at the 

desired cladding temperature. (Note: minimum temperature after quench was nominally 100oC; for 

tests conducted after January 2006, minimum temperature was reduced to 80±10oC.) 

The procedure for conducting in-cell oxidation-quench tests with high-burnup cladding samples is similar 

to the one described above for out-of-cell tests with non-irradiated cladding samples. However, the 

corroded, irradiated samples were not benchmarked directly because Argonne does not have the in-cell 

capability to locally remove the corrosion layer and spot-weld TCs to the cladding metal surface. With 

respect to the temperature history shown in Figure 7 for bare, non-irradiated cladding, it is expected that 

pretest cladding corrosion would slow down the initial oxidation rate and the heating rate associated with 

the exothermic oxidation reaction. The presence of the corrosion layer would basically affect the peak 

temperature reached during the very rapid heating ramp. This effect has been confirmed in out-of-cell 

tests using the following procedure to determine the temperature profile for high-burnup ZIRLO cladding 

with a 30-μm-thick corrosion layer and a 10-μm-thick fuel-cladding bond layer. These conditions apply to 

the current work: 

a) With TCs welded onto bare as-fabricated cladding, conduct the thermal benchmark test for a hold 

time selected to grow 10-to-40-μm-thick oxide layers on inner and outer surfaces of the cladding 

(see Fig. 7 for temperature history to grow 40-μm oxide layers).

b) Cool to 300oC and repeat thermal benchmark test using the same controller parameters as were used 

in (a); compare the two sets of results (see Fig. 8 for 40-μm pre-oxide layers) with emphasis on the 

maximum temperature at the end of the rapid temperature rise (first peak), the time to reach the 

hold temperature, and the hold temperature. 

c) If necessary, increase the holder control temperature to achieve the desired hold temperature for 

cladding with pre-transient oxide layers (not necessary for Fig. 8 results).

d) Install the calibrated test train in the in-cell furnace and conduct a metallurgical benchmark test 

with as-fabricated cladding without the TCs welded onto the cladding and with the hold time 

chosen to give a weight gain equivalent to about 10% CP-ECR; if the out-of-cell and in-cell weight 

gains and converted ECR values are in good agreement, then use of the test train in the in-cell 

furnace has been validated.

e) Use the results of the thermal benchmark in (b) or (c) – the one with oxide layers grown on the 

inner and outer surfaces prior to initiating the temperature ramp – to select test times to give the 

desired CP-ECR values and to interpret the experimental results for high-burnup cladding.
21

Figure 8. Thermal histories for bare and pre-oxidized (40-μm) ZIRLO cladding tested in the same test train w

parameters. At 150-s into the heating phase, the average temperature for the pre-oxidized sample wa

900

800

700

600

500

400

300

1300

1200

1100

1000

500 550 600 650 700 750 80

Temperature (oC)

Time (s)

ZLU#145: bare ZIRLO

ZLU#145A: pre-oxidized (40 μm) ZIRLO
22

Based on the results shown in Fig. 8, pre-oxidizing the cladding resulted in a 50°C decrease (from 1110°C 

to 1060°C) in the pre-1200°C peak and a longer time (150 s compared to 100 s) to reach 1200°C. 

Subsequently, it was verified that the results for 12-μm pre-oxide layers were essentially the same as the 

results for 40-μm pre-oxide layers (OD and ID). From a thermal perspective, it is likely sufficient to use a 

10-μm pre-oxide layer in thermal benchmark tests for high-burnup cladding. This would reduce the 

oxidation rate and the rapid heating rate of bare cladding due to the effects of the heat of oxidation. 

2.4 Ductility Determination Using Ring-Compression Tests

2.4.1 Background information and general description

The procedure for conducting ring-compression tests (RCTs) and the methodology for determining 

ductility are documented in Refs. 1–3. Reference 5 contains a detailed analysis of RCT loading and data 

interpretation. Some of this material is repeated below for the convenience of the reader.

Argonne has exposed as-fabricated, pre-hydrided, and high-burnup cladding samples to two-sided 

oxidation and quench. Short (≈8-mm-long) rings sectioned from these samples are subjected to ringcompression tests at RT and 135°C and at 2 mm/min (0.033 mm/s) displacement rate to determine the 

ductile-to-brittle transition oxidation level as a function of hydrogen content. 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.

Figure 9 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. Oxidized 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 (for rings with ≤1 tight crack) and/or by the offset displacement (δp) method (for rings with ≥1 

loose cracks). These displacements are normalized to the cladding metal outer diameter (Do for asfabricated and pre-hydrided cladding and Dmo for corroded cladding) prior to oxidation. For two-sided 

oxidized samples, cracking through the wall and along the length of the sample is equally probable at 

inner-surface locations under the load (12 o’clock position) and above the support plate (6 o’clock 

position). The maximum bending moment (M), tensile hoop stress (σθ), and tensile hoop strain (εθ) occur 

at these two locations. Bending stresses at these locations vary from tensile (inner surface) to compressive 

(outer surface). 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. These stresses vary from 

compressive (inner surface) to tensile (outer cladding). Cracking seldom initiates and propagates at the 3 

or 9 o’clock position, prior to cracking at the 12 and/or 6 o’clock position because of the lower tensile 

stress for intact rings. The primary bending stress is in the hoop direction. However, due to the large 

length-to-thickness ratio (L/h >10), a smaller axial stress is induced throughout the ring (excluding the 

free ends). This stress-inducing condition in the axial direction is referred to as plane-strain constraint.

Prior to conducting data-generating tests, RCTs are conducted with non-oxidized, as-fabricated cladding 

to ensure that the test machine is operating properly. Numerous benchmark tests have been conducted 

with 15×15 and 17×17 cladding rings at RT and 135°C to total displacements of up to 2 mm. As results 

for these homogeneous samples are easier to interpret, the data for one of these tests are used to illustrate 

methodology and measurement of key parameters. Figure 10 shows the RCT load-displacement curve for 

as-fabricated M5 cladding compressed to 1.7-mm displacement at RT. Dimensions of the ring were: L = 

7.7 mm, Do =9.48 mm, and h = 0.627 mm. Elastic properties for M5 at RT are: Young’s modulus E = 

91.9×104 MPa and Poisson’s ratio ν = 0.37. Figure 10 is discussed further in the following subsections. 
23

Figure 9. Schematic of ring-compression test sample and loading. The displacement (δ) rate is 

controlled and the response force (P) is measured.

Figure 10. Load-displacement curve for as-fabricated (AF) M5 cladding compressed at RT to 

1.7-mm total displacement.

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.0 0.5 1.0 1.5 2.0 2.5 3.0

Displacement (mm)

Load (kN)

KLM = 0.99 kN/mm

AF M5 Sample 116A3 at RT

7.7-mm Sample Length

Offset Displacement = 1.19 mm

Corrected Offset Displacement = 1.11 mm

Permanent Displacement = 1.05 mm

KLC = 1.30 kN/mm

1.19 mm

1.11 mm
24

2.4.2 Ductility and ductility criteria

The standard offset displacement methodology calls for mathematical unloading at the same slope as the 

measured loading slope (KLM) just before failure or at the end of the test if no failure occurs. This is 

illustrated in Fig. 10 by the dashed black line at the end of the loading portion of the test. The distance 

between these two lines is the offset displacement (δp), which is 1.19 mm in Fig. 10. However, the 

unloading slope is not precisely equal to the loading slope for compressed rings, and the standard 

methodology gives a higher measured plastic displacement than the actual permanent displacement (dp) 

measured after unloading (1.05 mm based on the difference between pre- and post-test micrometer 

measurements). For the load-displacement curve in Fig. 10, δp – dp = 0.14 mm. Normalizing δp and dp to 

the initial outer diameter (9.48 mm), gives offset and permanent strains of 12.6% and 11.1%, respectively. 

Based on over 50 benchmark tests with as-fabricated cladding, the inherent error between offset and 

permanent displacements was found to be ≤0.2 mm for total displacements up to 2 mm. This equates to an 

absolute strain error of ≤2%. As the total displacement decreases, along with δp, the error decreases.

For oxidized cladding tested to failure, the RCT is manually terminated after the first significant load drop 

(>30%) in an attempt to preserve the ring geometry. If, upon unloading, the ring has a single tight crack, 

then the post-RCT diameter can be measured with reasonable accuracy and dp can be determined from the 

difference between pre-test and post-test diameter along the loading direction. This situation is shown in 

the load-displacement curve (Fig. 11) for as-fabricated ZIRLO cladding oxidized (one-sided) at 1200°C 

to 17% CP-ECR prior to quench at 800°C. The test was conducted at reference temperature (135°C) and 

displacement-rate (0.033 mm/s) conditions. Offset and permanent displacements were 0.18 mm (1.9% 

strain) and 0.11 mm (1.2% strain). Based on the permanent-strain ductility criterion (dp/Do ≥ 1.0%) 

established in this program, the ring was classified as ductile. The 1.0% permanent strain criterion is 

based on an error assessment of post-RCT diameter readings, especially for corroded cladding with one 

tight crack. Corrosion-layer flake-off and mismatch of cracked surfaces during unloading account for 

most of the error in post-RCT diameter measurements for high-exposure cladding. 

Figure 11. Load-displacement curve for as-fabricated ZIRLO oxidized (one-sided) at 1200°C to 

17% CP-ECR. The sample had one tight through-wall crack at the 6 o’clock 

position.

0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0

Displacement (mm)

Load (kN)

0.18 mm

AF ZIRLO Oxidized to 17% CP-ECR

RCT at 135°C and 0.033 mm/s

Offset Strain = 1.9%

Permanent Strain = 1.2%

KLM = 0.97 kN/mm
25

Data trends were also considered in setting the permanent strain criterion for ductility. Figure 12 shows 

permanent-strain results for 17×17 Zry-4 cladding oxidized at 1200°C, quenched at 800°C, and 

compressed at RT and 135°C. The RT results show a sharp decrease in permanent strain from 20% at 

5.5% CP-ECR to 2.2% at 7.8% CP-ECR. The permanent strain drops to 0.5% at 10% CP-ECR and then 

oscillates within the range of 0.5±0.02% from 10% to 20% CP-ECR even though the average oxygen 

content in the cladding continues to increase to about 0.57% at 18% CP-ECR. The observation that the 

permanent strain does not go to zero as the cladding continues to embrittle supports the approach of 

setting a 1% limit on permanent strain, as values below 1% appear to be in the “noise” of uncertainty. The 

RCT results for 135°C test temperature show a sharp decrease in permanent strain from 11% to 13% CPECR and a more gradual decrease from 13% to 19% CP-ECR. The average permanent strain is 1% at 

17% and 19% CP-ECR. These results are consistent with the oxidation level at which the beta-layer 

oxygen content reaches the saturation level. The embrittlement oxidation level is assessed to be 19% for 

this material oxidized at 1200°C based on 3 data points at 17% and 2 data points at 19% CP-ECR. In Ref. 

3, it is recommended that at least three RCT data points be generated at 17%, 18%, 19%, and 20% CPECR to better determine the transition oxidation level to within 1% CP-ECR for as-fabricated cladding. 

Figure 12. Permanent strain vs. CP-ECR for 17×17 Zry-4 oxidized at 1200oC, cooled at ≈13oC/s 

to 800oC, quenched, and ring-compressed at RT and 135oC. 

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. 13, which gives offset minus permanent 

strain (δp/Do – dp/Do) as a function of oxidation level (in % CP-ECR) for rings with permanent strains in 

the range of 1.0% to 2.3% (near ductility threshold). By setting dp/Do to the 1% ductility limit and adding 

it to the 1-σ upper bound in Fig. 13, the offset-strain ductility criterion can be expressed as:

δp/Do ≥ 1.41% + 0.1082 CP-ECR (8) 

In using Eq. 8, the average measured 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. 8 (rounded to 

nearest 0.1%) for the material to be classified as ductile. It should be noted that use of the offset-strain 

criterion (Eq. 8) may result in lower transition oxidation levels than would be determined based on the 1% 

permanent-strain criterion. The offset strains corresponding to Fig. 12 were 2.8% at 17% CP-ECR 
26

Figure 13. Difference between offset and permanent strains as a function of oxidation level for 

as-fabricated, pre-hydrided, and high-exposure cladding alloys oxidized at 1200°C 

and compressed 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%.

(vs. 3.2% limit) and 2.5% at 19% CP-ECR (vs. 3.5% limit), which imply <17% transition CP-ECR. This 

conservatism is needed to accommodate the large data scatter seen in Fig. 13.

2.4.3 Measured vs. calculated loading slopes 

The ring in Fig. 9 behaves like a spring in the elastic deformation regime with spring constant KLC such 

that P = KLC δ. The spring constant can be calculated using advanced strength-of-materials methods:

KLC = {(E L)/[1.79 (1 – ν2

)]} (h/Rmid)

3

, (9) 

where Rmid = (Do – h)/2 is the mid-wall cladding radius. For the benchmark case shown in Fig. 10, KLC = 

1300 N/mm = 1.30 kN/mm. However, the stiffness measured from the load-displacement curve is less 

(0.99 kN/mm). While the measured load reflects the actual load on the sample, the measured 

displacement includes “machine” displacement, as well as sample displacement in the elastic deformation 

regime. The system acts like two springs in series. Let Kmach be the effective stiffness of the machine. The 

measured stiffness is:

KLM = (KLC/Kmach + 1)-1 KLC (10) 

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
27

For 0.99-kN/mm KLM and 1.30 kN/mm KLC, Kmach is calculated to be 4.1 kN/mm, which is close to the 

estimated 5 kN/mm for the Instron 8511 machine used in the current work. For stiffer machines and/or 

samples with lower stiffness values (e.g., 4-PBT samples discussed in Section 2.6), KLM → KLC. 

Using the standard methodology, the difference between measured and calculated stiffness values is not a 

factor in determining offset displacement and ductility. However, some caution should be exercised in 

integrating the area under RCT load-displacement curves to determine failure energy. Unless there is a 

correction to the measured slope, which is difficult to do for oxidized samples, the integration will result 

in higher calculated failure energy than is appropriate for the sample (e.g., as in Ref. 22). 

2.5 LOCA Integral Tests with Pressurized Cladding Samples

Out-of-cell LOCA integral tests with as-fabricated and pre-hydrided cladding samples are relatively 

straightforward with regard to sample preparation and test conduct. Each cladding sample was partially 

filled with ≈25-mm-long zirconia pellets to simulate the heat capacity of the fuel. There was a clearance 

of about 13 mm between the top zirconia pellet and the inner surface of the top end fixture to prohibit 

contact during sample bending. The test train was supported at the top to minimize specimen bowing. The 

quartz tube encasing the test train provided an enclosed volume for steam flow and water quench, both of 

which were introduced through the bottom of the unit. The test train was centered within the quartz tube 

by means of two perforated spacer disks. Centering is very important as four vertical infrared lamps are 

focused within the furnace to heat the specimen. Swagelok fittings were used above the specimen to 

connect to the high-pressure gas line and top pressure gauge and below the specimen to connect to the 

bottom pressure-gauge line. The total gas volume above the fuel column was 10 cm3

, most of which was 

outside the heated zone. Four Type S TC lead wires were fed in through the top. Two of the TCs were 

spot-welded at the specimen mid-plane, 180° apart. The other two were spot-welded 50 mm above and 50 

mm below the mid-plane at the same angular orientation as one of the mid-plane TCs. The signal from the 

top TC was used to control the furnace power to achieve the desired temperature ramp, hold temperature, 

and cooling rate prior to quench. Although the uniform heating zone of the furnace is nominally 250 mm, 

the actual uniform temperature region for a non-deformed sample ranged from 100-125 mm, centered on 

the sample mid-plane. For out-of-cell tests, hands-on assembly was used, and bare-wire TC beads were 

welded directly onto non-irradiated tubing. Reference 1 contains schematic diagrams, photographs, and a 

detailed description of the LOCA apparatus and integral test train. 

During early-phase testing of 17×17 ZIRLO cladding in the current work, changes were made to sample 

length, pellet stack height, location of the TCs, and bottom pressure transducer. It was observed that a 

secondary balloon appeared near the top plenum region in some tests. In order to eliminate this secondary 

balloon, the plenum region was moved to a higher elevation above the heated zone by extending the 

sample length from 305 mm to 318 mm. An additional half pellet was added to increase the pellet stack 

height and to maintain the 13-mm-long plenum. Also, the survival rate for the two mid-plane TCs was 

low due to the disruptive effects of ballooning and rupture. For mid-plane TCs that did survive, results 

indicated that the circumferential temperature variation was greater than the axial temperature variation. 

As such, in the current work two TCs were welded onto the sample about 57 mm above the mid-plane at 

0° and 180° orientations and two TCs were welded onto the sample at 44 mm below the mid-plane at 90° 

and 270°C. The TC above the mid-plane at 0° was used as the control TC. Also, the pressure transducer at 

the bottom of the test train was removed from the system. This pressure transducer was difficult to install, 

had a short lifetime, and was needed only for tests with fueled cladding samples in order to determine the 

gas communication between the top and bottom of the fuel column. 
28

Real-time data of interest included pressure and temperature histories, rupture temperature, and pressure 

at which rupture occurred. Following a test, the sample was photographed (low and high magnification), 

micrometer measurements were made along the length of the sample at two orientations to determine the 

outer-diameter profile, the maximum outer circumference was measured, and the rupture opening was 

characterized in terms of axial location, width, and length. Prior to destructive examination, samples (with 

the pellets removed) were subjected to four-point bend tests at 135°C to determine maximum bending 

moment, failure energy, and plastic deformation (see Section 2.6). Destructive examinations were then 

performed to determine oxide and alpha layer thicknesses for inner- and outer-cladding surfaces 

(metallography), circumferential and axial distribution of oxygen (LECO oxygen analyzer), and 

circumferential and axial distribution of hydrogen (LECO hydrogen analyzer). Additional details about 

the LOCA integral test setup and data are provided in Sections 5 and 6. 

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

The 4-PBT (Fig. 14) was used in the current work to determine post-quench strength, failure energy, and 

plastic deformation of ballooned, ruptured, oxidized, and quenched LOCA integral-test samples. It is 

described in more detail in Ref. 5. The 4-PBT has advantages over the 3-PBT for cladding samples with 

axial variation in 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.

Figure 14. Schematic of four-point-bend test (4-PBT) loading. The bending moment is constant 

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

measured. 

For the Argonne 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, M in N•m, is given in terms of the measured load, 

P 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. 14, 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. 
29

The 4-PBT sample is much more flexible than the RCT 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. Finite-element-analysis (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 FEA-calculated 

load-displacement curves for the 15×15 Zry-4 benchmark tests at RT were in excellent agreement with 

the measured load-displacement curves. 

Initially, the maximum elastic strain vs. displacement (δ) for the 4-PBT sample was set equal to the 

normalized RCT displacement (δ/Do) to determine that the 4-PBT displacement rate should be 2 mm/s. 

This was later reduced to 1 mm/s on the basis of 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). 

Thus, in order to conduct 4-PBTs at the same maximum elastic strain rate as was used in RCTs, the 

displacement rate was increased from 0.033 mm/s (RCT) to 1 mm/s (4-PBT).

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. 

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. Although loads were not measured during 

unloading, the subsequent reloading stiffness was measured and found to be within 1% of the original 

loading stiffness. 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 RCT offset displacement. 

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

sample by using elementary beam theory. However, the measurement of radius of curvature is tedious and 

cannot be done accurately for an oxidized sample tested to failure.

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

displacements beyond the elastic regime to FEA-calculated maximum plastic strains in an effort to 

determine the normalization length (Ln) to convert the measured plastic displacement to strain. The 

calculation was performed for both 15×15 and 17×17 Zry-4 geometries. The FEA results indicated that 

the normalization length varied with plastic strain. Unlike the RCT for which the outer diameter is used 

for normalization of plastic displacement, there is no equivalent normalization factor for the 4-PBT. 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. 14. 

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 deformation. Taking into 

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

of plastic deformation: δp ≥ 0.1 mm implies plastic deformation. In addition to plastic deformation, 

material strength (as measured by the maximum bending moment) and failure energy (area under loaddisplacement curve) can be determined.

31

3 EMBRITTLEMENT OF NON-BALLOONED CLADDING

In the previous work, ductile-to-brittle transition oxidation levels were determined for as-fabricated 

cladding alloys, for pre-hydrided Zry-4, and for defueled Zry-4, ZIRLO, and M5 from high-burnup fuel 

rods. These results were documented in Ref. 1, and they are summarized in the following subsections. 

The terms “ductile-to-brittle transition” and “embrittlement” are defined as the highest CP-ECR value 

(rounded to the nearest percent) for which the cladding maintains borderline ductility. Oxidation to levels 

greater than these CP-ECR values leads to brittle cladding behavior. Samples were oxidized at 1200oC 

and quenched at 800oC, prior to determining post-quench ductility (PQD) using ring-compression tests

conducted at 135oC and 0.033 mm/s. For as-fabricated 17×17 cladding alloys, LOCA oxidation tests were 

conducted with and without quench at 800oC to explore the effects of quench on PQD. The results 

indicated that quench at 800oC had no significant effect on the ductility of as-fabricated cladding alloys. 

However, for pre-hydrided and irradiated cladding, quench at 800oC significantly reduced ductility and 

the embrittlement oxidation level as compared to samples that were cooled without quench. 

This previous work did not include testing of high-burnup cladding samples with pre-test cladding-metal 

hydrogen contents (CHM) between 140 and 540 wppm. Thus, additional tests were conducted in the 

current work to determine the embrittlement oxidation level as a function of hydrogen content in the midrange region of 200 to 400 wppm. These tests were performed with pre-hydrided, non-irradiated ZIRLO

cladding and with defueled ZIRLO cladding from the same high-burnup rod used in the previous study. 

3.1 Embrittlement of As-fabricated Cladding

Table 7 gives embrittlement oxidation levels (CP-ECR values) for as-fabricated cladding alloys based on 

the permanent-strain criterion (<1% implies embrittlement and ≥1% implies ductility). For multiple tests 

conducted at the same CP-ECR level, the average permanent-strain value was compared to 1%. If 

interpolation or extrapolation was required, the embrittlement CP-ECR was determined to the nearest 1% 

oxidation level. Table 7 only includes ductility data at oxidation levels close to the embrittlement CPECR levels. Offset-strain values are also provided such that they can be compared to the criterion in Eq. 

8. These ductility data come from Ref. 1, where further details can be found at the locations indicated in 

the table. Where available, ductility data for slow cooling (i.e., no quench) were included to improve 

statistical significance because slow cooling in these tests had no observable effect on ductility. 

With the exception of the vintage HBR-type Zry-4 cladding, the modern cladding alloys exhibited an 

embrittlement oxidation level in the narrow range of 19-20% based on the permanent strain criterion. 

These results are significant in that the embrittlement threshold for these different cladding alloys and 

cladding materials (e.g., two different designs for Zry-4) was independent of alloy type and fabrication 

differences. An extensive database was generated for HBR-type 15×15 Zry-4, not all of which was 

tabulated in Ref. 1. Its embrittlement oxidation level was determined to be 16% on the basis of data 

summarized in Table 7 and presented in Fig. 15a. 

Except for 17×17 M5, use of the offset-strain ductility criterion (Eq. 8) resulted in an apparent decrease in 

embrittlement CP-ECR value: from 20% to 17% for 15×15 Zry-4, from 20% to 19% for 10×10 Zry-2, 

from 19% to 18% for 17×17 ZIRLO, from 19% to 16% for 17×17 Zry-4, and from 16% to 15% for HBRtype 15×15 Zry-4 (see Fig. 15b for offset strain vs. CP-ECR). For 17×17 Zry-4 additional offset-strain 

data (e.g., 7.5±3.1% offset strain at 13% CP-ECR) from Ref. 1 (Tables 11 and 43) were used to determine 

the 16% embrittlement oxidation level. Therefore, the use of offset strain in safety analyses to determine 

embrittlement threshold would appear to be conservative. 
32

Table 7. Ductile-to-brittle transition oxidation levels (embrittlement CP-ECR values) for asfabricated cladding alloys oxidized at 1200°C and ring-compressed at 135°C. Most data 

points used to determine ductility were for samples quenched at 800°C.

Cladding Material

Tested

CP-ECR,

%

Permanent

Strain,

%

No. of

Data

Points

Offset

Strain,

%

No. of

Data

Points

Embrittlement

CP-ECR,

%

17×17 M5

(34,45)a

16

20 

2.8

1.0 

1

2 

7.6±2.8

3.5±0.5 

(3.6%)b

4

2 

20

15×15 Zry-4

(16)a

17

20 

1.8±0.4

1.1±0.4 

2

2 

3.3±0.8

2.4±0.4 

(3.6)b

2

2 

20

10×10 Zry-2

(22)a

17

20 

2.8

1.0±0.1 

1

2 

6.4±1.3

1.7±0.2 

(3.6%)b

2

2 

20

17×17 ZIRLO

(27,44)a

17

19 

1.8±0.6

1.1±0.2 

5

2 

4.5±1.8

2.6±0.8 

(3.5%)b

7

2 

19

17×17 Zry-4

(11,43)a

17

19 

1.1±0.5

1.0±0.2 

5

2 

2.4±0.8

2.5±0.8 

(3.5%)b

5 

2 

19

15×15 Zry-4

HBR-type

(current Fig. 15)

15

16 

1.5±0.4

1.1±0.3 

6

7 

3.0±0.7

2.9±0.6 

(3.1%)b

6

7 

16 a

Table number(s) in Ref. 1 containing most of the data. 

b

Offset strain ductility criterion calculated with Eq. 8 at the indicated CP-ECR.

(a) Permanent strain (b) Offset strain

Figure 15. PQD results for 15×15 Zry-4 (HBR-type) following oxidation at 1200°C and quench 

at 800°C: (a) permanent strain and (b) offset strain.

0

1

2

3

4

5

12 13 14 15 16 17

CP-ECR (%)

Permmanent Strain (%)

HBR-type: 1204±16oC

HBR-type: 1200±16oC

HBR Archive: 1204±16oC

Ductility Limit

0

1

2

3

4

5

6

7

12 13 14 15 16 17

CP-ECR (%)

Offset Strain (%)

HBR-type: 1204±16oC

HBR-type: 1200±16oC

HBR Archive: 1204±16oC

Ductility Limit
33

3.2 Embrittlement of Pre-hydrided Cladding

3.2.1 Review of previous results

Table 8 summarizes previous PQD results for pre-hydrided 17×17 Zry-4 and HBR-type 15×15 Zry-4 

oxidized at 1200°C to 5–10% CP-ECR followed by quench at 800°C. LOCA oxidation tests were 

conducted at fixed oxidation levels for samples with variable hydrogen content in the cladding metal 

(CHM). For 17×17 Zry-4 oxidized to 10.3% CP-ECR, CHM was between 274 and 335 wppm. Based on 

extrapolation of offset strain data for 209 and 274 wppm, the embrittlement CHM was determined to be 

280 wppm at about 10% CP-ECR. For 17×17 Zry-4 oxidized to 7.9% CP-ECR, the embrittlement CHM

was between 314 and 412 wppm. 

At 7.7% CP-ECR, the embrittlement CHM for HBR-type Zry-4 was between 360 and 386 wppm. Although 

results at about 8% CP-ECR were consistent for these two materials, the more extensive database for 

HBR-type Zry-4 was used to determine the embrittlement CHM at about 8% CP-ECR: 370-wppm based on 

interpolation. PQD results depended on heating rate and time spent at elevated temperature. For example, 

HBR-type Zry-4 samples subjected to a fast heating rate with a brief overshoot to 1216°C (Fig. 109 [1]), 

were brittle with 335 wppm hydrogen following oxidation to 7.6% CP-ECR and quench at 800°C (Table 

51 [1]).

Table 8. Embrittlement pre-test hydrogen content (CHM) as a function of oxidation level (CPECR) for pre-hydrided Zry-4 cladding oxidized at ≈1200oC and quenched at 800oC. 

Material

Tested

CPECR,

%

Pre-test

CHM, 

wppm

Permanent

Strain,

% 

No. of

Data

Points

Offset

Strain, 

% 

No. of

Data

Points

Embrittlement

CHM, 

wppm

17×17 

Zry-4 

(46)a

10 209

274 

335

3.3

– 

0.4

1

0 

1

5.5

2.8 

(2.5)b

0.6

1

1 

1

280 

15×15 

Zry-4 

HBR-type

(48)a

8 360

386 

–

0.5±0.1 

0

2 

5.0

(2.3)b

1.2±0.7 

1

3 

370 

15×15 

Zry-4 

HBR-type

(47,53)a

5 470±8

500±16 

636 

–

– 

0.4 

0

0 

1 

4.2±1.5

3.0±2.3 

(1.7)b

0.5 

4

3 

1 

530 

a

Table number(s) in Ref. 1 containing the data. CP-ECR values include oxidation during cooldown. 

b

Offset strain criterion calculated with Eq. 8.

A large number of PQD tests were conducted with pre-hydrided, HBR-type 15×15 Zry-4 oxidized to 4.7–

5.7% CP-ECR. The oxidation tests were conducted in four separate campaigns using different test trains 

and temperature control parameters. PQD results were sensitive to heating rate from 1000°C to 1180°C 

with higher heating rates causing enhanced embrittlement. In general, the target CP-ECR value was 

reached during the slow ramp rate from 1180°C to 1200°C. For 5% CP-ECR samples, the extrapolated 

CHM leading to embrittlement was 530 wppm. This result is consistent with the embrittlement exhibited 

by the 5.2% CP-ECR sample with 636 wppm hydrogen. 
34

At the time Ref. 1 was published, there was an apparent discrepancy between CEA (at Saclay) and 

Argonne data for pre-hydrided (about 600 wppm) Zry-4 samples oxidized to about 6% CP-ECR and 

quenched at 600°C or 700°C: CEA observed a significant enhancement in ductility while Argonne did not 

(see Table 54 in Ref. 1). Argonne speculated that the difference was due to the very slow cooling rates in 

the CEA furnace (e.g., about 1000 s to cool from 1200°C to 800°C) compared to the more typical fast 

cooling rates characteristic of the Argonne furnace (e.g., about 40 s to cool from 1200°C to 800°C). 

Argonne samples were sent to CEA for detailed characterization. Concurrently, CEA modified their 

apparatus by adding He cooling following steam oxidation to achieve cooling rates (10°C/s) comparable 

to Argonne’s cooling rates. Using the faster cooling rate, CEA observed results comparable to Argonne 

results in that no enhancement in ductility was observed at the lower quench temperatures [11]. 

3.2.2 Current results

Oxidation and PQD tests were conducted with pre-hydrided ZIRLO samples after tests with high-burnup 

ZIRLO. However, results for pre-hydrided ZIRLO are presented before results are presented for highburnup ZIRLO (see Section 3.3). Tests with pre-hydrided samples were also conducted after the draft 

procedure [18] was submitted to NRC for conducting PQD tests. Thus, there was a dual purpose to these 

tests: (a) to generate PQD data for pre-hydrided ZIRLO with hydrogen contents comparable to the highburnup ZIRLO samples and (b) to test the adequacy of the detailed procedures documented in Ref. 18. 

Materials and sample preparation 

ZIRLO-2008 cladding was used to prepare pre-hydrided segments (≈80-mm long). Target hydrogen 

contents were 300±50 wppm. Based on trial efforts, it was very difficult to achieve these hydrogen levels 

with non-etched, belt-polished (outer surface), and sand-blasted (inner surface) segments. Hydrogen 

levels were either too low or too high. Although pre-etching segments with an HF-containing acid bath 

was not recommended in Ref. 18, additional trial runs were conducted with samples etched in a 1% HF 

solution (with equal parts of nitric acid and distilled water) for periods of 30 s to 120 s (reference time). 

This greatly improved the pre-hydriding process such that average hydrogen contents of 300 to 470 wppm 

(based on segment weight gain following pre-hydriding) were achieved. However, test samples (≈25-mm 

long) from etched as-fabricated cladding and pre-hydrided cladding experienced early breakaway 

oxidation and significant hydrogen pickup during relevant oxidation times at 1200°C. This artifact was 

eliminated by polishing the inner and outer surfaces of samples prior to conducting oxidation-quench 

tests. Polishing removed about 1 μm from these surfaces, which was sufficient to remove surface and 

near-surface fluoride impurities. As an extra caution, end caps were inserted into segments such that only 

the outer surface was exposed to the HF-containing bath. Both inner and outer surfaces were still polished 

to remove possible impurities that may have been picked up during the pre-hydriding process, which 

included a final long anneal in argon gas to achieve a more uniform hydrogen concentration. Based on the 

high purity levels of the gases used, polishing of the inner surface may have been unnecessary.

Parameters for hydrogen charging were modified over a five-year period to gain better control over the 

process. The hydrogen content of the gas was reduced from 100% to 30% (plus 70% He), and a transition 

was made from flowing to static hydrogen. The final process used to pre-hydride ZIRLO segments was: 

(a) evacuation of the system at RT, (b) out-gassing at elevated temperature (>400°C), (c) quick 

installation of cladding segment at RT, (d) purge of system at RT with flowing Ar gas (24 h), (e) 

evacuation at RT, (f) flowing hydrogen gas mixture through system at RT (4 h), (g) closure of valves to 

seal off hydrogen within furnace chamber, (h) controlled heating of furnace to 400°C hold temperature, 

(i) hold at 400°C in stagnant hydrogen mixture for target time, (j) turn off furnace at end of hold time, and 

(k) evacuation of system during initial cool down. Following pre-hydriding, cladding segments were 
35

subjected to flowing Ar at 400°C for about 24 h with five temperature cycles between 400°C and 300°C 

to decrease axial and circumferential gradients in hydrogen concentration. 

Pre-hydrided ZIRLO segments were sectioned as shown in Figs. 16a (A13) and 16b (A15) to produce 

oxidation test samples. Based on weight gain, average hydrogen contents were determined to be 302 

wppm and 471 wppm for segments A13 and A15, respectively. As shown in Fig. 16, hydrogen contents 

measured at four axial locations per segment tended to be less than weight-gain values. Hydrogen-content 

measurements on both sides of the oxidation samples were mass-averaged to estimate the CHM levels in 

the samples. Inner and outer sample surfaces were then polished. Length (L), outer diameter (Do), wall 

thickness (hm), and weight (W) were measured after polishing. Prior to weighing, the samples were 

thoroughly cleaned and dried. Table 9 summarizes the results of these measured parameters.

(a) Segment A13 

(b) Segment A15

Figure 16. Sectioning diagrams for pre-hydrided ZIRLO segments: (a) Segment A13 and (b) 

Segment A15. Nominal lengths are shown for the oxidation test samples.
36

Table 9. Pre-test estimates of CHM levels, measured dimensions, and measured weights for prehydrided ZIRLO oxidation test samples. 

Test ID

ZLUP# Sample ID

Estimated

CHM, wppm

L,

mm

Do,

mm

hm,

mm

W,

g 

6

5 

7 

A13C1

A13C3 

A13C5 

226±16

256±46 

263±36 

23.6

22.7 

27.3 

9.48

9.48 

9.50 

0.567

0.567 

0.560 

2.4341

2.3486 

2.8250 

4

3 

8 

A15C1

A15C3 

A15C5 

245±23

231±8 

314±91 

23.8

22.0 

22.8 

9.47

9.47 

9.47 

0.564

0.565 

0.560 

2.4500

2.2632 

2.3139 

Thermal and weight-gain benchmark results

The standard procedure described in Section 2.6 was used to perform thermal and weight-gain 

benchmarks. Controller parameters were determined earlier for oxidation tests with irradiated ZIRLO (see 

Section 3.3). This was done to allow good comparison of PQD results for pre-hydrided and irradiated 

cladding samples. An as-fabricated ZIRLO sample, with two TCs welded onto the sample, was used for 

the thermal benchmark test. The temperature histories for the two welded TCs are shown in Fig. 17. After 

151 s from the beginning of the temperature ramp, the average sample temperature was 1200±5°C. The 

test was repeated without the welded TCs. The measured weight gain (6.80 mg/cm3

) was only about 3% 

higher than the CP weight gain (6.62 mg/cm3

). The absence of temperature overshoot, the small 

circumferential temperature variation, and the excellent agreement between measured and calculated 

weight gain were well within limits established in Section 2.3. 

Oxidation tests

Choosing test times and CP-ECR values to bracket ductile and brittle behavior was very challenging for 

pre-hydrided oxidation test samples because no two samples appeared to have the same average CHM, the 

axial variation in CHM was neither linear nor predictable, and because of the overall uncertainty in the 

estimated CHM. The average (and 1–σ variation in) CHM could not be determined accurately until after the 

oxidation test was completed. Given the post-oxidation sample weight gain and LECO-measured 

hydrogen content, CHM was determined in the following manner for each cladding ring subjected to an 

RCT. 

In general, the LECO-measured hydrogen content is: CH = (HS/MS) × 106

, where HS is the mass of 

hydrogen evolved from the sample and MS is the mass of the sample. Therefore for pre-hydrided cladding 

samples before oxidation, the sample is all metal such that HS = HM, MS = MM and CH = CHM. But for 

samples after oxidation, the sample mass is MS = MM + WG, where WG is the oxidation weight gain in 

mass units. Hence, the LECO-measured hydrogen content for pre-hydrided samples after oxidation is 

(CH)pox = (HM/[MM+WG]) × 106

. This value can be multiplied by a mass-correction factor to determine 

CHM: CHM = (1 + WG/MM) (CH)pox = (1 + Wg/100) (CH)pox = fWg (CH)pox, where Wg is weight gain in percent 

and fWg is the weight-gain correction factor. 
37

Figure 17. Thermal history determined from benchmark test for oxidation tests to be conducted with pre-hydri

at 800°C, which was used for PQD tests, is not shown in this figure because benchmark samples are 

quench.

900

800

700

600

500

400

300

1300

1200

1100

1000

500 550 600 650 700 

Time (s)

Temperature (°C)

TC4 on sample

TC5 on sample

T = 1200±5oC

Δt = 151 s

CP-ECR = 10%
 

38 

Table 10 summarizes the six oxidation tests that were conducted in the CP-ECR range of 8 to 11%. RCT 

results are given in Table 11. Oxidation tests were run sequentially such that ductility and pre-test CHM

were determined for rings from the previous test prior to choosing the CP-ECR level for the subsequent 

test. Depending on sample length, short end pieces were removed and at least two RCT rings (6.7–8.6 

mm long) were sectioned from each oxidation sample. Rings were subjected to RCTs at 135°C and 2 

mm/min (0.033 mm/s) displacement rate. For 13 of the 16 rings, the permanent displacement could not be 

determined because simultaneous through-wall cracks occurred at multiple (2 to 4) circumferential 

positions. For the three RCT samples with only one through-wall crack, two were brittle and one was 

highly ductile. Thus, the ductile-to-brittle transition CP-ECR had to be determined based on the offsetstrain criterion (Eq. 8). 

The RCT data in Table 11 were reorganized and grouped according to CP-ECR and CHM values. Results 

for 240-wppm and 200-wppm CHM are given in Table 12. For LOCA samples oxidized to 8%, 9% and 

11% CP-ECR, three rings (A13C3B, A15C5C, and A13C5B) had average hydrogen contents of about 

240 wppm. Fourteen CHM values (4 to 6 per ring) were mass averaged to determine CHM for the three 

rings: 240±33 wppm. Minimum and maximum measured values were 169 and 287 wppm. Seven rings 

(A15C3B, A15C3C, A15C1B, A15C1C, A15C1D, A13C1B, and A13C1D) had average CHM values of 

about 200 wppm. Thirty-eight CHM values (4 to 6 per ring) were mass averaged to determine CHM for the 

three rings: 200±7 wppm. 

The RCT results for rings with 240 wppm are relatively straightforward to interpret. As shown in Table 

12 and Fig. 18, rings oxidized to 8% and 9% CP-ECR were highly ductile, while the ring oxidized to 11% 

CP-ECR was very brittle. The transition CP-ECR is >9% and <11%. Based on the trend curve shown in 

Fig. 18, 10% is a reasonable transition oxidation level for ZIRLO with 240-wppm CHM. If more samples 

had been available with 240 wppm CHM, a confirmation test at 10% CP-ECR would have been conducted. 

The RCT data set for 200-wppm ZIRLO is more robust at oxidation levels of 10% and 11% CP-ECR. The 

two high-ductility data points at 217 wppm and 8% CP-ECR can be included as a reasonable lower bound 

on the anticipated ductility of 200-wppm ZIRLO oxidized to 8% CP-ECR. On the basis of the data trend 

curve shown in Fig. 18, it is expected that 200-wppm ZIRLO would be brittle if it were oxidized to 12% 

CP-ECR. Thus, the transition CP-ECR for 200-wppm ZIRLO was chosen to be 11%. If additional 200-

wppm ZIRLO samples had been available, then a test would have been conducted at 13% CP-ECR 

(severe embrittlement expected) and a confirmation test would have been conducted at 12% CP-ECR. It is 

possible that 200-wppm ZIRLO has a transition oxidation level of 12%, but a conservative approach was 

adopted in choosing 11% for which ductility was clearly established. 

In addition to the ductile-to-brittle oxidation levels established for 200-wppm and 240-wppm ZIRLO, 

Table 11 includes one data point that indicates embrittlement at 286±33 wppm and 9% CP-ECR. Thus, 

ZIRLO with about 290-wppm pre-test CHM has a transition CP-ECR <9%.
 

39 

Table 10. Oxidation conditions and results for pre-hydrided ZIRLO samples. Wg is the measured 

change in sample weight (in %) and wg is weight gain normalized to surface area. 

Oxidation

Test 

ID

Sample 

ID

Pre-test

Estimated

CHM, wppm

Test

Timea

, s

CP-ECR,

% 

Meas.

ECR, %

Wg, 

% 

wg, 

mg/cm3

ZLUPH#3 A15C3 231±8 166 11.0 11.8 4.23 7.56

ZLUPH#4 A15C1 245±23 145 10.0 10.4 3.73 6.69

ZLUPH#5 A13C3 256±46 166 11.0 11.5 4.11 7.41

ZLUPH#6 A13C1 226±16 166 11.0 11.5 4.12 7.40

ZLUPH#7 A13C5 263±36 110 8.0 8.3 2.95 5.34

ZLUPH#8 A15C5 314±91 127 9.0 9.5 3.47 6.13

a

From beginning of temperature ramp to the end of the heating phase. 

Table 11. Oxidation levels, RCT results, and CHM values for pre-hydrided ZIRLO samples. 

Oxidation

Test ID

(Sample ID)

CP-ECR,

% 

RCT

Sample 

ID

RCT

Sample 

L, mm

Permanent 

Strain,

% 

Offset

Strain,

% 

(CH)POX,

wppm

CHM, 

wppm

ZLUPH#3

(A15C3) 

11

B 

C

8.2 

8.1

– 

–

5.5 

1.8

195±3 

189±3

203±2 

197±1

ZLUPH#4

(A15C1) 

10

B 

C 

D

6.7 

7.7 

6.5

– 

– 

–

6.2 

7.6 

3.5

197±3 

194±9 

190±1

204±3 

201±9 

197±1

ZLUPH#5

(A13C3) 

11

B 

C

7.7 

8.0

0.4 

0.8

0.5 

0.9

232±23 

270±27

242±24 

281±28

ZLUPH#6

(A13C1) 

11

B 

C 

D

7.6 

7.5 

6.4

– 

– 

–

3.6 

5.6 

4.4

195±5 

176±7 

184±10

202±5 

183±7 

192±11

ZLUPH#7

(A13C5) 

8

B 

C 

D

8.6 

8.2 

7.1

– 

– 

21

20 

31 

22

235±49 

210±20 

211±12

242±50 

217±21 

217±12

ZLUPH#8

(A15C5) 

9

B 

C 

Da

7.6 

7.7 

5.1

– 

– 

–

1.6 

9.4 

3.1

276±22 

228±20 

209±7

286±23 

236±21 

216±7 a

Short end piece with oxidation on one face; excluded from data analysis. 
 

40 

Table 12. Analysis of PQD data to determine ductile-to-brittle transition CP-ECR and CHM values. 

CP-ECR,

% 

CHM, 

wppm

Offset

Strain,

% 

No. of

Data 

Points

Ductility 

Limit,

% Embrittlement CP-ECR and CHM

11

(10) 

9 

8 

240±33

(240) 

240±33 

240±33 

0.5

(>2.5) 

9.4 

20 

1

– 

1 

1 

2.6

2.5 

2.4 

2.3 

Based on data trend shown in Fig. 18 

and interpolation of offset-strain 

data, embrittlement limit is 10% CPECR for 240 wppm hydrogen 

(12)

11 

10 

8 

(200)

200±7 

200±7 

217±6 

(≈1.8)

3.8±1.6 

5.8±2.1 

27±6 

–

4 

3 

2 

2.7

2.6 

2.5 

2.4 

Based on data trend shown in Fig. 18 

and extrapolation of offset-strain 

data, embrittlement limit is 11% CPECR for 200 wppm hydrogen 
41

Figure 18. RCT ductility (i.e., offset strain) results for pre-hydrided ZIRLO as a function of CP-ECR oxidation 

9

6

3

0

33

30

27

24

21

18

15

12

7 8 9 10 11 12 

CP-ECR (%)

Offset Strain (%)

183 wppm

200 wppm

240 wppm

286 wppm

217 wppm

Ductility Limit
 

42 

Clearly more than six oxidation tests would have to be conducted for samples sectioned from prehydrided ZIRLO segments with axial gradients in CHM. Oxidation times and target CP-ECR values were

chosen based on the estimated levels from interpolation of CHM values determined for thin (about 2-mm) 

rings adjacent to the ends of oxidation samples. As shown in Table 13, there is enough variation between 

pre-test estimated and post-test-determined values of CHM for PQD samples that about double (12) the 

number of oxidation tests would have to be conducted to determine transition CP-ECR levels at two pretest CHM levels. Based on recent results from an industry-coordinated round robin [23], it appears that 

significant improvements in pre-hydriding and oxidation test methods would be needed for some of the 

participating laboratories to reproduce the data documented in this report. Nevertheless, the combined 

industry results are consistent with our results. 

Table 13. Comparison between pre-test estimated CHM values for oxidation samples and CHM

values for RCT samples sectioned from oxidation test samples and determined from 

measured post-oxidation (CH)POX values. 

Test ID

(Sample ID)

CP-ECR,

% 

RCT

Sample 

ID

Pre-test

Estimated

CHM, wppm

Determined from

Measured (CH)POX

CHM, wppm

ZLUPH#3

(A15C3) 

11

B 

C 

231±8

203±2 

197±1 

ZLUPH#4

(A15C1) 

10

B 

C 

D 

245±23

204±3 

201±9 

197±1 

ZLUPH#5

(A13C3) 

11

B 

C 

256±46

242±24 

281±28 

ZLUPH#6

(A13C1) 

11

B 

C 

D 

226±16

202±5 

183±7 

192±11 

ZLUPH#7

(A13C5) 

8

B 

C 

D 

263±36

242±50 

217±21 

217±12 

ZLUPH#8

(A15C5) 

9

B 

C 

D 

314±91

286±23 

236±21 

216±7 
 

43 

3.3 Embrittlement of Irradiated Cladding

3.3.1 Review of previous results

Results are presented in Ref. 1 for characterization, thermal and weight-gain benchmark tests, oxidation 

tests, post-oxidation characterization, RCTs, and post-RCT characterization for Zry-4, M5, and ZIRLO 

cladding materials from high-burnup fuel rods. For convenience, these materials are referred to as “highburnup” cladding in the following. After publication of Ref. 1 (2008), improvements have been made in 

the methodology used to determine the pre-oxidation hydrogen content in the cladding metal and in the 

offset-strain ductility criterion (see Eq. 8) used to determine the ductile-to-brittle transition CP-ECR. 

Thus, the results presented in Ref. 1 have been re-analyzed using these improvements.

The standard approach for determining hydrogen content in irradiated cladding is to section a thin ring, 

weigh it, heat it to melting, and measure the hydrogen evolved to give hydrogen content in wppm. For 

determination of oxidation embrittlement levels vs. CHM, it is necessary to know the hydrogen content in 

the pre-test or pre-LOCA-transient cladding metal. However, there are several sources of uncertainty in 

standard approaches used to determine cladding hydrogen content. The dominant uncertainty is due to the 

large axial variation of hydrogen content, especially for hydrogen contents >400 wppm (see Table 8 in 

Ref. 1). Even if an accurate CHM measurement could be made using short rings on both sides of a 25-mmlong pre-test oxidation sample, linear interpolation would not necessarily give an accurate result for the 

average hydrogen content of the test sample. Also, as the technique used to measure hydrogen is 

destructive, there is no way that this approach could be used to pre-determine the average CHM values for 

an oxidation test sample or an 8-mm-long RCT sample used to determine PQD. For PQD determination, 

the average hydrogen content in RCT rings needs to be determined as accurately as possible. The only 

way to do this is to rely on post-oxidation and post-RCT measurements and to use mass correction factors 

to reference the hydrogen to the pre-test mass of the cladding metal.

With regard to the pre-test material (corroded and defueled cladding), the mass of the hydrogen-analysis 

sample consists of the corrosion-layer mass (MC), the cladding metal mass (MM) and the fuel-cladding 

bond mass (MB). For higher-burnup cladding, the mass of the bond layer can be significant if the 

defueling process is poor. The mass of total hydrogen evolved from the sample consists of corrosion-layer

hydrogen (HC), cladding-metal hydrogen (HM) and cladding-bond hydrogen (HB). The bond layer can pick 

up hydrogen evolved from the nitric acid defueling process. 

As before, CH (in wppm) is the total hydrogen (HS) evolved from the sample divided by the total mass of 

the sample (MS). Equation 11 represents what is generally reported for the hydrogen content in corroded 

cladding: 

CH = (HS/MS) × 106 = [(HC + HM + HB)/( MC + MM + MB)] × 106 (11) 

Based on Eq. 11, the hydrogen content in the cladding metal, CHM = HM/MM, may be greater or less than 

what is measured for corroded, irradiated samples. 

Argonne has made pre-test hydrogen-content measurements like this on short rings adjacent to locations 

from which oxidation test samples were sectioned, along with metallographic examination of one or two 

rings within each cladding segment (M5 and ZIRLO segments were 80-mm long). Such measurements 

were used in planning oxidation times and corresponding CP-ECR levels, but they were found to be not 

accurate enough to determine the hydrogen content in the cladding metal for oxidation samples and for 

PQD rings sectioned from post-oxidation samples. 
 

44 

The approach adopted by Argonne to determine CHM for oxidation samples and PQD rings is similar to 

the one used in Section 3.2.2 for non-irradiated, pre-hydrided cladding samples except that an additional 

correction factor is required to account for the effects of the masses of the corrosion and fuel cladding 

bond layers. During the oxidation test, hydrogen that was in the corrosion and bond layers is driven off. 

Let MS be the pre-test mass of the cladding sample, as defined in Eq. 11. The post-oxidation hydrogen 

measurement is then given by 

(CH)POX = (HM/MS) (1 + Wg/100)-1 × 10-6 = (HM/MS) (fWg)

-1 × 10-6 (12) 

An additional correction needs to be made to separate the cladding metal mass from the masses of the 

corrosion and bond layers. This can be performed by measuring the total mass and length of the pre-test 

oxidation sample to determine the mass per unit length (MS/L) and by measuring the sample outer 

diameter (micrometer), the corrosion layer thickness (metallographic examination), the cladding metal 

wall thickness (metallographic examination), and the cladding bond-layer thickness (metallographic 

examination). From these measurements, as well as the densities for cladding metal and corrosion-bond 

oxide layers, the cladding metal mass per unit length (MM/L) can be determined. Let the metal-mass 

correction factor be defined as fMM = (MS/L)/(MM/L). By inverting Eq. 12 and applying the metal-mass 

correction factor, CHM may be determined from the measured (CH)POX by: 

CHM = (fMM × fWg) (CH)POX (13) 

Based on Argonne experience, the metal mass correction factor is in the range of 1.06 to 1.10 for highburnup ZIRLO with corrosion layer thicknesses of about 25 μm to 45 μm. The factors do not scale 

directly with the corrosion layer thickness because the mass of the fuel-cladding bond varies from 

defueled segment to defueled segment. For high-burnup M5 with a corrosion layer in the range of 10-15 

μm, the factor can vary from 1.003 to 1.06, where the low value is for a segment whose inner-surface was 

wire brushed. 

Two reasonable assumptions have been made in order to calculate CHM from post-oxidation 

measurements. The first assumption was that no significant hydrogen pickup occurred during the 

oxidation test. Previous oxidation tests with bare, as-fabricated cladding materials indicated ≤14-wppm 

hydrogen pickup relative to the pre-test mass of the samples [1]. It is quite likely that this hydrogen 

pickup occurred during the initial oxidation of the bare cladding (i.e., during the heating ramp) prior to the 

formation of a protective tetragonal oxide layer. This level of hydrogen pickup is insignificant for the 

high-burnup Zry-4 and ZIRLO tested because of the high hydrogen content of these samples, as well as 

the large circumferential and axial variations within these samples. For high-burnup M5 with only about 

100-wppm hydrogen, hydrogen pickup could represent 10 to 15% of the measured post-oxidation 

hydrogen content if hydrogen could easily diffuse through the corrosion and bond layers. If such 

hydrogen pickup did occur in high-burnup M5 cladding, it would be about the same as the axial and 

circumferential variations in hydrogen content. 

The second assumption used was that the density of the cladding metal was not reduced by the presence 

of hydrides. This is a very good assumption for high-burnup M5 with low hydrogen content and a 

reasonable approximation for high-burnup Zry-4 and ZIRLO. 

Previous results for high-burnup (HBU) Zry-4 cladding

Testing of 15×15 Zry-4 from HBU (63 GWd/t) HBR fuel rods was conducted in 2005. These samples 

were not thoroughly defueled. As the mass of the fuel adherent to the fuel-cladding bond was not known 
 

45 

and likely varied from sample to sample, pre-test hydrogen measurements at axial locations between 

oxidation samples, along with post-test hydrogen measurements, could not be used to determine reliable 

values for CHM. An additional segment was sectioned from an axial location adjacent to the oxidation 

samples, carefully defueled, and sectioned into 8 rings for hydrogen determination. The hydrogen content 

(550±100 wppm) was relatively constant from ring to ring. Also, the total mass of the rings was 1.52 g. 

Thus, 550±100 wppm is a reliable characterization of the pre-test hydrogen content in the corroded 

cladding. The corrosion layer was ≈70-μm thick and the cladding metal wall was 710-μm thick. 

Including the bond layer, the metal-mass correction factor was calculated to be 1.09. However, the 

corrosion layer had about 10% porosity and may have contained a significant amount of hydrogen. As no 

reliable post-oxidation hydrogen data are available for the Zry-4 LOCA samples, it is assumed that the 

mass-correction factor and the corrosion-layer-hydrogen correction factor balance out to give 550±100 

wppm of hydrogen in the cladding metal.

High-burnup Zry-4 samples with 550±100 wppm CHM were oxidized to CP-ECR levels of 2.9% to 9.4% 

(includes oxidation during cooling). Samples were slow-cooled without quench. RCT results are 

summarized in Table 14. In selecting the transition CP-ECR level, the embrittling effects of quench on

ductility were taken into account. Rings tested at 6.6% and 7.6% are expected to be brittle following 

quench based on results for pre-hydrided Zry-4 with comparable hydrogen contents. The 4.5% CP-ECR 

ring had enough ductility such that it should be ductile after quench. The transition CP-ECR is estimated 

to be 5% for the high-burnup Zry-4 tested. The peak oxidation temperature was ≈1180°C at 5% CP-ECR.

Table 14. RCT results for HBU Zry-4 samples oxidized at ≤1200°C and cooled without quench. 

CHM was ≈550±100 wppm for these samples. 

Oxidation

Test IDa

CP-ECR,

% 

Maximum

Oxidation

T, °C

Offset 

Strain, %

Offset-Strain 

Ductility Limit, 

% 

Permanent 

Strain, %

HBRI#2 2.9 1110 >45 1.7 >43

HBRI#1 4.5 1169 >12

≈37

1.9 –

– 

HBRI#3 6.6 1196 4.0 2.1 2.6

HBRI#5 7.6 1197 4.0 2.2 2.9

HBRI#4 9.4 1198 0.5 2.4 0.6

a

HBRI#6 sample (740 wppm) is not included because it was not used to determine the transition CP-ECR.

Previous results for high-burnup M5 cladding

M5 cladding samples from a fuel rod irradiated to high burnup (63 GWd/t) in one of the Ringhals 

Reactors were used for oxidation and PQD tests. Tests MI#1 to MI#4 were slow-cooled without quench 

to determine an upper-bound transition CP-ECR for samples cooled with quench. Tests MI#5 to MI#7 

were cooled with quench at 800°C and were used to determine PQD. The corrosion layer thickness was 

12 μm for one of two adjacent segments used to section oxidation test samples. Measured values for preand post-test hydrogen content for these test samples are listed in Table 15. Results for CHM were higher 

than expected for six of the seven oxidation samples based on data given in Ref. 24.
 

46 

Table 15. Determination of CHM based on post-oxidation hydrogen measurements [(CH)POX] for 

high-burnup M5 samples used for oxidation tests; fWg is post-test-weight/pre-test-weight 

and fMM is pre-test sample-mass/metal-mass. Sample masses are listed in parentheses.

Oxidation

Test #

(ID#)

CP-ECR, 

% 

(CH)POX, 

wppm

(mass, g) fWg

Meas.

Pre-test

mass/L,

g/cm fMM

CHM, 

wppm

Measured

Pre-test

CH, wppm

(mass, g) 

MI#1

(645A1) 

13.4 85±7 

(0.35) 

1.040 1.036 1.028 91±8 101±21 

(0.1 g) 

MI#2

(645A4) 

15.1 107±14 

(0.38) 

1.040 1.042 1.035 115±15 101±21 

(0.1 g) 

MI#3

(645A3) 

17.0 128±19 

(0.24) 

≈1.045a ≈1.04 ≈1.04 ≈140±20 101±21 

(0.1 g) 

MI#4

(645B6) 

18.1 109±9 

(0.26) 

1.052 1.066 1.059 122±10 108±5 

(0.22 g) 

MI#5

(645B2) 

16.1 130±11 

(0.33) 

1.047 1.039 1.033 140±12 108±5 

(0.22 g) 

MI#6

(645C3) 

15.1 138±15 

(0.25 g) 

1.047 1.044 1.025 148±14 108±10 

(0.40 g) 

MI#7

(645C6) 

13.4 130±15 

(1.31 g) 

1.039 1.007 1.003 135±16 108±10 

(0.40 g) 

a

Pre- and post-test masses were not measured. The fWg value was estimated by means of interpolation.

Post-oxidation and post-quench ductility results for high-burnup M5 are summarized in Table 16. Results 

for samples cooled without quench are interesting in that they show little effect of hydrogen content on 

the transition CP-ECR. Based on permanent strain, the transition CP-ECR is ≈18% for ≈120-wppm CHM, 

as compared to 20% CP-ECR for as-fabricated M5 with only 5-wppm CHM. However, for quenched 

samples the hydrogen content caused a significant reduction in the transition CP-ECR. On the basis of 

offset strains, quenched RCT samples were ductile (3 data points) at 13.4% CP-ECR and brittle (3 data 

points) at 15.1% CP-ECR. Using interpolation, the transition CP-ECR for HBU M5 is estimated to be 

14.5% for 140-wppm CHM. In order to be conservative in data interpretation and consistent with the 

embrittlement observed at 15.1% CP-ECR, 14.5% is rounded down to 14% for the transition CP-ECR. 

Consistent with the definition used for transition CP-ECR, embrittlement of M5 samples with ≥140 wppm 

CHM is expected to occur for oxidation levels >14% CP-ECR. 
 

47 

Table 16. RCT results for HBU M5 samples oxidized at 1200°C and cooled without and with 

quench (see Table 74 in Ref. 1). 

Oxidation

Test ID

CP-ECR,

% 

CHM, 

wppm

Average

Offset 

Strain, %

Offset-Strain 

Ductility Limit, 

% 

Permanent 

Strain, %

MI#1

Slow Cooled

13.4 91±8 12.0±0.2 2.9 6.5±0.3

MI#2

Slow Cooled

15.1 115±15 8.5±1.4 3.0 4.4

MI#3

Slow Cooled

17.0 140±20 5.8±0.6 3.2 1.3

MI#4

Slow Cooled

18.1 122±10 4.8±1.0 3.4 1.1

MI#5

800°C Quench

16.1 140±20 2.9±0.9 3.2 –

MI#6

800°C Quench

15.1 148±14 2.8±0.7 3.0 –

MI#7

800°C Quench

13.4 135±16 3.4±1.3 2.9 –

Previous results for high-burnup ZIRLO cladding

ZIRLO cladding samples from a fuel rod irradiated to high-burnup (70 GWd/t) in the North Anna 

Reactors were used for oxidation and PQD tests. Tests ZLI#1 to ZLI#3 were slow-cooled without quench 

to determine an upper bound on the transition CP-ECR. Tests ZLI#4 to ZLI#6 were cooled with quench at 

800°C and were used to determine PQD. The corrosion layer thickness was measured to be 42±3 μm for 

one of two adjacent segments from which oxidation samples were sectioned [1]. Measured values for preand post-test hydrogen content for rings from the three oxidation tests with quench are listed in Table 17. 

The average CHM values were about 50 wppm less than the hydrogen content measured for the corroded 

pre-test sample. This is less than the measured circumferential and axial variations. 

Table 17. Determination of CHM based on measured post-oxidation hydrogen content for highburnup ZIRLO samples used for oxidation tests and PQD determination; fWg = post-testweight/pre-test-weight, fMM = pre-test sample-mass/metal-mass. Sample masses are given 

in parentheses.

Oxidation

Test #

(ID#)

CP-ECR, 

%

(CH)POX, 

wppm

(mass, g) fWg

Meas.

Pre-test

mass/L,

g/cm fMM

CHM, 

wppm

Measured

Pre-test

CH, wppm

(mass, g)

ZLI#4

(648F2)

6.3 505±107

(0.26)

0.996 1.074 1.074 540±114 620±140

(0.55)

ZLI#5

(648F6)

4.0 505±107

(0.27)

0.990 1.074 1.096 593±115 600±150

(0.36)

ZLI#6

(648E2)

5.1 414±115

(0.79)

1.004 1.073 1.080 451±101 500±120

(0.24)
 

48 

Post-oxidation results for slow-cooled samples are presented in Ref. 1 (see Table 72). Although pre- and 

post-test sample weights were not measured for ZLI#1 through ZLI#3 test samples, pre- and post-test 

hydrogen contents were measured. As weight gain is expected to be negligible for these samples (see fWg

in Table 17 for quenched samples), post-test hydrogen measurements are considered to be more reliable 

than pre-test measurements, especially as the masses of the post-test samples were generally larger than 

those for the pre-test samples. Offset strains indicated high ductility (19±10%) for the 540±100 wppm 

sample oxidized to 6.3% (1176°C peak oxidation temperature), moderate ductility (4.5±0.5%) for the 

510±100 wppm sample oxidized to 8.1% (1191°C peak oxidation temperature), no ductility (1% offset 

strain and 0.3% permanent strain) for the 490±120 wppm sample oxidized to 10.2% (1200°C peak 

oxidation temperature). Given the expected decrease in ductility due to quench at 800°C, oxidation tests

with quench (ZLI#4 through ZLI#6) were initiated at 6.3% CP-ECR. RCT results for these oxidized and 

quenched samples are summarized in Table 18. 

Table 18. RCT results for high-burnup ZIRLO samples oxidized at <1200°C and cooled with 

quench at 800°C. 

Oxidation

Test ID

(Sample ID)

CP-ECR,

%

Maximum

Oxidation

T, °C

CHM, 

wppm

Average 

Offset 

Strain, %

Offset-Strain 

Ductility 

Limit, 

%

Permanent 

Strain, %

ZLI#4

(648F2)

6.3 1176 540±114 0.8±0.4

(2 data points)

2.1 0.2

(1 data point)

ZLI#5

(648F6)

4.0 1132 593±115 >43

(2 data points)

1.9 >44

(2 data points)

ZLI#6

(648E2)

5.1 1162 451±101 40±27

(3 data points)

2.0 –

The PQD results for HBU ZIRLO with high-hydrogen content were more difficult to interpret than the 

PQD results for pre-hydrided ZIRLO with low-hydrogen content because: (a) only three oxidation with 

quench tests were conducted; (b) the average CHM was not constant (about 90 wppm variation); (c) the 

circumferential variation in CHM was large (about 100 wppm); and (d) the peak oxidation temperature 

decreased significantly as the target CP-ECR decreased. However, experience gained in testing prehydrided and high-burnup Zry-4 was used to assess and interpret the data in Table 18. The PQD sample 

with 540±114 wppm CHM was very brittle at 6.3% CP-ECR based on both offset and permanent strains, 

while the PQD sample with higher CHM (593±115 wppm) was highly ductile at 4.0% CP-ECR. It is 

reasonable to infer that the transition CP-ECR is about 5% for high-burnup ZIRLO with 540-wppm CHM

and a peak oxidation temperature of 1160°C. Conversely, the higher CHM sample (593 wppm) would be 

very brittle at 6% CP-ECR and have lower ductility at 5% CP-ECR than the 540 wppm sample. 

Engineering judgment was used to assess the transition CP-ECR (to the nearest 1%) to be 4% for highburnup ZIRLO with about 590 wppm CHM and a peak oxidation temperature of about 1130°C. 

Stronger conclusions could have been reached if all three LOCA samples had the same average CHM (e.g., 

540 wppm). However, cladding with high-hydrogen content has large axial and circumferential variations 

in CHM over relatively short distances along fuel rods. ZLI#4 and ZLI#5 test samples were sectioned from 

the same 80-mm-long cladding segment with a mid-span to mid-span distance between of only 17 mm. 

The ZLI#6 LOCA test sample was from an adjacent 80-mm-long segment. Pre-test hydrogen-content 

measurements for the corroded cladding were 464±71 wppm and 563±145 wppm for 2-mm-long rings 

spaced 52 mm apart (see Table 68 [1]). These variations are very significant and should be accounted for 

in correlations for hydrogen pickup vs. corrosion layer thickness. 
 

49 

3.3.2 Current PQD results for high-burnup ZIRLO

Characterization

Figure 161 of Ref. 1 shows axial locations of ZIRLO Rod AM2-L17 irradiated to HBU (70 GWd/t) in the 

North Anna Reactors. Studsvik sectioned and defueled 80-mm-long cladding segments from this rod for 

ANL. Segments 5, 6, and 8 with high hydrogen concentrations were tested earlier and results are 

summarized in Section 3.3.1. Segments 1 and 2 with lower corrosion-layer thickness (25–30 μm based on 

eddy current) and expected medium hydrogen concentrations were selected for oxidation and RCT 

testing. Corresponding Argonne ID numbers for these segments are 648A and 648B, respectively. 

Figure 19 shows micrographs at one of eight circumferential locations for: (a) cladding cross section; (b) 

corrosion layer; and (c) fuel-cladding bond layer. The measured thicknesses were 26±1 μm for the 

corrosion layer (δc), 551±5 μm for the metal wall (hm), and ≈7±1 μm for the fuel-cladding bond (δb). 

However, as shown in Fig. 19c, local regions of the bond layer were about 20-25 μm thick. These thicker 

bond regions are consistent with the higher (factor of 2) gamma dose-rate readings for 648A and 648B, as 

compared to the dose rate reading for segment 648F (Segment 6) used in previous oxidation tests. 

Figure 20 shows the hydride morphology at two circumferential locations adjacent to the sample with 

318±30 wppm of hydrogen in the corroded cladding. As is typical of seven of the eight locations imaged, 

hydrides were primarily circumferential and clustered in a dense rim near the corrosion layer (Fig. 20a). 

Figure 20b shows one location that had long radial hydrides within the inner third of the cladding wall. 

For LOCA events, the radial variation in hydrogen content and hydride morphology are not important 

because hydrogen homogenizes quickly across the beta phase of the cladding. 

Table 19 summarizes characterization results for δc, hm, δb, and CH. For the LECO hydrogen-content 

measurements (CH), short cladding rings (≈2-mm-long) were sectioned and then snipped into four quarter 

rings. Hydrogen content and mass were measured for each quarter ring. The hydrogen content included 

the hydrogen evolved from the cladding metal and the corrosion layer, as well as the bond layer. The 

sample mass included masses of the corrosion layer, the cladding metal, and the fuel-cladding bond. The 

four measurements per ring were also used to determine the one-sigma (1-σ) circumferential variation in 

hydrogen content. The weight-averaged hydrogen contents of the two segments were 310±32 wppm for 

648A and 366±30 wppm for 648B. Two oxidation samples were sectioned from 648A, and three

oxidation samples were sectioned from 648B. 

Pre-oxidation hydrogen data were used for choosing target oxidation levels (CP-ECR values). However, it

was also realized that the hydrogen content in the cladding metal could be higher or lower than what was 

measured. LOCA samples were weighed before and after oxidation tests. Post-oxidation hydrogen 

contents were corrected for weight gain (1-2%). A final correction was made to account for the ratio of 

pre-test sample mass (measured) and metal mass (calculated from metallographic results). This factor 

(1.06 to 1.10) varied from sample to sample and reflected the added weight of the pre-test sample due to 

the fuel-cladding bond, which included fission products and actinides in or adherent to the fuel-cladding 

bond. The methodology for determining the metal hydrogen content (CHM) is described in detail in 

Section 3.3.1. As indicated in Table 19, CHM values were close to measured CH values. Also, axial and 

circumferential variations in both CHM and CH were much smaller for these lower hydrogen-content 

samples than what was observed previously for high hydrogen-content samples. The results suggest that 

for 25–30 μm corrosion-layer thicknesses the additional hydrogen in the corrosion layer is compensated 

by the additional mass of the corrosion layer such that CHM ≈ CH. 
50 

(a) Cladding cross section

(b) Corrosion layer

(c) Fuel-cladding bond layer

Figure 19. Micrographs taken from a sample sectioned near the middle of ZIRLO segment 

648A: (a) cladding cross section; (b) corrosion layer; and (c) fuel-cladding bond

layer.
51 

(a) Circumferential hydrides

(b) Circumferential and radial hydrides

Figure 20. Hydride morphology for high-burnup ZIRLO sample 648A5: (a) circumferential 

hydrides (typical); and (b) circumferential and radial hydrides. Based on measured 

values for an adjacent ring (648A4), the CH was 318±30 wppm. 
52 

Table 19. Characterization of high-burnup ZIRLO segments used for current oxidation tests. 

ANL

Segment

ID

δc from

Eddy Cur., 

μm

δc from

Metallography, 

μm

hm, 

μm

δb

μm

Measured

Pre-Test

CH, wppm

Estimated

Pre-Test

CHM, wppm

648A 25-28 

– 

26±1 

– 

– 

551±5 

– 

– 

7±1a

– 

294±29 

318±30 

317±38 

296±33 

648B 25-30 

– 

– 

– 

– 

– 

– 

– 

– 

– 

361±36 

367±29 

370±33 

340±34 

a

Local regions of bond layer were as thick as 20-25 μm. 

Thermal and metallurgical benchmark tests for high-burnup ZIRLO 

For in-cell tests with high-burnup ZIRLO, thermal benchmark tests were first conducted in the out-of-cell 

LOCA apparatus. Two thermocouples (TCs) were welded onto as-fabricated ZIRLO samples, in addition 

to the three TCs permanently welded to the Inconel holder just above the sample. Tests were repeated 

with fresh samples until furnace control parameters were established to give the desired temperature ramp 

and hold temperature. The results of benchmark test ZLU#145 with a bare ZIRLO sample are shown in 

Fig. 21. The time for the sample to reach 1200oC was about 100 s. At 150 s, the average of the two 

sample TCs was 1202±7oC. The heating phase of the test was 200 s. The corresponding CP weight gain 

and CP-ECR, including the cooling phase, were 8.27 mg/cm2 and 12.7%, respectively. The CP-calculated 

oxide layer grown on the inner and outer surfaces of the cladding was about 40-μm thick. 

For high-burnup ZIRLO cladding, the corrosion and bond oxide layers will result in a slower temperature 

ramp rate as compared to the rate for bare ZIRLO. At the completion of benchmark test ZLU#145, the 

oxidized sample was again ramped to the target temperature using the same control parameters as were 

used for ZLU#145. To distinguish these two tests, the one with the pre-oxidized sample was called 

ZLU#145A. Figure 22 shows the comparison between the average cladding temperatures for bare and 

pre-oxidized ZIRLO. The small temperature pre-peak was reduced from ≈1110oC to 1060oC. The 

average ramp rates (≈2.2°C/s) from 1080oC to 1180 oC were comparable. However, the time to reach 

1200oC was longer for the pre-oxidized cladding (150 s vs. 100 s for bare cladding). The average 

temperature of the pre-oxidized cladding was 1200±3oC at 150 s during the heating phase. It was later 

verified that there was essentially no difference in the thermal response of pre-oxidized cladding samples 

with thin (≈12 μm) and thick (≈40 μm) oxide layers. Thus, the ZLU#145A thermal history was 

determined to be appropriate one for high-burnup cladding oxidation tests.

The test train benchmarked in the out-of-cell LOCA apparatus was installed in the in-cell apparatus 

located in Cell #4 of the IML with the same three holder TCs welded onto the sample holder. A weightgain benchmark test was conducted in-cell with bare ZIRLO. The same temperature history and heatingphase shown in Fig. 21 were used to conduct this weight-gain benchmark test (ZLU#146). The measured 

weight gain (7.91 mg/cm2

) was within 5% of the CP-calculated weight gain (8.27 mg/cm2). These results 

validated the use of the ZLU#145A thermal history for in-cell tests with high-burnup ZIRLO samples.
53

Figure 21. Thermal history for bare ZIRLO cladding determined from test ZLU#145. At 150 s into the heating 

sample temperature was 1202±7oC.

900

800

700

600

500

400

300

1300

1200

1100

1000

500 550 600 650 700 750 

Time (s)

Temperature (°C)

Control TC1 on holder

TC2 on holder

TC3 on holder

TC4 on sample

TC5 on Sample
54

Figure 22. Thermal history for bare ZIRLO (ZLU#145) and pre-oxidized (ZLU#145A) cladding. At 150 s into t

average temperature for the pre-oxidized sample was 1200±3oC. 

900

800

700

600

500

400

300

1300

1200

1100

1000

500 550 600 650 700 750 

Time (s)

Temperature (oC)

ZLU#145: bare ZIRLO

ZLU#145A: pre-oxidized (40 μm) ZIRLO
55 

Ductility Results for High-burnup ZIRLO

Four oxidation tests and 9 ring-compression ductility tests were conducted with samples sectioned from 

high-burnup ZIRLO segments 648A and 648B. Test conditions and PQD data, along with ductility data 

for one LOCA test cooled without quench, are summarized in Table 20. 

For planning purposes, the measured cladding wall thickness (0.55 mm) was used to derive the 

relationship between CP-ECR (in %) and wg (in mg/cm2

):

CP-ECR = 1.594 wg (14) 

Test ZLI#7 was conducted without quench with a sample containing about 340 wppm CHM to ensure 

adequate ductility at 10% CP-ECR. The heating-phase test time (153 s) and corresponding CP-ECR 

(10%) are shown in Fig. 23. The pre-test sample outer diameter was 9.50 mm. Subtracting twice the 

corrosion layer thickness measured at a neighboring location gave a pre-test cladding-metal outer 

diameter of about 9.45 mm. This pre-test metal diameter was used to normalize offset and permanent 

displacements to determine offset and permanent strains. As shown in Table 20, the offset and permanent 

strains at 135oC were in the range of 10-13% and 5-9%, respectively. These results indicate significant 

ductility for the slow-cooled LOCA sample following oxidation to 10% CP-ECR. The results were used 

to plan the first test with quench. 

For test ZLI#8, a lower hydrogen content (≈300 wppm) sample from segment 648A was used. The test 

time and CP-ECR were kept the same as for test ZLI#7. However, the sample was quenched at 800oC. 

The effects of quench were significant in that the rings sectioned from the LOCA sample were brittle 

even though the sample had less hydrogen in the metal than the ZLI#7 sample. The offset strains (0.7% 

and 2.2%) were less than the ductility criterion (2.5%) at 10% CP-ECR. Permanent displacements and 

strains could not be determined because the rings cracked into two pieces. 

In order to bracket the ductile-to-brittle transition oxidation level, test ZLI#9, also with ≈300 wppm 

hydrogen, was conducted for 118-s heating-phase time, which corresponded to 8% CP-ECR. As shown in 

Fig. 23, the peak oxidation temperature at the end of the heating phase was 1190oC. The first two rings 

compressed at 135oC also cracked at top and bottom locations with adequate offset strains of 2.9% and 

4.6% indicative of ductile behavior (≥2.3%). In order to confirm ductile behavior, the load was terminated

for the third ring before cracking occurred and after sufficient offset strain (≥2%). The offset and 

permanent displacements of the uncracked ring were 0.19 mm and 0.11 mm, respectively. These 

correspond to 2.0% offset and 1.2% permanent strains. The difference (0.8%) is less than the difference 

(1.3%) used for the offset-strain ductility criterion at 8% CP-ECR. These results are encouraging because 

the offset-strain criterion is based on the one-sigma upper bound for ductility. 

Results in Table 20 indicate that the ductile-to-brittle transition oxidation level was 9% CP-ECR for 300-

wppm CHM. Following this determination, the remaining two oxidation samples with 340-wppm CHM

were used for oxidation tests. During the ZLI#10 ramp, the control thermocouple failed and the test was 

terminated after the other two thermocouples on the holder recorded temperatures above 1250oC. The 

remaining sample was oxidized to 8% CP-ECR in test ZLI#11 and quenched at 800oC. As indicated in 

Table 20, the two rings sectioned from this oxidation sample had very low offset strains (0.6–1.0%) 

indicating embrittlement. Although the ductile-to-brittle transition CP-ECR could not be determined 

precisely, the data point is useful in establishing that the transition CP-ECR is <8% for 340-wppm CHM.
56

Figure 23. Thermal history and test times for high-burnup ZIRLO tests conducted to 8% (ZLI#9 & 11) and 10

900

800

700

600

500

400

300

ECR.

1300

1200

1100

1000

500 550 600 650 700 750 

Time (s)

Temperature (oC)

10% CP-ECR

8% CP-ECR
57 

Table 20. Post-test ductility results for high-burnup ZIRLO cladding oxidized at ≤1200oC, cooled 

with or without (SC) quench at 800oC, and ring-compressed at 135oC. 

Oxidation

Test ID

CHM, wppm

(mass, g)

Test

Time,a

s 

Peak

T,

oC

CP-ECR,

% 

Offset

Strain,

% 

Offset Strain

Ductility

Criterion, % 

Permanent

Strain,

% 

ZLI#7

SC

340±34

(0.28)

153 1200 10 9.9

13.4

2.5 5.0

9.4

ZLI#8

Quench

294±37

(0.76)

153 1200 10 0.7

2.2

2.5 ---

---

ZLI#9

Quench 

296±35

(0.72) 

118 1190 8 2.9

4.6 

>2.0

2.3 ---

---

>1.2

ZLI#11

Quench

340±34

(0.28)

118 1190 8 0.6

1.0

2.3 ---

--- a

From beginning of ramp at 300oC to end of heating phase. 

3.4 Discussion of PQD Embrittlement Results

Figure 24 shows the oxidation embrittlement threshold vs. hydrogen content for as-fabricated cladding 

alloys, pre-hydrided Zry-4 and ZIRLO, and high-burnup Zry-4, ZIRLO, and M5. The bilinear fit to the 

data in Fig. 24 for high-burnup cladding alloys is more of a guide than a best fit. For CHM ≤400 wppm, it 

is based on one high-burnup M5 data point (140 wppm and 14% CP-ECR) and two high-burnup ZIRLO 

data points (300 wppm and 9% CP-ECR; 340 wppm and <8% CP-ECR). For CHM ≤400 wppm: 

CP-ECR = 18.5 – 0.03125 CHM (15) 

where CHM is in wppm and CP-ECR is in %. Equation 15 gives 6% CP-ECR for 400 wppm. The choice of 

400 wppm for the slope change is somewhat arbitrary as there are no data for high-burnup cladding 

between 340 wppm (1190°C peak temperature for high-burnup ZIRLO) and 540 wppm (1160°C peak 

temperature for high-burnup ZIRLO). As CHM increases beyond 340 wppm, the peak oxidation 

temperature decreases to as low as 1130°C for high-burnup ZIRLO at 4% CP-ECR. For CHM >400 wppm: 

CP-ECR = 6.0 – 0.00865 (CHM – 400 wppm) (16) 

For as-fabricated cladding, the embrittlement oxidation level was found to be independent of quench 

temperature and of cladding alloy for modern belt- or wheel-polished materials. We speculate that the 

embrittlement threshold for as-fabricated cladding is also relatively insensitive to higher heating rates

(>3°C/s from 1080°C to 1180°C) and slower cooling rates (<10°C/s) to quench. For pre-hydrided and 

high-burnup cladding alloys, the embrittlement threshold exhibited a high sensitivity to heating rate, peak 

oxidation temperature, cooling rate prior to quench, and quench temperature. 

The results in Fig. 24 suggest that pre-hydrided cladding is a good surrogate for high-burnup cladding in 

terms of PQD. A second fit was performed by including all data points and using a decaying exponential 

function. In one curve fitting attempt, high-burnup data points were assigned a weight of 4 while all other 

data points were assigned a weight of 1. In a second curve fitting attempt, all data points were assigned a 

weighting factor of 1. The difference in curve-fitting parameters was negligible for these two approaches. 

The best-fit curve and data are shown in Fig. 25. The best-fit exponential function is given in Eq. 17. 
58

Figure 24. Embrittlement oxidation level (CP-ECR) as a function of pre-transient hydrogen content in cladding 

oxidation temperatures of 1190–1200°C for <400 wppm hydrogen content and 1130–1190°C for >40

content. 

8

6

4

2

0

20

18

16

14

12

10

0 100 200 300 400 500 600 700 

Embrittlement Oxidation Level (%)

Hydrogen Content (wppm)

 

15x15 Vintage Zry-4

15x15 Low-Sn Zry-4

17x17 Low-Sn Zry-4

10x10 Zr-lined Zry-2

17x17 ZIRLO

17x17 M5

 

 

63 GWd/t

70 GWd/t

70 GWd/t

64 GWd/t

Ductile

Brittle

Previous 10CFR50.46 Limit
59

Figure 25. Exponential fit to data for embrittlement CP-ECR as a function of pre-transient hydrogen content in 

8

6

4

2

0

20

18

16

14

12

10

0 100 200 300 400 500 600 700 

Embrittlement Oxidation Level (%)

Hydrogen Content (wppm)

 

15x15 Vintage Zry-4

15x15 Low-Sn Zry-4

17x17 Low-Sn Zry-4

10x10 Zr-lined Zry-2

17x17 ZIRLO

17x17 M5

 

 

63 GWd/t

70 GWd/t

70 GWd/t

64 GWd/t

Ductile

Brittle

Previous 10CFR50.46 Limit
60 

CP-ECR = 19.38 exp (-0.0026 CHM) (17) 

For the best-estimate curve using the weighted average of data points, in which high-burnup data were 

given a weighting factor of 4, the pre-exponential factor was 19.42 and the exponent factor was still 

-0.0026. The exponential fit (Eq. 17) gives better agreement with the data than the bilinear fit (Eqs. 15–

16) and it is easier to use the continuous exponential equation. Also, the exponential fit is based on more 

data points by using the assumption that non-irradiated, pre-hydrided samples are a good surrogate for 

high-burnup samples for determining ductility and embrittlement threshold. The data generated in this 

program support the position that embrittlement is governed by hydrogen content in the cladding metal 

and oxidation level independent of whether the hydrogen pickup occurred during reactor operation or 

during the pre-hydriding process. 

Equation 17 ignores the differences in peak cladding temperature for LOCA samples with high hydrogen 

content. Thus, it is not more mechanistic than the bilinear representation of the data. The 4% to 6% CPECR values attained for the double-sided oxidized samples with about 2°C/s temperature increase from

1000°C to the peak oxidation temperature <1200°C could have been obtained at 1200°C peak oxidation 

temperature by using faster heating rates, one-sided oxidation, and direct quench from 1200°C. CEA ran a 

number of tests with as-fabricated and pre-hydrided samples using these oxidation test parameters [9-11]. 

The hold time at 1200°C was as low as 50 s, which resulted in about 3% CP-ECR. Pre-hydrided (600 

wppm) Zry-4 was found to be brittle after the CEA oxidation tests. However, the CEA test conditions 

were less LOCA-relevant than the Argonne test conditions. 

For typical fuel management schemes, assemblies containing high-burnup fuel rods are placed closer to 

the edges of the core than the center of the core. It is unlikely that these high-burnup fuel assemblies 

would reach 1200°C during a LOCA. Thus, the decaying exponential curve (Eq. 17) is very useful for 

high-burnup fuel as long as the peak cladding temperature does not exceed 1130°C. 

In order to allow independent assessment of the PQD data for high-burnup ZIRLO cladding, the RCT

load-displacement curves are presented in Figs. 26–34. 
61

Figure 26. Load-displacement curve for HBU ZIRLO cladding ring 648B2B oxidized to 10% CP-ECR and cool

0.45

0.40

0.35

0.30

0.25

0.20

0.15

0.10

0.05

0.00

0.0 0.5 1.0 1.5 2.0 

Displacement (mm)

Load (kN)

0.94 mm

ZLI#7 Ring 648B2B

340±34 wppm H

10% CP-ECR, no Quench

9.9% Offset Strain

5.0% Permanent Strain
62

Figure 27. Load-displacement curve for high-burnup ZIRLO cladding ring 648B2D oxidized to 10% CP-ECR a

quench.

0.50

0.45

0.40

0.35

0.30

0.25

0.20

0.15

0.10

0.05

0.00

0.0 0.5 1.0 1.5 2.0 

Displacement (mm)

Load (kN)

1.29 mm

ZLI#7 Ring 648B2D

340±34 wppm H

10% CP-ECR, no Quench

13.7% Offset Strain

9.4% Permanent Strain
63

Figure 28. Load-displacement curve for high-burnup ZIRLO cladding ring 648A3B oxidized to 10% CP-ECR a

800oC.

0.30

0.25

0.20

0.15

0.10

0.05

0.00

0.0 0.2 0.4 0.6 0.8 1.0 

Displacement (mm)

Load (kN)

0.07 mm 

ZLI#8 Ring 648A3B

300±33 wppm H

10% CP-ECR, Quench

0.7% Offset Strain
64

Figure 29. Load-displacement curve for high-burnup ZIRLO cladding ring 648A3C oxidized to 10% CP-ECR 

800oC.

0.40

0.35

0.30

0.25

0.20

0.15

0.10

0.05

0.00

0.0 0.2 0.4 0.6 0.8 1.0 

Displacement (mm)

Load (kN)

0.21 mm

ZLI#8 Ring 648A3C

300±33 wppm H

10% CP-ECR, Quench

2.2% Offset Strain
65

Figure 30. Load-displacement curve for high-burnup ZIRLO cladding ring 648A6B oxidized to 8% CP-ECR an

0.50

0.45

0.40

0.35

0.30

0.25

0.20

0.15

0.10

0.05

0.00

0.0 0.2 0.4 0.6 0.8 1.0 1

Displacement (mm)

Load (kN)

0.27 mm

ZLI#9 Ring 648A6B

300±33 wppm H

8% CP-ECR, Quench

2.9% Offset Strain
66

Figure 31. Load-displacement curve for high-burnup ZIRLO cladding ring 648A6C oxidized to 8% CP-ECR a

0.60

0.50

0.40

0.30

0.20

0.10

0.00

0.0 0.2 0.4 0.6 0.8 1.0 1.2 

Displacement (mm)

) Load (kN

0.43 mm

ZLI#9 Ring 648A6C

300±33 wppm H

8% CP-ECR, Quench

4.6% Offset Strain
67

Figure 32. Load-displacement curve for high-burnup ZIRLO cladding ring 648A6D oxidized to 8% CP-ECR a

0.35

0.30

0.25

0.20

0.15

0.10

0.05

0.00

0.0 0.2 0.4 0.6 0.8 1.0 1

Displacement (mm)

Load (kN)

0.19 mm

ZLI#9 Ring 648A6D

300±33 wppm H

8% CP-ECR, Quench

2.0% Offset Strain

1.2% Permanent Strain
68

Figure 33. Load-displacement curve for high-burnup ZIRLO cladding ring 648B6B oxidized to 8% CP-ECR an

0.30

0.25

0.20

0.15

0.10

0.05

0.00

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 

Displacement (mm)

Load (kN)

0.06 mm

ZLI#11 Ring 648B6B

340±34 wppm H

8% CP-ECR, Quench

0.6% Offset Strain
69

Figure 34. Load-displacement curve for HBU ZIRLO cladding ring 648B6C oxidized to 8% CP-ECR and quen

0.30

0.25

0.20

0.15

0.10

0.05

0.00

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 

Displacement (mm)

Load (kN)

0.09 mm

ZLI#11 Ring 648B6C

340±34 wppm H

8% CP-ECR, Quench

1.0% Offset Strain

0.5% Permanent Strain

71

4 BREAKAWAY OXIDATION

Breakaway oxidation occurs in all zirconium-alloy cladding materials if the cladding is held for long 

times (minutes to hours) at constant temperatures between 650 and 1100°C. Under these conditions, the 

crystal structure of the zirconium dioxide can change to a form that is not protective, and hydrogen from 

the metal-steam reaction during LOCA oxidation can then be absorbed rapidly into the cladding metal 

thus affecting cladding embrittlement. Breakaway oxidation is often apparent from the visual appearance 

of the outer surface of the oxide layer. The time at which breakaway oxidation occurs is commonly

determined from the increase in weight-gain rate. However, a more reliable and relevant indicator is the

measured increase in hydrogen concentration. This is particularly true if breakaway oxidation is due to the 

presence of certain impurities such as fluorine or fluorides, which can result in high hydrogen uptake 

without appreciable increase in weight-gain rate [19]. In Ref. 1, 200-wppm hydrogen pickup was used to 

determine breakaway oxidation time. This metric is also used in the current work. Leistikow and Schanz 

[20] conducted a comprehensive study of breakaway oxidation during the 1980s, but its degrading effect 

on cladding ductility was not known at that time. Results are presented in Ref. 1 for cladding ductility 

decrease with increasing hydrogen pickup and for hydrogen levels at which embrittement was observed. 

Breakaway oxidation time has been shown to be sensitive to surface finish and methods used to produce 

zirconium ingots. Sensitivity to surface finish (i.e., roughness, scratches, impurities) was demonstrated in 

the Argonne test program [1]. Sensitivity to zirconium-ingot preparation was demonstrated in the Russian 

test program [25] sponsored by NRC and IRSN and in subsequent work [26]. 

Measured minimum breakaway-oxidation times were reported in Ref. 1 for a number of cladding 

materials. These are summarized in Table 21. 

Table 21. Minimum breakaway oxidation times and corresponding oxidation temperatures 

reported in Ref. 1 for as-fabricated cladding alloys.

Cladding Material Minimum Breakaway Time, s Oxidation Temperature, °C

Zry-4 (vintage, unpolished) 3800 985°C

Zry-4 (modern, wheel polished) ≈5000 985°C

Zry-2 (modern, belt polished) >5000 970–1000°C

ZIRLO (modern, belt polished) 3100±300 970–985°C

M5 (modern, wheel polished) ≈6400 [9] 1000°C

E110 (vintage, unpolished, pickled) 500-700 1000°C

In the current work, additional tests have been conducted with ZIRLO and M5 cladding oxidized at 

constant temperatures and with ZIRLO cladding oxidized under transient temperature conditions to 

improve understanding of these phenomena for conditions relevant to LOCA safety analyses.
72

4.1 Breakaway Oxidation for ZIRLO at Constant Oxidation Temperature

As the previously reported breakaway time for ZIRLO was relatively short compared with other modern 

cladding materials, additional measurements were performed along with careful reanalysis of data 

generated previously. New measurements were made with some ANL-cleaned specimens and some 

Westinghouse (WEC or W)-cleaned specimens, and these new tests will be highlighted below. To present 

the new measurements in the context of the older measurements, all relevant results are summarized in 

the following sections of this report.

The ANL criterion for breakaway oxidation is the time corresponding to 200-wppm hydrogen pickup 

(ΔCHM), where the reference mass for this parameter is the pre-oxidation cladding metal mass. However, 

Zry-4 and ZIRLO materials retained ductility at higher hydrogen contents. Ring compression tests at 

135oC indicated ductility at ≤440±110 wppm ΔCHM and embrittlement at ≥770±120 wppm ΔCHM. The ± 

value represents one standard deviation based on hydrogen measurements for multiple arc segments of 8-

mm-long compressed rings. Because of the rapid hydrogen pickup with time, the transition between 

ductile and brittle behavior was as short as 200 s at 1000oC. Thus, 200-wppm hydrogen pickup is a 

reasonably conservative criterion for breakaway oxidation based on embrittlement. 

The cladding materials used for previous studies included ZIRLO received from WEC in 2003 (ZIRLO2003) and in 2006 (ZIRLO-2006), ZIRLO-2006 modified by ANL by machining a 20-μm deep scratch 

into the cladding wall along the length of the sample (ZIRLO-2006MS), and ZIRLO received from WEC

with a thin (<1 μm) oxide film grown at low temperature in pressurized water (ZIRLO-PF). Results for 

ZIRLO-2003 are not included in this section because this material was highly susceptible to breakaway 

oxidation on the inner surface of the cladding. Excluding the short ballooned region of the cladding, only 

the outer surface of the cladding is exposed to steam oxidation. Thus, it is important to demonstrate that 

breakaway oxidation occurs at the cladding outer surface and that the measured hydrogen pickup includes 

only hydrogen absorbed at the cladding outer surface. All results reported in Ref. 1 were generated using 

a long quartz-tube steam chamber (686-mm long). At the time Ref. 1 was published, ANL was not able to 

demonstrate the early breakaway times for ZIRLO in a short (610-mm long) steam chamber, based on a 

very limited number of tests conducted. 

At 1000oC, the minimum breakaway time for ZIRLO was ≈4000 s based on tests at 3600s (60 wppm 

ΔCHM), at 4000 s (130 wppm ΔCHM), and at 4200 s (630 wppm ΔCHM). As breakaway oxidation is an 

instability phenomenon due to local transformation from the protective, tetragonal oxide phase (lustrous 

black in appearance) to the non-protective, monoclinic oxide phase (yellow in appearance for ZIRLO), 

multiple tests under the same time-temperature conditions would have generated scatter of at least ±200 s 

at 1000oC. Thus, the time for ZIRLO breakaway at 1000oC was assessed to be 4000±200 s based on the 

limited data set generated. It should be noted that test times reported by ANL include the duration from 

the initiation of the temperature ramp at 300oC to the end of the heating phase. Typical ANL heating 

curves consisted of a very fast ramp from 300oC to 950oC in <20 s and a slower ramp from 950oC to 

1000oC in ≈60 s to limit the short-time temperature overshoot to ≤10oC. Thus, breakaway oxidation times 

reported in Ref. 1 included about 80-s heating ramps from 300oC to 1000oC 

Tests at higher (1015oC) and lower (950-985oC and 800oC) temperatures were conducted to determine the 

temperature sensitivity of breakaway oxidation and the minimum breakaway oxidation time. Based on a 

more extensive data set for as-fabricated, machine-scratched, and pre-filmed ZIRLO, 3100±300 s was 

determined to be the minimum breakaway time within the temperature range of 970-985oC. These earlier

studies were conducted in two campaigns with separate test trains and associated thermal benchmarks: 

tests with ZIRLO-2006 and ZIRLO-2006MS within a wide temperature range and tests with ZIRLO-2006 
73

and ZIRLO-PF within a narrow temperature range (970-985oC). Results for two samples tested with a 

machined scratch indicated that local breakaway occurred early along the scratch, but the presence of the 

scratch had little effect on the reduction in breakaway time (<200 s) based on the 200-wppm criterion. 

Tests with ZIRLO-PF demonstrated that the presence of a thin oxide layer grown at reactor-relevant 

temperature had no significant effect on breakaway oxidation time. 

WEC [27] generated independent data sets for 15×15 Zry-4 and 17×17 ZIRLO. The WEC-determined 

minimum breakaway oxidation times for Zry-4 (<4400 s at 980oC for one bare sample and ≈4400 s at 

1010oC for one pre-filmed sample) were consistent with ANL results (3800 s for vintage material and 

5000 s for modern material at 985oC, both 15×15 Zry-4 manufactured by AREVA). WEC results for 

ZIRLO oxidized at 1000oC were in agreement (within data scatter) with ANL results. However, WEC 

results indicated longer breakaway oxidation times than ANL results at <1000oC. In an attempt to resolve

the differences, WEC subsequently provided ANL with additional ZIRLO specimens that had been 

cleaned by WEC using their experimental procedures: chemical detergent (Alconox) treatment prior to 

water rinsing. ANL used an organic solvent (ethanol) prior to water rinsing. Although it was not 

anticipated that this would account for differences in results, it was the easiest factor to test and possibly 

eliminate. ANL had used ethanol, which never came in contact with the metal, to clean ZIRLO-PF 

samples prior to oxidation. Therefore, early breakaway times observed for ZIRLO-PF samples could not 

have resulted from an ethanol-metal surface reaction during cleaning. 

WEC supplied ANL with ten samples of W-cleaned ZIRLO (ZIRLO-WC), as well as a new supply of 

ZIRLO (ZIRLO-2008). The following sections describe the results of new tests performed by ANL with 

both W-cleaned (Alconox) and ANL-cleaned (ethanol) ZIRLO samples. To help resolve the issue of long 

vs. short steam chambers, these new tests were conducted using both steam chambers. 

4.1.1. ZIRLO materials for breakaway oxidation 

Table 5 summarizes geometrical parameters, oxygen content and hydrogen content measured by ANL for 

the larger lots of W-supplied ZIRLO. The ten W-cleaned ZIRLO samples were very similar to the ZIRLO 

received in 2006 to 2010: 9.50-mm OD, 0.584±0.003-mm wall thickness, 0.16±0.01-μm OD surface 

roughness, and 22±2-wppm hydrogen. 

4.1.2. Breakaway oxidation test parameters and results

Tests at 1000oC in long steam chamber

Two new oxidation tests were conducted in the long steam chamber with a standard-length test train. Wand ANL-cleaned samples were oxidized for 4000 s at 1000oC. The thermal benchmark for these tests is 

shown in Fig. 35. The benchmark was performed for a long-time hold temperature of 975±5oC, a 

temperature at which additional tests were to be conducted. The actual test time (3100 s from ramp 

initiation to the end of the hold time) was longer than the time (1200 s) shown in Fig. 35. Temperature 

readings were recorded every 0.1 s and EXCEL has a limit (32000) on the number of data points that can 

be plotted for a single curve. The 10oC circumferential temperature variation is based on readings of two 

TCs welded onto the sample. These TCs reached 975oC in ≈80 s, experienced a 10oC overshoot from 80 s 

to 160 s, and then decreased to a hold temperature of 975oC for the time duration of about 1500 s. The 

slow decrease from 980 to 975oC occurred during a period in which the control thermocouple on the 

sample holder had a constant reading. As such, in addition to ±5oC circumferential temperature variation, 

there may have been a 5oC uncertainty in long-time sample TC readings. For the current tests, the furnace 

power was increased to achieve a long-time hold temperature of 1000oC.
74

Figure 35. Thermal benchmark results with two TCs welded onto as-fabricated, 17×17 ZIRLO-2008 cladding. T

hold temperature was 975±5oC. For data-generating tests, the furnace power was increased to give 1

temperature.

900

800

700

600

500

400

300

1100

1000

500 700 900 1100 1300 1500 170

C) ° Temperature (

Time (s)

TC1: on holder

TC2: on holder

TC3: on holder

TC4: on Sample

TC5: on Sample
75 

Figure 36 shows the visual appearance of the two test samples. It is clear that the ANL-cleaned sample 

experienced more extensive breakaway oxidation than the W-cleaned sample. However, the W-cleaned 

sample had small yellow spots on the outer surface indicating the initiation of local breakaway oxidation. 

Figure 36. Surface appearance of W- and ANL-cleaned ZIRLO samples after 4000 s at 1000oC.

The ANL-cleaned sample, with the extensive yellow regions, picked up 400±240 wppm hydrogen near 

the sample mid-span, while the W-cleaned sample picked up 120±110 wppm hydrogen near the mid-span. 

The hydrogen results for the W-cleaned sample indicated that local breakaway oxidation had occurred. 

Metallographic examination was performed to verify that local breakaway of the outer surface oxide was 

responsible for the hydrogen pickup in the W-cleaned sample. Images were taken at eight locations 

corresponding to angular orientations separated by ≈45o. One location indicated breakaway of the outersurface oxide (Fig. 37a) which included both cracked monoclinic outer-surface oxide and tetragonal 

oxide. For the other locations, no tetragonal-to-monoclinic phase change was observed at the outer 

surface. For all images, the inner-surface oxide showed no signs of breakaway oxidation. The 

metallographic results are consistent with the local hydrogen readings: high hydrogen content under 

yellow spots and low hydrogen content in regions under lustrous black oxide. Based on ANL experience, 

the W-cleaned sample was <200 s away from picking up >200 wppm hydrogen. As shown in Table 22, 

this result is consistent with results previously obtained for ANL-cleaned samples. Thus, for an oxidation 

temperature of 1000oC, no significant effects of Alconox vs. ethanol cleaning were observed. 
76 

(a) Area 8: local hydrogen content >280 wppm under yellow outer-surface spot

(b) Area 7 (c) Area 1

Figure 37. Metallographic images (at the same cross section) for W-cleaned sample oxidized 

for 4000 s at 1000oC: (a) Area 8, which has outer-surface monoclinc oxide (≈0.4 mm 

long); (b) Area 7 with tetragonal oxide; and (c) Area 1 with tetragonal oxide. 
77

Table 22. Results of ZIRLO breakaway oxidation tests at 1000oC; new results are in bold font. 

Material

Test Timea

, 

s 

Measured

wg, 

mg/cm2

Measured

(CH)POX, 

wppm

ΔCHM

 b, 

wppm Comment

ZIRLO-2006 1500 8.7 15 5 ---

ZIRLO-2006 3600 11.4 68 60 ---

ZIRLO-2006 4000 12.6 131±12 130±10 ---

ZIRLO-2008 4000 13.3 390±220 400±240 See Fig. 36

ZIRLO-WC 4000 12.6 130±100 120±110 See Fig. 36

ZIRLO-2006 4200 14.3 610±370 630±400 

a

Time from ramp initiation at 300oC to end of hold temperature; ramp time is ≈80 s.

b

ΔCHM is referenced to as-fabricated sample mass: ΔCHM = (1 + 5.4×10-3 wg) (CH)POX - CHi, where CHi is 

the hydrogen content in as-fabricated ZIRLO. 

The hydrogen content shown in Fig. 38 for the new ANL-cleaned sample was measured for a 2-mm-long 

ring sectioned from near the center of the sample. An 8-mm-long ring was sectioned from the adjacent 

area and subjected to a ring-compression test at 135oC. The sample proved to be extremely brittle and 

failed at a much lower load than other samples tested. The hydrogen content measured for the 8-mm-long, 

compressed ring (555±133 wppm) was even higher than for the central 2-mm-long ring (390±220 wppm). 

The corresponding hydrogen pickup for the 8-mm-long ring was 600±150 wppm. This ring-compression

data point is plotted in Fig. 39, along with the data in NUREG/CR-6967 Fig. 63, to better define the 

hydrogen range for ductile-to-brittle transition for breakaway oxidation samples. Based on Fig. 39, the 

ductile-to-brittle transition hydrogen level for samples oxidized at 970-1000oC is between 440 and 600 

wppm, which gives ≈500±50 wppm for the transition. Also, the circumferential variation in hydrogen 

content (greater than ±100 wppm) was very large for average hydrogen contents in this range (Fig. 40). 

Breakaway oxidation tests were terminated by furnace cooling rather than quench at 800oC. There were 

two reasons for conducting these tests with slow cooling from 800oC: (a) for postulated LOCAs of long 

duration, the temperature range during which rapid cooling might occur due to wetting may be much 

lower than 800oC and (b) the vibration and thermal stresses during quench might cause some oxide to 

flake off, particularly oxide that had experienced breakaway. From the experimental point of view, slow 

cooling was used to preserve the oxide layer for visual and metallographic examination. Based on the 

cooling method used for these tests, it is not clear if the data points shown in Fig. 39 for hydrogen levels 

>440 wppm would shift somewhat to the left if the samples had been quenched at 400 to 800oC. Thus, it 

is recommended that the lower-bound value of ≈450 wppm be used for the ductile-to-brittle transition.
78

Figure 38. Hydrogen pickup for ANL-cleaned and W-cleaned ZIRLO samples oxidized at 

1000oC. 

Figure 39. RCT offset strains at 135oC for Zry-4 and ZIRLO rings following oxidation at 970-

1000oC, local-to-global breakaway oxidation, and cooling without quench. 

0

200

400

600

800

3500 3700 3900 4100 4300

Hydrogen Pickup (wppm)

Oxidation Time (s)

ANL-Cleaned Samples

W-Cleaned Sample

Best Linear Fit

0

1

2

3

4

5

6

7

0 200 400 600 800 1000 1200 1400

Offset Strain (%)

Hydrogen Pickup (wppm)

BP 15x15 Zry-4

HBR-type 15x15 Zry-4

BP 17x17 ZIRLO
79

Figure 40. Circumferential variation in hydrogen content measured after ring-compression 

testing for an 8-mm-long ring sectioned from the ANL-cleaned ZIRLO sample 

oxidized at 1000oC for 4000 s. 

Tests at ≈980oC in short steam chamber with temperature overshoot

Concurrent with the new breakaway oxidation studies, a test train for the shorter steam chamber was 

thermally benchmarked for a hold temperature of 1200oC to conduct high-temperature embrittlement 

tests. It was time-consuming to switch back and forth between test trains and chambers for breakaway and 

high-temperature embrittlement studies. Therefore, the short test chamber was used for the next sequence 

of new breakaway tests, and the furnace power was scaled to give a long-time hold temperature of 

≈985oC. Two tests were run for 3200 s with W-cleaned samples, and one test was run for 3400 s with an 

ANL-cleaned sample. All three samples exhibited significant yellow areas on the outer surfaces indicating 

extensive breakaway. Following these tests, TCs were welded onto a fresh ZIRLO sample to determine 

more precisely the temperature history. As shown in Fig. 41, the scaling of furnace power resulted in a 

long-time hold temperature of 981±4oC (≈980±5oC). However, unlike other histories, this one had 

considerable overshoot early in the transient. The cladding temperature rose rapidly (18 s) to ≈1065oC and 

was above 985oC for ≈20 s. Other than a faster rise to high temperature, it is not clear what effect an early 

temperature excursion would have on breakaway oxidation occurring ≈3000 s later. As the purpose of the 

tests was to compare the performance of ANL- and W-cleaned ZIRLO samples exposed to the same 

steam flow and temperature conditions, the results from this series of tests are considered valid for that 

purpose. The hydrogen pickup was high (≈700 wppm) for the samples and well beyond the 200-wppm 

ΔCHM criterion for breakaway. These three new results are significant because they showed no apparent 

difference between the W- and ANL-cleaned samples (Fig. 42). Also, for the first time at ANL, 

breakaway oxidation was observed using the short steam chamber. 
80

Figure 41. Thermal benchmark results for test ZLU#129. Peak temperature was ≈1065oC and long-time hold te

≈980±5oC.

900

800

700

600

500

400

300

1100

1000

500 600 700 800 900 1000 1100 

Temperature (oC)

Time (s)

TC4 on sample

TC5 on sample
81 

(a) ANL-cleaned sample after 3400 s at 980oC 

(b) W-cleaned sample #1 after 3200 s at 980oC 

(c) W-cleaned sample #2 after 3200 s at 980oC 

Figure 42. Surface appearance of ANL-cleaned and W-cleaned samples after oxidation at 

980oC: (a) ANL-cleaned sample after 3400 s; (b) W-cleaned sample #1 after 3200 s; 

and (c) W-cleaned sample #2 after 3200 s. 
82 

Tests at 985oC in short steam chamber with no temperature overshoot

An additional set of new tests was conducted with the same short test train and steam chamber as was 

used to generate the previous results. However, furnace control parameters were reset to eliminate 

temperature overshoot. The temperature history determined from the new thermal benchmark is shown in 

Fig. 43. The long-time hold temperature was 975±3oC. Also shown in Fig. 43 are the histories of the 

control TC and one other TC, both welded to the sample holder. The control TC readings were only a few 

degrees lower than the sample readings. As the data-generating tests were to be conducted at 985oC, the 

controller was adjusted to increase the control TC by 10oC and to correspondingly increase the sample 

temperature to 985oC (985±5oC). One ANL-cleaned sample was oxidized for 3500 s and had a post-test 

hydrogen pickup of 720 wppm. The W-cleaned samples picked up 20 wppm after 3500 s and 270 wppm 

after 3800 s. Although breakaway oxidation was somewhat delayed for the W-cleaned samples, the 

results were within data scatter for this oxidation temperature.

Table 23 summarizes the six new results generated in the temperature range of 980-985oC (short steam 

chamber), as well as previous results generated for as-fabricated, machine-scratched, and pre-filmed 

ZIRLO in the temperature range of 970-985oC (long steam chamber). Hydrogen pickup values are plotted 

in Fig. 44 as a function of test time for all data in Table 23. Based on these results, as well as expected 

data scatter for an instability phenomenon such as breakaway oxidation, it is clear that there were no 

measurable effects of Alconox (WEC) vs. ethanol (ANL) pre-test cleaning.

In parallel with the ANL testing, WEC conducted a series of breakaway oxidation tests with ANL-cleaned 

samples and WEC-cleaned samples (ZIRLO-2008 and another lot of ZIRLO) [28]. WEC tests were 

performed in a resistance-heating furnace for which the circumferential temperature variation is expected 

to be small. Test temperatures were between 960oC and 1020oC based on measured sample-holder 

temperatures. Ramp times ranged from 60 to 270 s and hold times ranged from about 3700 to 4500 s. 

Breakaway oxidation was assessed based on the appearance of the cladding outer surface. For 1000oC

tests, breakaway was observed to have occurred for three samples at ≈4400 s (60 s ramp with 4330 s hold 

time), while no breakaway was observed at ≈4550 s (130 s ramp with 4410 s hold time). WEC samples 

oxidized with the faster ramp exhibited large discolored regions indicating that the minimum breakaway 

time was <4400 s. These results were consistent with ANL results for minimum breakaway oxidation 

time (4000±200 s). WEC speculated that there was a significant effect of temperature ramp rate on 

breakaway oxidation because the longer ramp-rate tests showed no evidence of breakaway at a total test 

time of 4550 s. However, for both “fast” and “slow” ramp times, the ramp rate was about the same up to 

980oC. It is difficult to comprehend why an extra 70 s to progress from 980oC to 1000oC would make a 

difference in results observed about 4000 s later. Also, it is not clear if apparent differences in three slowramp samples and three fast-ramp samples were due to data scatter or to differences in thermal history.

The WEC results for hold temperatures of 980oC and 990oC were quite different from ANL results. For 

the lower temperature regime, WEC observed no breakaway for total test times of 3800 to 4600 s. Based 

on a larger data set, ANL determined 3100±300 s for average hold temperatures of 970oC to 985oC. One 

possible explanation for these different observations is that ANL samples experienced the full range of 

temperatures between 965oC and 990oC while WEC samples were subjected to discrete temperatures. If 

the minimum breakaway time for ZIRLO is highly temperature sensitive, then this could explain why 

ANL and WEC data do not appear to be consistent in the lower temperature range. WEC did not measure 

sample temperatures directly nor did WEC perform weight-gain benchmarks prior to data generation. 

Weight-gain validation tests were performed successfully by WEC about two years later [29] after 

modifications were made to the furnace.
83

Figure 43. Thermal benchmark results for test ZLU#132. Long-time hold temperature is 975±3oC. The holder c

was increased by 10oC for data generating tests, which increased the long-time sample temperature t

900

800

700

600

500

400

300

1100

1000

490 990 1490 1990 2490 2990 

C) ° Temperature (

Time (s)

 

TC2 on Holder (Control)

TC3 on Holder

TC4 on Sample at 0o

TC5 on Sample at 120o
84 

Table 23. Results of ZIRLO breakaway oxidation tests at 970−985oC; new results are in bold font. 

Material

Test T

oC

Test 

Timea

, 

s 

Measured

wg, 

mg/cm2

Measured

(CH)POX, 

wppm

ΔCHM

 b, 

wppm Comment

ZIRLO-2006MS 970 2600 7.8 50 40 Scratched

ZIRLO-2006 970 3000 8.1 200±150 200±160 ---

ZIRLO-2006 970 3400 8.6 570±90 580±90 

ZIRLO-2006 975 3000 9.0 40 30 ---

ZIRLO-2006 975 3500 8.7 110±50 100±50 ---

ZIRLO-2006 980 3000 11.8 350±120 370±130 

ZIRLO-PF 980 3200 8.7 130±120 120±120 

ZIRLO-PF 980 3200 8.8 100±100 100±100 

ZIRLO-WC 980 3200 8.4 690±90 700±100 W-cleaned

T-overshoot 

ZIRLO-WC 980 3200 8.4 630±120 650±120 W-cleaned

T-overshoot 

ZIRLO-2008 980 3400 11.4 --- --- Breakaway 

T-overshoot 

ZIRLO-2006 985 3000 9.1 100±40 100±40 ---

ZIRLO-PF 985 3000 9.0 60 50 Pre-filmed

ZIRLO-2006 985 3400 9.6 50 40 ---

ZIRLO-2006MS 985 3400 9.9 180±140 170±150 Scratched

ZIRLO-2006 985 3400 9.4 300±160 310±170 

ZIRLO-PF 985 3400 9.3 260±140 260±150 Pre-filmed

ZIRLO-2008 985 3500 9.5 700±120 720±120 ---

ZIRLO-WC 985 3500 9.2 40 20 W-cleaned

ZIRLO-2006 985 3600 10.0 280±160 280±170 ---

ZIRLO-WC 985 3800 8.9 280±60 270±60 W-cleaned

ZIRLO-PF 985 4000 11.9 1070±90 1130±100 Pre-filmed

ZIRLO-2006 985 4000 11.8 850±140 890±150 ---

a

Time from ramp initiation at 300oC to end of hold temperature; ramp time is ≈80 s.

b

ΔCHM is referenced to as-fabricated sample mass: ΔCHM = (1 + 5.4×10-3 wg) (CH)POX - CHi, where CHi is 

the hydrogen content in as-fabricated ZIRLO. 
85 

Figure 44. Hydrogen pickup for ANL- and W-cleaned samples oxidized at 970−985oC. ANLcleaned samples included as-fabricated, MS (20-μm depth), and PF (<1-μm oxide) 

ZIRLO.

4.2 Breakaway Oxidation for M5 Cladding

The purpose of these tests was to determine if as-fabricated M5 cladding would experience breakaway 

oxidation at ≤5000 s for oxidation temperatures in the range of 985-1000°C. ANL results for oxidation of 

M5 cladding at 1000°C for ≤4100 s are documented in Ref. 1. Based on the large data set generated by 

CEA at Saclay [9], the breakaway oxidation time (at ≈200-wppm H pickup) is ≈6400 s for M5 oxidized at 

1000oC. M5 exhibits very slow growth rate of inner- and outer-surface oxide layers at 1000oC, which may 

contribute to the longer breakaway-oxidation time for M5 as compared to the CEA-value reported for 

Zry-4 (≈5400 s at 1000oC). The Ref. 9 data were generated by very different testing methods (e.g., onesided oxidation, fast heating ramp, etc.) than the ones used at ANL. Also, breakaway-oxidation times for 

M5 at <1000oC oxidation temperatures have not been reported. 

4.2.1 M5 materials for breakaway oxidation 

Cladding materials used for the new ANL tests included as-fabricated 17×17 M5 tubing received from 

Framatome in 2003 (M5-2003) and M5 received from AREVA (formerly Framatome) in 2007 (M5-

2007). Table 24 shows the comparison between the two materials based on ANL measurements.

0

200

400

600

800

1000

1200

2400 2600 2800 3000 3200 3400 3600 3800 4000

Hydrogen Pickup (wppm)

Oxidation Time (s)

ANL-Cleaned Samples

W-Cleaned Samples

Best Linear Fit
86 

Table 24. Characterization of 17×17 M5 cladding used in the ANL testing program.

Material

Outer Diameter, 

mm

Wall thickness, 

mm

OD Surface 

Roughness, μm 

H Content, 

wppm

O Content, 

wppm

M5-2003 9.50 0.61 0.12 5 1450 

M5-2007 9.50 0.57 -- 17 --

4.2.2 Test parameters and results for M5 breakaway oxidation 

Results for M5 oxidation tests conducted at 1000oC for ≤4100 s are reported in Ref. 1. Three additional 

tests were conducted at ≈985oC for 5000 s during the current test program.

Previous test results for M5 at 1000oC 

Thermal benchmark results for M5 (Validation Lot with 0.57-mm h) cladding are shown in Fig. 45 for 

1000oC tests. The results were used for M5-2003 with 0.61-mm h because the results for 0.57-mm 

cladding were so close to the holder temperatures. Table 25 lists the breakaway-oxidation test results for 

M5 cladding samples. For 1000oC oxidation, M5 exhibited significantly lower weight gain and slower 

oxide-layer growth as compared to Zry-4 and to CP-predicted values. Also, above ≈10% CP-ECR (≈960 s 

equivalent isothermal time), weight gain did not increase in a consistent manner with either the oxidation 

time or with the square root of oxidation time (CP correlation). All samples had lustrous black outer 

surfaces. The measured hydrogen pickup was low. Breakaway oxidation was not observed at 1000oC for 

test times ≤4100 s.

New Tests at 980oC and 985oC

During previous studies, minimum breakaway oxidation times were observed by Argonne for ZIRLO for 

temperatures in the range of 970-985oC. Unlike results at 1000oC, there was considerable scatter in 

hydrogen pickup vs. oxidation time within this highly sensitive temperature range. M5-2007 breakaway 

oxidation tests were therefore performed at 980oC and 985oC for 5000 s in parallel with those conducted 

using ZIRLO cladding. 

Both ZIRLO and M5-2007 samples had the same outer-diameter (OD) and wall thickness. Therefore,

thermal benchmark temperature histories determined for ZIRLO were also used for M5 (see Fig. 41 for 

test MU#80 and Fig. 43 for test MU#83). For test MU#81, control parameters were adjusted to give a 

slower ramp rate to the target temperature.

Table 26 lists the weight gain results for M5-2007 for long-time hold temperatures of 980oC and 985oC. 

Breakaway oxidation was not observed for test times as long as 5000 s, which was confirmed by 

metallographic examination of the MU#80 test sample (see Fig. 46). 
87 

Figure 45. Thermal benchmark results for 17×17 M5 (Validation Lot) oxidation tests at 

1000oC. 

Table 25. Results of previous M5 oxidation tests at 1000oC in long test train. 

Test ID Material

Test Timea, 

s 

CP-predicted

wg, 

mg/cm2

Measured

wg, 

mg/cm2

(CH)POX, 

wppm

ΔCHM

 b, 

wppm,

MU#52 M5-2003 207 3.3 3.0 --- ---

MU#53 M5-2003 810 6.5 5.3 --- ---

MU#54 M5-2003 1880 9.8 7.4 --- ---

MU#55 M5-2003 2440 11.1 8.0 --- ---

MU#56 M5-2003 3380 13.1 9.2 26 22 

MU#61 M5-2003 4100 14.4 8.6 13 8 

a

Time from ramp initiation at 300oC to end of hold temperature; ramp time is ≈80 s.

b

ΔCHM is referenced to as-fabricated sample mass: ΔCHM = (1 + 5.4×10-3 wg) (CH)POX - CHi, where CHi is 

the hydrogen content in as-fabricated M5. 

200

300

400

500

600

700

800

900

1000

1100

200 300 400 500 600 700 800 900 1000 1100 1200 1300

Temperature (°C)

Time (s)

TE0 (on M5 at 0°)

TE1 (on holder at 0°)

TE3 (on holder at 240°)

TE4 (on M5 at 240°)
88 

Table 26. M5-2007 breakaway oxidation test results for long-time hold temperatures of 980oC and 

985oC in the short test train. 

Test ID

Long-Time

Hold T,

°C

Test Timeb

,

s 

CP-predicted

wg, 

mg/cm2

Measured

wg, 

mg/cm2

 

(CH)POX, 

wppm

ΔCHM c, 

wppm,

MU#80 985 5000 14.4 7.9 32 28 

MU#81 980 5000 ≈14.4 7.7 --- ---

MU#83a 985 5000 ≈14.4 7.4 --- ---

a

The ramp rate for this test was close to the one shown in Fig. 44.

b

Time from ramp initiation at 300oC to end of hold temperature. 

c

ΔCHM is referenced to as-fabricated sample mass: ΔCHM = (1 + 5.4×10-3 wg) (CH)POX - CHi, where CHi is 

the hydrogen content in as-fabricated M5. 

4.3 Discussion on Breakaway Oxidation under Isothermal Conditions

Leistikow and Schanz [20] observed a transition in the contour of the oxide-metal interface from stable, 

tetragonal oxide growth to breakaway oxidation. For Zry-4, the interface progressed from smooth to 

scalloped to smooth. Thus, they referred to the scalloped interface as a pre-cursor to breakaway oxidation. 

They further speculated that the oxide was growing stably under compressive stress in the thicker regions, 

while the thinner regions would become unstable under a tensile (or less compressive) stress because of 

the contour. Thus, destabilization or breakaway would initiate at the metal-oxide interface under the thin 

regions and propagate in the radial, circumferential and axial direction. However, the time between 

initiation of the scalloped interface and breakaway was not established. Additional metallographic 

examination was performed with Zry-4, ZIRLO, and M5 test samples to shed some light on the amplitude 

and frequency of scallops prior to breakaway oxidation or close to the initiation of breakaway oxidation.

Figure 47 shows the evolution of the outer-surface oxide-layer/metal interface for ZIRLO oxidized at 

1000oC. After 1500 s (Fig. 47a), the interface was relatively smooth with the appearance of very lowamplitude scallops. The amplitude of the scallops grew from 1500 s to 2440 s (Fig. 47b). At 3480 s, the 

high magnification image in Fig. 47c shows rather pronounced scallops. The image at 4000 s is from the 

W-cleaned ZIRLO sample that experienced local breakaway oxidation at a location about 140° from the 

location shown in Fig. 47d. Although not enough data were generated to directly correlate the amplitude 

of the scallops with initiation of local breakaway oxidation, the data trend shown in Fig. 47 is consistent 

with the Leistikow and Schanz speculation that scalloping is a precursor to breakaway oxidation. The 

trend is also consistent with the ANL hypothesis that breakaway initiation occurs at a critical scallop 

amplitude. 

It should be noted that scalloping appears to result in measured weight gain and average oxide layer 

thickness values that are less than those predicted by the CP correlations. The results presented in Ref. 1 

for ZIRLO-2003 were misleading because the inner surface (66 μm vs. 57 μm outer-surface oxide layer) 

experienced breakaway at 1000°C prior to 3480 s (20% CP-ECR). The measured weight gains for prebreakaway ZIRLO oxidized at 970 to 1000oC were less than CP-predicted weight gains. 
89

Figure 46. Metallographic images around the cross section of the MU#80 test sample, which show wavy interfac

metal layers. However, no local breakaway oxidation was observed after 5000 s at 985oC.
90 

(a) 1500 s 

(b) 2440 s 

(c) 3480 s 

(d) 4000 s 

Figure 47. Evolution of outer-surface oxide-layer/metal interface for ZIRLO oxidized at 

1000°C for (a) 1500 s; (b) 2440 s; (c) 2480 s; and (d) 4000 s. 
91 

M5 exhibited slower oxide layer growth during oxidation at ≈1000°C relative to ZIRLO and Zry-4. Thus, 

the average thickness of the oxide layer, as well as the amplitude of the scallops at the oxide/metal 

interface, grew slowly with oxidation time within a temperature range of about 975°C to 1000°C. The 

slow oxidation kinetics appeared to contribute to the longer breakaway oxidation time for M5 relative to 

Zry-4 and ZIRLO. Figure 48 shows the oxide/metal interface for M5 oxidized at 1000°C for 3440 s (highmagnification image in Fig. 48a) and at 985°C for 5000 s (low- and high-magnification images in Figs. 48 

b and c, respectively). 

Based on hydrogen-content data in Ref. 9, it was estimated that the breakaway oxidation time (i.e., time to 

reach 200 wppm ΔCHM) would be ≈6400 s at an oxidation temperature of 1000°C. More recently, CEA, 

EdF, and AREVA collaborated to do a more extensive study of the breakaway oxidation time for standard 

M5, as well M5 with variations in Nb, Fe, O, S, and Hf content [30]. For standard M5, the minimum 

breakaway oxidation time for 200-wppm hydrogen was in the range of ≈6500±500 s. Variations in alloy 

and impurity contents had little effect on breakaway oxidation time within the data scatter for standard 

M5. It was also observed that both the weight-gain and hydrogen-pickup rates were relatively slow during 

the transition between pre- and post-breakaway oxidation. 

For polished Zry-4 oxidized at 1000°C, the outer-surface oxide-layer/metal interface was relatively 

smooth for test times up to ≈3400 s. Figure 49 shows the outer-surface oxide-layer/metal interface for 

Zry-4 oxidized for: (a) 3440 s at 1000°C and (b) 5000 s at 985°C. Pronounced scallops can be observed at 

the higher oxidation time for which local breakaway oxidation occurred at a location away from the one 

shown in the figure. The local breakaway oxidation is identified in Fig. 50 at one of eight metallographic 

images around the cross-section.

To develop a correlation between scallop amplitude and breakaway-oxidation initiation, additional 

oxidation tests would have to be performed at intermediate oxidation times. Also, metallographic samples 

would have to be prepared carefully (i.e., additional polishing steps with finer-grit paper) to avoid local 

fallout of the brittle oxide and/or oxygen-stabilized alpha material. Fallout tends to exaggerate the 

amplitude of the scallops. Examples of high-quality metallographic images are shown in Figs. 47a and c, 

Fig. 48a and the right half of Fig. 48c, and Fig. 49a. 
92 

(a) 3440 s at 1000°C 

(b) 5000 s at 985°C (low magnification) 

(c) 5000 s at 985°C 

Figure 48. Evolution of outer-surface oxide-layer/metal interface for M5 cladding oxidized for: 

(a) 3440 s at 1000°C; (b) 5000 s at 985°C (low-magnification); and (c) 5000 s at 

985°C.
93 

(a) 3440 s at 1000°C 

(b) 5000 s at 985°C 

Figure 49. Evolution of oxide-layer/metal interface for polished Zry-4 oxidized for: (a) 3440 s 

at 1000°C and (b) 5000 s at 985°C. 
94

Figure 50. Metallographic images at eight locations around the cross section of polished Zry-4 sample oxidized 

which show scalloped (wavy) oxide-layer/metal interfaces. Local breakaway oxidation was observed 
95 

4.4 Breakaway Oxidation under Transient Temperature Conditions

The purpose of the transient oxidation tests was to determine if temperature changes might decrease 

breakaway oxidation time. The study was not intended to simulate a temperature history for a particular 

LOCA, which would be difficult to do as there is no standard temperature history, especially for smallbreak LOCAs. As such, temperature histories were chosen such that they might disturb the stress state at 

the oxide-metal interface and thereby induce lower breakaway oxidation times. The approach was based 

on hypotheses of Leistikow and Schanz [20]: (a) the transition from a smooth oxide-metal interface to a 

wavy oxide-metal interface is a precursor to breakaway oxidation and (b) the variation in stress state from 

thick oxide regions (compressive) to thin oxide regions (less compressive and perhaps tensile) facilitates 

local breakaway oxidation in regions where the oxide layer is thin. The results of Argonne’s isothermal 

tests support their hypotheses. 

4.4.1 Transient temperature histories

Transient breakaway oxidation tests were conducted with three temperature profiles. Schematics of these 

three temperature profiles are shown in Fig. 51. 

The first transient test was designed to have a temperature history corresponding to Transient #1 in Fig. 

51. This history included a single step change from 1045°C to 980°C: (a) hold at 1045°C for ≈1000 s, (b) 

decrease temperature to 980°C and (c) hold for ≈3000 s at 980°C. Test ZLU#138 was conducted as a 

thermal benchmark with 2 TCs welded directly onto the ZIRLO sample outer surface. Benchmark results 

are shown in Fig. 52. The long-time hold temperatures were 1041±1oC during the last 500 s of step (a) 

and 981±1oC during the last 500 s of step (b). Except for the initial temperature ramp, the holder control 

TC measurements were only a few degrees lower than the sample TCs. The first 500 s (not shown) was 

the preconditioning phase to stabilize the temperature at 300oC. The ramp time to 1045oC was ≈80 s, the 

hold time was ≈920 s for step (a), the temperature-decrease duration was ≈40 s, and the lowertemperature hold time was ≈1960 s to give 3000-s total test time from ramp initiation at 300oC to the end 

of the heating phase. The long-time step (b) temperatures were stable enough that the benchmark was 

terminated 1000 s earlier than what was specified for Transient #1. Data were recorded every 0.1 s. Due 

to the 32000 data point limit of EXCEL, the data were filtered at the end of the heating phase such that 

data are plotted every 50 s from 3200 s to 3500 s (see thinner lines in Fig. 52). 

Prior to conducting Test ZLU#139, the furnace power was adjusted slightly to give long-time hold 

temperatures of 1045°C and 980°C and the total test time was extended to 4000 s. As Test ZLU#139 was 

a data-generating test, it was conducted without TCs welded directly onto the sample. Figure 53 shows 

the temperature history for the control TC welded to the sample holder. The thinner part of the 

temperature trace represents data points recorded every 50 s.

Test ZLU#140 was conducted with a temperature history corresponding to Transient #2 in Fig. 51. This 

was also a single step-change test: (a) hold at 975°C for 1500 s, (b) increase temperature to 1000°C, and 

(c) hold for 2500 s at 1000°C. In order to achieve long-time hold temperatures of 975oC and 1000oC, the 

furnace power was adjusted based on the benchmark relationships established among the furnace power, 

the holder control TC, and the sample TC. The control TC temperature history for Test ZLU#140 is 

shown in Fig. 54. As before the thinner parts of the temperature trace represent data plotted every 50 s.
96 

Figure 51. Schematics of transient temperature histories for breakaway oxidation tests with 

ZIRLO cladding samples.

Temperature Profile for Transient #1

0

200

400

600

800

1000

1200

0 1000 2000 3000 4000 5000

Time (s)

Temperature (°C)

1045°C

980°C

300°C

Temperature Profile for Transient #2

0

200

400

600

800

1000

1200

0 1000 2000 3000 4000 5000

Time (s)

Temperature (°C)

975°C 1000°C

300°C

Temperature Profile for Transient #3

0

200

400

600

800

1000

1200

0 1000 2000 3000 4000 5000

Time (s)

Temperature (°C)

980°C

1030°C

930°C

300°C
97 

Figure 52. Thermal benchmark results for Test ZLU#138 with as-fabricated ZIRLO-2008. 

Figure 53. Temperature history of the control TC for Test ZLU#139.

300

400

500

600

700

800

900

1000

1100

500 1000 1500 2000 2500 3000 3500

Temperature (°C)

Time (s)

TC4 on Sample

TC5 on Sample

300

400

500

600

700

800

900

1000

1100

500 1500 2500 3500 4500

Temperature (°C)

Time (s)

Control TC
98 

Figure 54. Temperature history of the control TC for Test ZLU#140.

Two thermal-cycling tests were performed for Transient #3. Test ZLU#142 was run with ZIRLO cladding 

under the following conditions: (a) 980°C for 1500 s, (b) 5 cycles from 930°C to 1030°C for about 400 s, 

and (c) 980°C for 1400 s. For ZLU#143, the temperature profile was modified slightly: the initial hold 

time at 980°C was increased from 1500 s to 2000 s, the temperature was cycled during the 2000-2400 s 

time interval, and the sample was held at 980°C during the remaining 2400-2800 s time interval. The 

control TC temperature histories for Tests ZLU#142 and ZLU#143 are shown in Figs. 55 and 56, 

respectively.

4.4.2 Outer-surface appearance of transient oxidation samples

The outer surfaces of the ZLU#139 and ZLU#140 samples were lustrous black (see Fig. 57a). There were 

no indications of breakaway behavior after 4000-s total oxidation time for these samples subjected to 

temperature decrease (from 1045°C to 980°C) and increase (975°C to 1000°C). 

Outer surface discolorations were observed for both samples subjected to temperature cycling (see Figs. 

57c and 57d). The ZLU#142 sample had small, but significant, yellow-gray areas on the outer surface and 

extensive (middle 75%) yellow-gray areas on the inner-surface after 3300-s total test time. The region just 

to the right of the ZLU#142 sample mid-plane had more discoloration than the mid-plane due to yellowgray areas around the back of the sample, which are not visible in Fig. 57c. The reverse was true for the 

ZLU#143 sample after 2800-s total test time: extensive yellow-gray area on the outer surface and yellowgray spots on the inner surface. 

300

400

500

600

700

800

900

1000

1100

500 1500 2500 3500 4500

Temperature (°C)

Time (s)

Control TC
99 

Figure 55. Temperature history of control TC for Test ZLU#142. 1400-s hold time following 

cycling. 

Figure 56. Temperature history of control TC for Test ZLU#143. 400-s hold time following 

cycling. 

300

400

500

600

700

800

900

1000

1100

1200

500 1000 1500 2000 2500 3000 3500

Temperature (°C)

Time (s)

Control TC

300

400

500

600

700

800

900

1000

1100

1200

500 900 1300 1700 2100 2500 2900 3300

Temperature (°C)

Time (s)

ControlTC
100 

Figure 57. Appearance of as-received ZIRLO cladding outer-surface after steam oxidation 

tests (a) ZLU#139 for Transient #1, (b) ZLU#140 for Transient #2, (c) ZLU#142 for 

Transient #3, and (d) ZLU#143 for a modified Transient #3. Photos are in gray 

scale. 

4.4.3 Weight gain and hydrogen pickup

As indicated in Table 27, the measured weight gains were less than the CP-predicted weight gains for all 

four samples oxidized under transient conditions. This is consistent with weight-gain results from 

isothermal oxidation tests conducted at 970 to 1000°C for long times. The hydrogen contents of ZLU#139 

and ZLU#130 test samples were not measured because both outer and inner surfaces were lustrous black.

Breakaway oxidation for both temperature-cycled samples was confirmed by outer-surface discoloration 

and hydrogen pickup (see Table 27): 427±36 wppm after 3300-s total test time for the ZLU#142 sample 

and 225±79 wppm after 2800-s total test time for the ZLU#143 sample. However, some of the hydrogen 

in the ZLU#142 sample came from inner-surface breakaway oxidation. Previous results obtained for 

ZIRLO-2003 oxidized at 1000oC [1] indicated that hydrogen pickup from inner-surface breakaway was 

lower than hydrogen pickup from the outer surface: 30 wppm after 2440 s and 100 wppm after 3480 s 

even though the inner surface was in full breakaway (see Table 29 and Fig. 55b in Ref. 1). Therefore, 

most of the hydrogen pickup for the ZLU#142 sample was from outer-surface breakaway. 

4.4.4 Discussion of transient oxidation results

In general, breakaway oxidation of zirconium-based alloys is associated with local formation of 

monoclinic oxide at the metal-oxide interface during oxidation in steam. It has been reported that the 

monoclinic-to-tetragonal and tetragonal-to-monoclinic transformations do not occur at a fixed 

temperature [31]. The ranges of the transformation vary for both the forward and the reverse 

transformations, depending on impurities (type and concentration), grain size, stoichiometry, and thermal 

history. In addition, the transformation temperature can be reduced as the compressive stress increases. 
101

Table 27. Results of ZIRLO-2008 transient breakaway-oxidation tests. 

Test ID

Transient

Type

Total Test

Timea

, 

s 

CP-predicted

wg,

mg/cm2

Measured

wg, 

mg/cm2

Measured

(CH)POX, 

wppm

ΔCHM

b

, 

wppm 

ZLU#139 1 4000 14.3 9.5 --- --- Lu

ZLU#140 2 4000 13.5 8.1 --- --- Lu

ZLU#142 3 3300 11.6 8.5 425±35 427±36 Small 

extens

ZLU#143 3 2800 10.8 8.2 232±76 225±79 Extensi

smal

a

Total time from ramp initiation at 300oC to end of hold temperature; ramp time is ≈80 s.

b

ΔCHM is referenced to as-fabricated sample mass: ΔCHM = (1 + 5.4×10-3 wg) (CH)POX - CHi, where CHi is hydrogen content i
102 

Although it is not clear at which temperature the oxide phase transformation takes place for the 

ZIRLO alloy, previous ANL breakaway oxidation results for ZIRLO cladding subjected to 

isothermal oxidation indicate minimum breakaway oxidation times occur within the range of 

3100±300 s for isothermal oxidation temperatures of 970-985oC. At a slightly higher oxidation 

temperature (1000oC), the minimum breakaway oxidation time for ZIRLO was longer (4000±200 

s). We also found that a rapid temperature rise with ≈60oC overshoot for a very short time appeared 

to have some effect on long-time breakaway. However, no conclusions could be drawn because the 

hydrogen-pickup results were within data scatter for long-time temperatures of 970-985oC. 

Therefore, it was of interest to determine if these step-change transients (i.e., Transients #1 and #2)

could lead to shorter breakaway times. As indicated in Table 27, step changes from higher 

(1045oC) to lower (980oC) and from lower (975oC) to higher (1000oC) temperatures did not result 

in shorter breakaway oxidation times (>4000 s total test time).

The monoclinic-to-tetragonal transformation is accompanied by a sudden volume change and by 

anisotropic thermal expansion behavior of the lattice parameter, which disrupt the oxide. The 

purpose of the cycling tests under Transient #3 conditions was to determine whether localized 

stresses induced by the phase transformation and differential thermal expansion could disrupt the 

tetragonal layer near the metal-oxide interface, which in turn might reduce the breakaway time near 

the critical breakaway temperature. Test results listed in Table 27 indicate a marginal reduction 

(≈300 s) in breakaway time due to temperature cycling: 225±79 wppm ΔCHM after 2800 s. 

Additional characterization of this sample from Test ZLU#143, along with the ZLU#142 samples, 

is presented in the following.

Figure 58 shows the yellow breakaway regions on the ZLU#143 sample outer surface. Some 

grayish/white areas were also observed, and some of these were in the center of the yellow areas. 

Since the oxide color is strongly influenced by stoichiometry, as well as impurities, this change 

might indicate an increase in the O-to-Zr ratio. A high-magnification image of a small yellow area 

(Fig. 58b) shows the morphology of the monoclinic phase on the ZLU#143 sample outer surface. 

The center area appears grayish/white surrounded by yellow areas. Figure 59 shows the transition 

area from black to yellow in the region marked “T” in Fig. 58a. The cladding longitudinal axis is 

delineated by an arrow in Figs. 58a and 59a. It is interesting to observe that the yellow areas have a 

rougher outer surface than the black areas. This feature was also observed within the small area of 

yellow, monoclinic oxide shown in Fig. 58b. 
103 

Figure 58. (a) ZIRLO sample after steam-oxidation at Transient #3 conditions for 2800 s 

total test time with cycling from 930°C to 1030°C (Test ZLU#143), and (b) 

high-magnification.

For the temperature-cycled ZLU#142 sample oxidized for longer time (3300 s), there was 

extensive discoloration within the middle 75% of the sample inner surface, and the additional 500 s 

allowed for more axial and circumferential diffusion of hydrogen (see Fig. 59). As shown in Fig. 

60, rings were sectioned from an axial region that was ≈3 to 11 mm from the center of the sample. 

Hydrogen pickup values were determined to be 427±36 wppm, 312±31 wppm, 221±41 wppm, and 

153±38 wppm, with the highest value for a thin ring closest to the center of the sample. The highest 

ΔCHM value corresponded to an area with yellow-gray oxide over the whole inner surface and with 

yellow-gray oxide over about 40% of the outer surface. The lowest ΔCHM value was for a ring near 

the end with lustrous black oxide on both inner and outer surfaces. Given the relatively small 

circumferential variation and the relatively large hydrogen content in the ring with black inner- and 

outer-surface oxides, it is clear that significant hydrogen diffusion occurred during the additional 

500 s test time. 
104

Figure 59. High-magnification images of Area T in Fig. 58a showing different outer-surface 

areas of the ZLU#143 sample: (a) lustrous black area, (b) transition area from 

black-to-yellow, and (c) yellow area. 

(a) 

(b)

Figure 60. Axial distribution of hydrogen pickup for the ZLU#142 test sample. 312±31 wppm 221±41 wppm 153±38 wppm End
105

The 2800-s breakaway time for temperature-cycled ZIRLO is comparable to the breakaway time for a 

pre-scratched ZIRLO sample oxidized under isothermal conditions at 970°C [1]. In the earlier testing, a 

20-μm-deep, rectangular-shaped scratch was machined along the length of the outer surface. After 

oxidation, a yellow region was observed along the middle third of the scratch. The average hydrogen 

pickup at the sample mid span was 44 wppm. However, the local hydrogen pickup under the yellow 

scratched region was 440 wppm. Based on the large hydrogen-pickup data set for smooth cladding, it was 

estimated that the sample would have picked up >200-wppm hydrogen by ≈2800 s. Thus, the effects of 

temperature cycling on breakaway oxidation time appear to be comparable to the effects of a machined 

surface scratch. Each of these conditions reduced the average breakaway time by about 300 s, which is at 

the lower end of the data scatter.

The total data set for oxidation under isothermal and transient conditions, with long-time hold 

temperatures in the range of 970 to 985°C, is shown in Fig. 61. As can be seen, this larger data base 

confirms our earlier conclusion that breakaway oxidation of ZIRLO occurs at 3100±300 s (as determined 

by pickup of ≈200 wppm hydrogen) in the temperature range of 970 to 985oC. 

Figure 61. Hydrogen-pickup vs. time data for ZIRLO oxidized in steam at 970−985°C: smooth 

samples (isothermal), pre-scratched (≈20-μm depth) samples (isothermal), preoxidized samples (isothermal) and smooth samples (transient) oxidized with five 

cycles from 930 to 1030°C for 400 s and 980°C isothermal temperature.

0

200

400

600

800

1000

1200

2400 2600 2800 3000 3200 3400 3600 3800 4000

Hydrogen Pickup (wppm)

Oxidation Time (s)

Isothermal: Smooth

Isothermal: Scratched

Transient: Smooth

Linear Fit

 

107 

5 BEHAVIOR OF 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 [1]. Consistent with current licensing 

requirements in 10 CFR 50.46, all CP-ECR values reported in this section are for the rupture-region cross 

section with the minimum cross-section averaged wall thickness prior to oxidation.

Two 4-point bend tests (4-PBTs) were conducted previously at RT on as-fabricated Zry-2 specimens after 

LOCA integral testing [1]. Inferences were made from those results, but not enough tests were conducted 

to determine quantitative trends and effects of 4-PBT temperature. More recently, a large number of postrupture 4-PBTs were performed at 135°C on ZIRLO specimens to evaluate results quantitatively.

LOCA integral tests were conducted with as-fabricated (AF) and pre-hydrided (PH) ZIRLO-2010 

cladding samples. Methods for conducting these tests, along with sample geometry, are described in 

Section 2.5 and in Ref. 1. Following LOCA oxidation, 4-PBTs were conducted to determine the 

maximum bending moment (measure of strength), the applied energy prior to failure (measure of 

toughness), and the offset displacement (measure of plastic deformation). It was anticipated that all three 

of these metrics would decrease with increasing oxidation level and pre-test hydrogen content. Of 

particular interest was the determination of whether or not the embrittlement limits established for nondeformed cladding (see Fig. 25) would be sufficient to protect the ballooned and ruptured region from 

excessive degradation. The results of this investigation were documented in Ref. 5 for AF ZIRLO 

cladding. However, detailed results from LOCA integral tests (e.g., temperature and pressure histories) 

and 4-PBTs (e.g., load-displacement curves) were not documented in Ref. 5. This section presents a 

comprehensive documentation of test results for AF and PH ZIRLO cladding samples.

The testing sequence consisted of thermal and metallurgical benchmark tests, scoping tests generally 

conducted without quench and with variable internal gas pressures and oxidation test times, and datageneration tests. Relevant raw data are presented in the Appendices: (A) Temperature and Pressure 

Histories, (B) Axial Profiles of Post-Rupture Cladding Strains, (C) Photographs of Post-Oxidation LOCA 

Integral Samples, (D) 4-PBT Load-Displacement Curves, and (E) Photographs of Post-Bend LOCA 

Integral Samples.

Symbols used to describe measured and calculated geometrical parameters are defined as they are used 

throughout Sections 5.1−5.3. As a convenience to the reader, these symbols are also defined here with 

some of the more complicated ones illustrated in the Fig. 62 schematic. Dimensions of the AF cladding 

are used as reference pre-test parameters. Nominal dimensions for this cladding are indicated by the 

subscript “i” for initial condition: 9.50-mm outer diameter (Doi), 0.57-mm wall thickness (hi), 29.85-mm 

outer-wall circumference (Coi), 28.05-mm mid-wall circumference (Cmi), and 16.0-mm2 cross-sectional 

area (Ai). With the exception of Ai, these parameters change significantly during ballooning and rupture 

(R). Routinely measured and/or calculated parameters for the rupture cross section with the largest 

rupture strain are the (see Fig. 62): outer-surface diameter (DmaxR) measured from the rupture tips to the 

back of the rupture cross section, width (δR) of the rupture opening, outer-surface diameter (DminR) 

measured at an orientation perpendicular to DmaxR, outer-surface circumference (CoR) excluding δR, and 

mid-wall circumference (CmR) excluding δR. The parameter DavgR, which is used in correlations, is simply 

the average of DmaxR and DminR. 
 

108 

Figure 62. Schematic of ballooned and ruptured cladding, which includes the parameters 

routinely measured and/or calculated following a LOCA integral test. The sketch is 

not to scale.

5.1 Thermal and Metallurgical Benchmark Test Results

Prior to conducting tests with ZIRLO cladding, four tests were run with zirconia-filled Inconel tubing to 

optimize temperature control parameters for ramp rate and hold temperature, to examine the effects of 

sample axial location within the furnace, and to examine the effects of temperature ramp rate on axial and 

circumferential temperature variations during the heating ramp and the hold time. These tests were 

conducted with temperature ramp rates of 5°C/s and 10°C/s and hold temperatures of 1200°C and 

1050°C. Multiple tests were conducted with a single Inconel tube because of its relatively low oxidation 

rate in steam. The TC positions and orientations were the same as used in the testing previously reported 

[1]. Relative to the sample mid-span, TCs were positioned at 0° orientation and axial locations of 51 mm, 

0 mm, and -51 mm. In addition, a 4th TC was located at the mid-span (0 mm) at the 180° orientation. The 

reference heating rate was chosen to be 5°C/s because better temperature uniformity was achieved at this 

heating rate. The reference cooling rate was maintained at 3°C/s between the end of the hold time and 

800°C. Also, because of the need for data at or near the licensing limit, 1200°C was chosen as the 

reference hold temperature. Given these test parameters, it was possible to conduct ramp-to-rupture tests 

at ≈0% CP-ECR and oxidation tests with ballooned-and-ruptured cladding at >10% CP-ECR. For 

ballooned-and-ruptured ZIRLO cladding with reduced wall thickness prior to oxidation, >10% CP-ECR 

was realized for temperature histories with essentially zero hold time because of oxidation during the 

heat-up and cool-down phases of the tests.
 

109 

The 1st benchmark test with ZIRLO cladding (OCZL#5) was conducted without internal pressurization 

and without quench. The temperature history for this test is shown in Fig. A.1. During the temperature 

ramp, the variations in TC readings were quite small at average temperatures of about 700°C, 800°C, and 

900°C. The programmed hold time was 60 s for this test. However, the temperature controller is 

programmed to slow down the ramp rate at about 1180°C as 1200°C is approached in order to avoid 

overshoot. On the average, the target temperature is reached after about 20 s to 25 s at the slow ramp rate. 

The average hold temperature was 1206±9°C about half way through the programmed 60-s hold time. 

These temperature values are well within acceptable ranges for LOCA integral tests.

The metallurgical part of the benchmarking consisted of measuring the outer-surface oxide layer 

thickness values at two axial locations and comparing them to the CP-predicted values for the 

corresponding temperature histories. On the basis of three measurements at each of eight circumferential 

locations at the mid-span (0 mm) and at 51 mm above the mid-span (location of control TC), the 

corresponding oxide layer thickness values were 41±2 μm and 40±3 μm, respectively, where the ± values 

represent one standard deviation in the data. The CP-predicted values for oxide layer thickness values at 

these locations were 39 μm and 42 μm for the two mid-span TC thermal histories and 38 μm for the TC 

history at 51 mm above the mid-span. The average of the 16 (times 3) data points was 40±2 μm. The 

average CP-predicted value for the four TC readings was also 40±2 μm. The oxide-layer-thickness data 

have an uncertainty of about ±1 μm. Thus, the agreement between measured and calculated values for 

oxide layer thickness was excellent, which confirmed the accuracy of the temperature measurements.

The 2nd thermal-metallurgical benchmark test (OCZL#6) was conducted with AF ZIRLO pre-pressurized 

to 8.28 MPa (1200 psig) and subjected to a 60-s hold time at a target hold temperature of 1200°C. It was 

anticipated that temperature uniformity would be good prior to ballooning initiation and that temperature 

variation would be more substantial after rupture. Ballooning and rupture cause a significant change in 

geometry, and they challenge the integrity of welded TCs. The temperature-pressure histories are shown 

in Fig. A.2, and a post-oxidation-test photograph is shown in Fig. C.1. Just prior to rupture, the mid-plane 

TC at 180° began to read higher than the other TCs. Figure C.1 shows that the rupture occurred at a 

location between the mid-span and the lower TC at -51 mm. The average rupture temperature was 

750±30°C, with most of the variation due to the high reading of the 0-mm/180° TC. Half way through the 

hold time, there was excellent temperature uniformity for the three TCs at 0° orientation (1204±7°C). 

However, the corresponding 0-mm/180° TC reading was unacceptably high (1276°C). The average 

temperature for the effective hold time (82 s) was 1219±33°C. When one TC behaves significantly 

different from the other TCs, there is a distinct possibility that the measurement is unreliable. 

Measurements of outer-surface oxide-layer thickness at the mid-span cross section were used to confirm 

that all four TC readings were reliable. Oxide-layer thickness values varied from 50 μm at the orientation 

of the 0-mm/180° TC to 36 μm at the orientation of the 0-mm/0° TC. Corresponding CP-predicted values 

were 52 μm and 39 μm. Although the 0-mm/180° TC reading was unacceptably high from a temperature 

uniformity viewpoint, the accuracy of the temperature readings was confirmed.

Axial profiles of cladding diametral strain at two orientations are shown in Fig. B.1 for the post-oxidation 

OCZL#6 test sample. The designated orientations of 0° and 90° differ from TC orientations and are 

defined as follows: 0° corresponds to the caliper-measured diameter (Dmax) along the length of the sample 

at the orientation defined by the line from the rupture-opening mid-span to the back of the cross section 

containing the widest rupture opening (δR) and 90° corresponds to the caliper-measured diameter (Dmin) at 

an orientation perpendicular to Dmax. Peak values within the cross section containing δR are labeled as 

DmaxR and DminR. The average of these diameters, which is used in Section 5.3, is DavgR. In Fig. B.1, as 

well as in subsequent Appendix B figures, the differences (ΔDmax and ΔDmin) between the post- and pretest outer diameters were normalized to the pre-test outer diameter (Doi) to give the plotted diametral 
 

110 

strains (ΔDmax/Doi and ΔDmin/Doi). The diametral strains shown in Fig. B.1 are useful for identifying the 

presence of a secondary balloon at an axial location between the primary balloon and the top of the 

sample and for identifying a tertiary balloon near the top of the sample (i.e., gas-plenum region). 

The presence of secondary and tertiary balloons observed in Fig. B.1 was undesirable, but it was decided 

to proceed with scoping tests to determine if these were trends or abnormalities. After completion of 

scoping and data-generation tests with AF ZIRLO cladding, additional benchmark tests (OCZL#36 and 

OCZL#37) were conducted with AF ZIRLO prior to testing PH ZIRLO. The temperature and pressure 

histories for these tests are shown in Figs. A.22 and A.23. Test OCZL#36 was essentially a repeat of 

OCZL#6, except that the sample length was increased from 305 mm to 318 mm to move the plenum 

region outside the heating zone of the furnace and the axial and circumferential positions for the TCs 

were modified. These modifications followed the scoping tests and are discussed in Section 5.2. Thermal 

benchmark results for test OCZL#37 are presented in the following paragraph. 

As indicated in Fig. A.23, the longer hold-time test (150-s hold, 172-s effective hold) gave an average 

hold temperature of 1211±16°C. Although no TC was located at the sample mid-span, metallographic 

examination was performed to determine the outer-surface oxide layer thickness at this location. On the 

basis of images at 10 circumferential locations, the oxide layer thickness was measured to be 52±7 μm. 

The CP-predicted oxide layer thicknesses for the TC measurements above and below the mid-span were 

51±3 μm. This good agreement justified the subsequent (for data-generation tests) relocation of the two 

TCs near the sample mid-span to locations above and below the mid-span with angular orientations of 0°

(upper), 90° (lower), 180° (upper control TC), and 270° (lower). Also (see Fig. B.20), subsequent 

movement of the plenum region to a location above the furnace heating zone eliminated the tertiary 

balloon observed in the OCZL#6 sample. 

5.2 Scoping Test Results

Seven scoping tests were conducted to determine the effects of internal gas pressure on the: (a) rupture 

strain (ΔCmR/Cmi) and temperature (TR), (b) minimum wall thickness (hR) for CP-ECR calculations, (c) 

TC durability and reliability, (d) temperatures (Th and The) during programmed (th) and effective (the) hold 

times, and (e) 4-PBT performance. Four ramp-to-rupture tests, for which the furnace was shut down after 

rupture, were conducted. Three additional tests were heated to ≈1200°C for 60 s to 70 s, cooled at 3°C/s 

to about 800°C, and further cooled without quench. Table 28 summarizes test conditions and results for 

the scoping tests.

Ramp-to-rupture tests

Ramp-to-rupture tests were conducted with 300°C fill pressures of: 400 psig, 600 psig, 1400 psig, and 

1600 psig. The determination of maximum rupture strains (ΔCmR/Cmi) and corresponding hR is 

documented in Section 5.3. Data trends are comparable to what has been observed for AF Zry-4 [32]: (a) 

high rupture strains at 715°C (likely close to the temperature at which the peak strain is observed to occur 

in the alpha phase), (b) low rupture strains at 845°C (close to the mixed alpha-beta phase temperature at 

which the minimum strain is observed to occur) and (c) increasing rupture strains at >845°C. 

Three ramp-to-rupture samples were subjected to 4-PBTs. None of the samples cracked during the 14-mm 

displacement, indicating high values for strength, absorbed energy, and plastic deformation for nonoxidized AF ZIRLO. The degree of out-of-plane bending and secondary-tertiary ballooning did vary from 

sample to sample following the ramp-to-rupture tests. For test OCZL#11 pressurized to 1400 psig, 
111

Table 28. Summary of results for LOCA integral and post-LOCA-bend scoping tests with AF ZIRLO cladding. Fil

300°C. Oxidation tests were conducted with a target hold temperature of 1200°C and without quench. 4-P

at 135°C and 2 mm/s displacement rate to 14-mm maximum displacement. Failure locations are referenc

the rupture opening.

Test

ID

OCZL#

Fill

Pressure,

psig (MPa)

ΔCmR/Cmi,% 

(TR, °C) hR, mm

CP-ECR, % 

(The, °C)

Stress in

Rupture 

Node 

Failure

Location 

Max.

Bending 

Moment, 

N•m

Max

Energ

J 

7 800 

(5.52) 

22

(810±30) 

0.47 15.7 

(1226±28) 

─ ─ ─ ─ 

8 600 

(4.18) 

22 

(845±25) 

0.47 0 

(─) 

Maximum 

tension 

No

cracking

20.9 >8.4 

9 400 

(2.76) 

33 

(875±15) 

0.43 0 

(─) 

Maximum 

tension 

No

cracking

20.6 >8.3 

10 1600 

(11.0) 

70 

(715±10) 

0.33 0 

(─) 

Maximum 

tension 

No

cracking

19.7 >7.7 

11 1400 

(9.66) 

40

(750) 

0.41 0 

(─) 

─ ─ ─ ─ 

12 1000 

(6.90) 

33 

(805±20) 

0.43 14.5 

(1191±17) 

Maximum 

compression

-40 mm

+33 mm

10.5 0.78 

13 1200 

(8.28) 

42 

(741±15) 

0.40 14.4 

(1162±21) 

Maximum 

tension 

0 mm 8.8 0.58 
112 

Figures B.6 (profilometry) and C.6 (photographs) show a pronounced secondary balloon that occurred 

near the top of the sample with a strain comparable to that of the primary balloon. A smaller tertiary 

balloon occurred below the primary balloon. Figure C.6a also shows the out-of-plane bending of the 

sample, which would cause alignment issues during the 4-PBT. On the basis of the limited results for 

ramp-to-rupture tests, 600-psig fill pressure at 300°C was chosen as the reference pressure for tests 

designed to achieve oxidation levels of 10% to 12% CP-ECR because of the relatively high hR value for 

this pressure.

Metallographic examination was performed on the OCZL#11 sample to determine the rupture node crosssection shape, variation in cladding thickness around the circumference, and circumferential strain at the 

cladding mid-plane. Figure 63 shows the cross section at the rupture mid-span location. The 

circumference (CmR) around the cladding mid-wall was 39.3 mm. Relative to the mid-wall circumference 

(Cmi = 28.05 mm) for AF ZIRLO, the rupture strain (ΔCmR/Cmi = [CmR – Cmi]/Cmi) was 40%. 

Figure 63. Image of ramp-to-rupture test OCZL#11 sample cross section through rupturenode mid-span. Rupture width (δR) was 6.2-mm and ΔCmR/Cmi was 40%.
113 

Scoping oxidation test results 

Three oxidation tests were conducted at 300°C fill pressures of 800 psig (OCZL#7), 1000 psig 

(OCZL#12) and 1200 psig (OCZL#13) Temperature and pressure histories are presented in Figs. A.3, 

A.8, and A.9, respectively. For the OCZL#7 test sample, three of the four TCs survived ballooning and 

rupture. Half way through the 60-s hold time, the average of the mid-span and top TC readings, both at 0° 

orientation, was 1211±13°C. However, the temperature recorded by the mid-span TC at 180° (close to the 

edge of the rupture opening) was considerably higher (1260°C). The profilometry results shown in Fig. 

B.2 indicated a relatively small secondary balloon near the top of the sample. Low-magnification 

metallographic examination was performed to determine the rupture cross-section shape and the cladding 

mid-wall circumference at the mid-span of the rupture node. Figure 64 shows the rupture cross-section, 

along with the variation in metal wall thickness from very thin at the rupture tips to relatively thick at the 

back of the rupture node. The shape of the OCZL#7 sample is similar to the shape of the OCZL#11 

sample shape shown in Fig. 63. However, the rupture strain (22%) for the OCZL#7 sample was lower. 

Figure 64. Low magnification image of the cross section through rupture-opening mid-span for 

the OCZL#7 sample with 22% rupture strain prior to 16% CP-ECR oxidation. 

For the OCZL#7 sample, as well as the OCZL#11 sample, the cladding wall thickness varied from very 

small at the rupture tips to a maximum value near the back (top of figure) of the rupture node. In the Fig. 

64 orientation, the thin rupture tips would be subjected to maximum axial tensile stresses in response to 4-

PBT loading while the thicker back region of the cross section would be subjected to maximum 

compressive stresses. This reference orientation is conducive to crack initiation at the brittle rupture tips 

and rapid crack growth through at least half of the cross section. Whether or not the crack would blunt in 

the thicker “more ductile” back half of the cladding would depend on the fracture toughness of the 

material.

For the OCZL#12 test, all four TCs behaved well up until rupture. The temperature readings for the midspan TC at 0° were unreliable after rupture. It is surprising that the control TC (51 mm/0°) behaved as 

well as it did. Figure C.7 shows that the top of the rupture tip was very close to this TC and at the same
114 

orientation. Better temperature uniformity was achieved during the hold time for the three remaining TCs. 

Half way through the 60-s hold time, the average hold temperature was 1196±16°C, which is acceptable. 

The CP-ECR was about 15% (14.5%). However, as shown in Fig. B.7, there was a pronounced secondary 

balloon about 51 mm below the mid-span, which perhaps explains the failure of the TC at this location. 

The average diametral strain (23%) at this location was large compared to the average diametral strain 

(39%) at the rupture mid-span. Also, a tertiary balloon was observed near the top of the sample (close to 

the plenum region). 

Following oxidation and cooling, the OCZL#12 sample was subjected to a “reverse” 4-PBT at 135°C 

with the maximum axial tensile stress occurring at the back of the rupture cross section and with the 

rupture tips in axial compression. The sample failed at two locations between the rupture edges and the 

balloon necks (see Fig. E.2) at relatively low failure energy and offset displacement compared to the 

ramp-to-rupture samples. The bending moment just prior to failure was about half of what was achieved 

by the ramp-to-rupture samples prior to test termination. Additional post-bend characterization was 

performed to determine the axial profile of hydrogen content (also known as secondary hydriding). Figure 

65 shows the measured hydrogen content and CP-ECR at the rupture mid-span, as well as the hydrogen 

contents and measured ECRs (based on metallographic examinations) at the two failure locations. The 

hydrogen content at the rupture mid-span was only 25 wppm, as compared to 22 wppm for AF ZIRLO. 

As anticipated, the hydrogen pickup at the rupture mid-span was insignificant. However, the hydrogen 

content was 1700 wppm at the two failure locations. Figure 66 shows the hydrogen concentration vs. 

axial location relative to the rupture mid-span location, along with the hydrogen peaks (≈2400 wppm) 

near the balloon necks. 

Figure 65. Hydrogen content, ballooning strain, and oxidation level at rupture mid-span (CPECR) and failure locations (measured ECR) for the OCZL#12 sample after the 

reverse 4-PBT.

For the OCZL#13 test, the sample was moved 13 mm higher in the furnace, the 300°C internal pressure 

was increased to 1200 psig and the hold time was increased to 70 s. The goals were to generate a 4-PBT 

sample exposed to oxidation temperatures closer to 1200°C and had less secondary ballooning. Previous 

experience with TC location and orientation suggested that circumferential variations in temperature were 

greater than axial variations. Thus, TCs were oriented at 4 circumferential orientations and positioned at 

±51 mm from the center of the sample: 0° and 180° (both upper) and 90° and 270° (both lower). Figure 

A.9 shows that all four TCs responded reliably throughout the ramp, hold time, and cool down. The 

average temperature of the control TC was 1200°C during the 70-s hold time. However, the average 

temperature of the other three TCs was considerably lower (1150±8°C) during the same time period, such 

that the overall average was only 1162±21°C during the 92-s effective hold time. These results suggested 
115 

Figure 66. Axial profile of hydrogen content for the OCZL#12 sample following oxidation to 

14.5% CP-ECR.

that additional refinements in axial location within the furnace were required before conducting datageneration tests. Some improvement was observed in the diametral strain profile. Movement of the 

plenum region about 13 mm higher appeared to eliminate secondary and tertiary ballooning near the top 

of the sample (see Fig. B.8). Another improvement was that no out-of-plane bending was observed, which 

is desirable for the 4-PBTs. Table 28 summarizes the results for the 4-PBT. The 4-PBT sample failed 

abruptly (see load-displacement curve in Fig. D.5) at the rupture mid-span (see Fig. E.3) with little 

evidence of offset displacement (0.13% measured compared to ≥0.1% criterion for plastic deformation).

Results from the scoping tests were used to refine sample length, axial position of the sample within the 

furnace, temperature control parameters, and TC locations. With respect to sample length, this was 

increased from 305 mm to 318 mm in order to relocate the short plenum region above the uniform heating 

zone of the furnace. The benchmark and scoping tests were also used to establish reference 300°C internal 

pressures of 600 psig and 1200 psig for the data-generation tests.

5.3 Determination of Rupture Strains and Post-Rupture Cladding Wall 

Thickness

For ballooned and ruptured cladding, the CP-ECR is based on the predicted average wall thickness (hR) 

prior to oxidation for the cross section containing the rupture node. The instructions given in 10 CFR 

50.46 for calculating hR are: “...the unoxidized cladding thickness shall be defined as the cladding crosssectional area, taken at a horizontal plane at the elevation of rupture...divided by the average 

circumference at that elevation. The average circumference does not include the rupture opening.” 

It is common to report rupture strain as the maximum post-test outer-surface circumference (CoR) of nonoxidized cladding compared with the pre-test outer-surface circumference (Coi): 

ΔCoR/Coi = (CoR – Coi)/Coi (18) 
116 

To be consistent with 10 CFR 50.46, as well as to improve the accuracy in hR determination, the mid-wall 

circumferential strain is used in the current work to calculate hR. However, the measurement of crosssectional area would be tedious for the cross sections shown in Figs. 63 and 64. A good approximation for 

determining hR for CP-ECR calculations is:

hR = hi/(CmR/Cmi) = hi/(1 + ΔCmR/Cmi) (19) 

where hi is the pre-test wall thickness (0.57 mm) and the subscript “m” indicates mid-wall values.

Equation 19 is consistent with observations from pressurized-tube tests for isotropic materials: (a) the 

axial plastic strain is negligible and (b) the cross-sectional area of the expanded tube is constant and equal 

to the pre-test cross-sectional area. For the AF ZIRLO cladding used in this study, the cross-sectional area 

(Ai) was 16.0 mm2 and the outer and mid-wall circumferences were 29.85 mm and 28.05 mm, 

respectively. The product of Cmi and hi agrees quite well with Ai, while Coi × hi over-predicts Ai by 6%. 

Although the expansion of the cross section with no change in area thins the wall and results in better 

agreement between CoR and CmR (e.g., 2% difference at 60% strain), we will use the more accurate midwall values. 

The initial approach used in the current work was to develop a correlation for post-rupture circumference 

(i.e., CmR) as a function of easily measureable parameters at the rupture mid-span: DminR, DmaxR, DavgR, and 

δR. During efforts to develop such a correlation, Argonne modeled the rupture-region cross section with 

two basic shapes with the same cross-sectional area: a partial circle and a partial ellipse with the arc 

length of the truncated portion associated with a cord of length δR subtracted from the full circle or 

ellipse. Neither approach gave good agreement with CmR data determined from low-magnification images 

(e.g., see Figs. 63 and 64) or determined from direct measurements of CoR. This is partly due to the wide 

range of rupture cross-section shapes and rupture-opening widths. Thus, it was determined that an 

empirical approach for determining CmR from measured parameters would yield better results than a semianalytical approach.

In addition to shapes shown in Figs. 63 and 64, Fig. 67 shows the post-oxidation, rupture mid-span cross 

section for data-generation test OCZL#18, which had a measured mid-wall circumference of 40 mm (43% 

rupture strain) and a 5.4-mm rupture opening. Data from these images were very useful in developing the 

empirical correlation because of the ranges of CmR (34−40 mm) and δR (3.7−6.2 mm) values.

Concurrent with the LOCA integral testing at Argonne, Studsvik conducted similar tests with AF ZIRLO 

[33] and ZIRLO from fuel rods irradiated to high burnup [34]. Metallographic results for Tests 185 (AF 

ZIRLO) and 192 (irradiated ZIRLO with ≈230 wppm hydrogen) were analyzed by Argonne [35]. Rupture 

cross-section shapes for these two tests are shown in Figs. 68 and 69. The addition of the Studsvik data 

for correlation development expanded the databases for CmR and δR to (34.1−43.5 mm) and (3.0−7.8 mm), 

respectively.

In addition to five mid-wall circumferences measured from low-magnification images, Argonne measured 

the post-oxidation CoR values for seven AF and six PH ZIRLO test samples. The measurements were 

performed with a relatively stiff, thin string. The mid-wall circumference was calculated by solving a 

quadratic equation that included the subtraction of wall thickness from the diameter of an equivalent 

hollow-cylinder cross section. The quadratic equation comes from combining Eq. 19 with: 

CmR = CoR – πhR (20) 
117 

Figure 67. Low magnification image of cross section through rupture mid-span of the 

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

CP-ECR.

Figure 68. Rupture cross section for Studsvik Test 185 with AF ZIRLO sample pressurized to 

1200 psig at 300°C. 
118 

Figure 69. Rupture cross section for Studsvik Test 192 with irradiated ZIRLO pressurized to 

1200 psig at 300°C. The central scale bar is 8 mm. 

For large rupture strains, the mid-wall circumference was only 2−3% less than the outer-surface 

circumference. The oxide layer thickness, which has a very small effect on outer-surface circumference, 

was not accounted for in this correction. 

Complete data sets are listed in Table 29. These data were used to develop a correlation for CmR as a 

function of measured DavgR and δR. The best linear fit to the decrease in CmR/(πDavgR) with the increase in

δR/DavgR is given by:

CmR/(πDavgR) = 1.00 – 0.207 (δR/DavgR) (21) 

Although Eq. 21 is the best fit to both the AF and PH ZIRLO data, it is not accurate as δR/DavgR

approaches zero. For ballooned cladding outside the rupture region, δR is zero and CmR/(πDavgR) should be 

<1. Also, for very small values of δR and large values of DavgR, CmR/(πDavgR) should be <1. 

As is shown in Fig. 70, the data points for PH ZIRLO are higher, relative to predicted values, than the 

data points for AF and irradiated ZIRLO. An alternate fit to the AF and irradiated ZIRLO data is given by 

Eq. 22 and shown as the solid line in Fig. 70. 

CmR/(πDavgR) = 0.980 – 0.225 (δR/DavgR) (22) 

As the correlation is needed in the current work only for AF ZIRLO samples for which no direct 

circumferential measurements were made, Eq. 22 was chosen to calculate rupture strain and minimum 

cladding wall thickness for these samples. The correlation can be rewritten as:
119 

Table 29. Summary of data for cross sections at rupture mid-spans for ZIRLO samples subjected 

to ramp-to-rupture tests and samples subjected to LOCA oxidation tests.

OCZL# 

ID Cladding

Cmi, 

mm

DavgR, 

mm

δR, 

mm

CoR, 

mm

CmR, 

mm

hR, 

mm

7 AF 28.05 12.45 3.7 ─ 34.1b 0.47 

11 AF 28.05 14.26 6.2 ─ 39.3b 0.41 

18 AF 28.05 14.60 5.4 ─ 40.0b 0.40 

185a AF 28.05 14.85 3.0 ─ 43.5b 0.37 

192a Irrad. ≈28 15.64 7.8 ─ 41.0b 0.38 

25 AF 28.05 14.95 8.7 41.0 39.7c 0.40 

29 AF 28.05 14.82 5.9 43.0 41.8c 0.38 

30 AF 28.05 14.51 4.5 41.0 39.7c 0.40 

32 AF 28.05 14.90 7.2 43.0 41.8c 0.38 

36 AF 28.02 15.04 10.2 44.0 42.8c 0.37 

37 AF 28.02 14.35 5.2 40.9 39.7c 0.39 

43 AF 27.99 14.04 6.1 43.0 42.8c 0.38 

39 PH 28.02 13.98 1.9 45.2 44.1c 0.36 

40 PH 27.99 13.51 4.2 42.3 41.1c 0.39 

41 PH 28.02 13.83 3.0 44.8 43.6c 0.37 

42 PH 28.02 14.88 1.2 49.0 48.0c 0.33 

44 PH 27.99 14.25 8.2 44.8 43.6c 0.37 

45 PH 27.99 14.94 5.0 48.0 46.9c 0.34 a

Studsvik tests [33−35].

b

Measured from low-magnification met images. c

Calculated from CoR data.

CmR = [0.980 – 0.225 (δR/DavgR)] (πDavgR) (23) 

Figure 71 shows the comparison of the CmR values determined from measurements and the calculated 

values (from Eq. 23). Given the inherent scatter in the database, the agreement is reasonably good for 10 

data points while the correlation-predicted values are low by about 4 mm for two data points. 

Table 30 lists the test ID#, initial fill pressure at 300°C, calculated or measured (in bold font) rupture 

strain, rupture temperature, average wall thickness calculated from rupture strain, and CP-ECR for tests 

conducted by Argonne with AF and PH ZIRLO. The CP-ECR values agree to within a fraction of a 

percent to previously published values [5] based on an older correlation. The CP-ECR values should be 

rounded to the nearest percent, as temperatures for calculating CP-predicted weight gain and wall 

thickness values are not accurate enough to justify three significant figures for CP-ECR. The purpose of 

listing CP-ECR values in Table 30 to the nearest 0.1% is to distinguish data points in subsequent plots of 

parameters vs. CP-ECR.
120 

Figure 70. Variation of CmR/(πDavgR) with δR/DavgR for AF, PH, and irradiated (Irrad.) ZIRLO

subjected to pressurized ramp-to-rupture tests and LOCA oxidation tests. 

Figure 71. Measured and calculated (Eq. 23) mid-plane circumferences in the rupture region 

for AF and irradiated (Irad.) ZIRLO following ramp-to-rupture and oxidation tests.

0.80

0.85

0.90

0.95

1.00

1.05

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7

CmR/(

π

DavgR) 

δR/DavgR

 AF ZIRLO Data

PH ZIRLO Data

Irrad. ZIRLO

Equation 21

Equation 22

30

32

34

36

38

40

42

44

46

48

50

30 32 34 36 38 40 42 44 46 48 50

Measured CmR (mm)

Calculated CmR (mm)

AF ZIRLO

Irad. ZIRLO
121 

Table 30. Internal pressure (Pi[300°C]), rupture strain (ΔCmR/Cmi, measured values in bold font), 

rupture temperature (TR), wall thickness (hR), and CP-ECR for AF and PH ZIRLO tests.

OCZL#

(Cladding)

Pi(300°C), 

psig

ΔCmR/Cmi, 

% 

TR, 

°C

hR, 

mm

CP-ECR,

% 

6 (AF) 1200 42 750±30 0.40 18.2 

7 (AF) 800 22 810±30 0.47 15.7 

8 (AF) 600 22 845±25 0.47 ≈0

9 (AF) 400 33 875±15 0.43 ≈0

10 (AF) 1600 70 715±10 0.33 ≈0

11 (AF) 1400 40 750 0.41 ≈0

12 (AF) 1000 33 805±20 0.43 14.5 

13 (AF) 1200 42 741±15 0.40 14.4 

14 (AF) 1200 48 735±6 0.39 18.3 

15 (AF) 1200 52 755±23 0.37 18.4 

16 (AF) 1200 35 734±15 0.42 13.9 

17 (AF) 1200 45 750±17 0.39 12.7 

18 (AF) 1200 43 748±4 0.40 12.3 

19 (AF) 600 25 840±12 0.46 17.3 

21 (AF) 600 28 850±10 0.44 10.6 

22 (AF) 600 23 837±12 0.46 11.6 

25 (AF) 1200 42 757±21 0.40 16.1 

26 (AF) 1200 34 765±39 0.42 15.9 

27 (AF) 1200 43 760±23 0.40 17.8 

29 (AF) 1200 49 746±19 0.38 17.3 

30 (AF) 1200 42 746±32 0.40 ─

32 (AF) 1200 49 748±8 0.38 15.9 

36 (AF) 1200 53 750±30 0.37 18.3 

37 (AF) 1200 46 755±30 0.39 23.1 

43 (AF) 1200 50 738±16 0.38 17.6 

39 (PH) 600 57 742±15 0.36 13.2 

40 (PH) 1200 47 690±10 0.39 11.9 

41 (PH) 1200 56 730±32 0.37 15.6 

42 (PH) 1200 71 682±5 0.33 14.4 

44 (PH) 1200 56 673±2 0.37 13.5 

45 (PH) 1200 68 737±11 0.34 13.2 
122 

The rupture-strain data in Table 30 are plotted against rupture temperature in Fig. 72. With the exception 

of two tests (OCZL#16 and OCZL#26), neither of which is included in Fig. 72, the agreement between 

correlation-calculated values and data is quite satisfactory for AF ZIRLO. For Test OCZL#16, the control 

TC experienced some electrical noise during ballooning and after rupture (see Fig. A.12) leading to 

rupture-temperature uncertainty. Also, the sample failed during disassembly just above the rupture region. 

OCLZ# 26 test results were excluded because of high uncertainty (about ±40°C) in rupture temperature.

Figure 72. Rupture strain vs. rupture temperature following ramp-to-rupture tests (AF 

ZIRLO) and post-rupture oxidation at ≈1200°C (AF and PH ZIRLO).

For AF ZIRLO heated at 5°C/s, the maximum rupture strain (70%) occurred at the minimum rupture 

temperature (715°C). For Zry-4, this temperature would correspond to the alpha phase regime. The alpha 

to alpha-beta transition temperature for standard (≈1.5 wt.% Sn) Zry-4 is about 810°C for equilibrium 

conditions [7]. This transition temperature decreases with decreasing Sn, oxygen, and hydrogen content, 

and it increases with increasing heating rate. Brachet et al. [36] report the following for the phasetransition temperature of low-Sn (1.3 wt. %) Zry-4: 790°C under equilibrium conditions and 860°C for a 

heating rate of 10°C/s. The alpha-phase transition temperature is lower for M5 (Zr-1Nb) with 0.14 wt.% 

oxygen: estimated equilibrium phase change temperature as low as 650°C and 840°C for a heating rate of 

10°C/s. Although the phase change temperature for ZIRLO (Zr-1Nb-1Sn) is not reported in the literature, 

it is expected to be between the phase-change temperatures for low-tin Zry-4 and M5 because of the 

intermediate tin content (1 wt. %). For the 5°C/s heating rate used in the current test program, 715°C 

could be within the alpha phase for ZIRLO where the peak rupture strain is expected to occur. The 

0

10

20

30

40

50

60

70

80

650 700 750 800 850 900 950

Rupture Strain (%)

Rupture Temperature (°C)

AF ZIRLO; 400-1600 psig

220-wppm ZIRLO, 1200 psig

390-wppm ZIRLO, 1200 psig

520-wppm ZIRLO, 1200 psig

660-wppm ZIRLO, 600 psig

700-wppm ZIRLO, 1200 psig
123 

rupture strain shown in Fig. 72 decreases monotonically for rupture temperatures <800°C, levels off at 

≈830°C, and increases for temperatures >840°C. Similar behavior has been observed for Zry-4 [32] 

although the transition temperatures are different due to differences in phase-change temperatures. For the 

reference fill pressures (at 300°C) used in this test program, there is remarkable consistency in results: (a) 

46±4% at 749±7°C based on 14 data points for 1200-psig fill pressure and (b) 25±3% at 843±6°C based 

on 4 data points for 600-psig fill pressure. 

Considerably higher rupture strains and lower rupture temperatures were observed for PH ZIRLO with 

220−700 wppm hydrogen. For the PH (660 wppm) sample with 600-psig fill pressure: the rupture strain 

was 57% (significantly higher than the 25% for AF ZIRLO) and the rupture temperature was 742°C 

(significantly lower than the 843°C for AF ZIRLO). Although the presence of hydrogen lowers the phase 

transition temperature (36), the results for PH ZIRLO cannot be explained based on phase change 

temperature alone. It appears that the presence of hydrogen weakens the non-oxidized cladding relative to 

the strength of AF ZIRLO such that the PH ZIRLO ruptures at a lower temperature and higher strain. 

While the trends shown in Fig. 72 are instructive, the values of rupture strain are likely particular to the 

parameters of the sample and the LOCA integral apparatus used. Although the uniform heating zone of 

the furnace is 254 mm, the axial extent for which the cladding temperatures are relatively “uniform” is 

only 100−125 mm. The axial temperature distribution along the test sample would limit the axial extent of 

the balloon region and likely reduce the rupture strain. Also, circumferential temperature variations just

prior to and at rupture were as high as ±32°C. These circumferential temperature variations would also 

limit the extent of rupture strain. It is highly likely that tests with longer uniform temperature regions, 

more uniform temperatures, and lower temperature ramp rates would result in considerably higher 

ballooning strains, as indicated in Ref. 32. However, in the current work, the purpose of generating 

rupture-strain data was to determine the wall thickness values for CP-ECR calculations. Thus, the rupture 

strain data in Table 30 and Fig. 72 have high significance for CP-ECR determination even though the data 

may not be applicable to what would be obtained using different sample lengths tested in different 

furnaces.

Although there are uncertainties in the data used to determine hR, the values presented in Table 30 are 

much more accurate than those determined from modeling codes, which assume symmetrical balloons 

and which use empirical correlations to determine rupture time. Ballooning and rupture codes cannot 

predict the cross-section shapes shown in Figs. 63−64 and 67−69. As the current effort is experimental 

rather than analytical, it was appropriate to determine hR as accurately as possible in order to establish 

data trends such as the variation of 4-PBT data with CP-ECR, which is based on hR. 

In addition to the data already presented in this section, metallographic examinations were performed to 

verify that the cross-sectional area remained constant during plastic deformation associated with 

ballooning and rupture. This assumption is the basis for Eq. 19. An example of the results of this effort is 

given in the following. For the OCZL#18 sample, the low-magnification image shown in Fig. 67 is very 

useful for measuring CmR, but it is inadequate for measuring hR. Higher magnification images are needed 

to measure the cladding metal thickness (100X) and the oxide layer thicknesses (200X). About 40 100X 

images would be needed to cover the whole cross section for this sample. As a compromise, Argonne did 

extensive metallographic examinations at 10 circumferential locations covering 25% of the cross section. 

On the basis of the 100X images, the average metal wall thickness was measured to be 0.35±0.12 mm. 

The inner- and outer-surface oxide layers were relatively uniform at 29±1 μm (0.058 mm total oxide 

thickness). For a Pilling-Bedworth ratio of 1.56, the cladding metal thickness prior to oxidation would be 

0.39±0.12 mm. Using the measured CmR value and the area-conservation assumption, Eq. 19 predicted hR

= 0.40 mm, which is in excellent agreement with the data. The data for average cladding metal thickness 
124 

prior to oxidation could be improved by analyzing 40 images, which would cover the whole cross section. 

However, the results for 10 images were adequate to demonstrate conservation of area. 

5.4 Bend-Test Results for AF ZIRLO

Samples used for 4-PBT data generation were subjected to the full LOCA integral test sequence, 

including quench. Reference conditions for LOCA integral tests were: 1200 psig or 600 psig prepressurization at 300°C, 5°C/s temperature ramp rate from 300°C to the target temperature of 1200°C 

(based on the control TC output), hold at temperature Th (based on average of TC readings) for specified 

times (th), 3°C/s temperature cooling rate from 1200°C to 800°C (based on control TC output), and 

quench at 800°C. Reference conditions for 4-PBTs were: (a) sample positioned such that the rupture tips 

were exposed to the maximum axial bending stress, (b) 135°C test temperature, (c) 2 mm/s (subsequently 

reduced to 1 mm/s) displacement rate, and (d) 14-mm maximum displacement. Under these LOCA 

integral test conditions, CP-ECR values of about 11% to 23% were achieved. 4-PBT parameters of 

interest were the maximum bending moment (Mmax) prior to failure, maximum applied energy (Emax) prior 

to failure, and offset displacement (δp) prior to failure. 4-PBT failure is defined as the complete severing 

of one or more cross sections and is easily discerned by the load drop to zero. Mmax is a measure of postquench strength, Emax is a measure of post-quench toughness, and δp is a measure of post-quench plastic 

deformation (for δp ≥0.1 mm). 

Raw data for all LOCA integral tests conducted are given in Appendices A (temperature histories), B 

(axial profiles of post-test outer diameters), and C (photographs of post-test samples). The raw data for 

4-PBTs are presented in Appendices D (load-displacement curves) and E (photographs of post-bend 

samples). In addition to these data, hydrogen-content measurements and metallographic examinations 

were routinely performed at the failure locations. For some test samples, both metallographic examination 

and oxygen-content measurements were conducted. For two of the data-generation test samples, the axial 

profile of hydrogen content was measured, and the axial profile of oxygen content was measured for one 

of these samples.

On the basis of scoping-test results, it is evident that several regions of the post-quench sample were 

vulnerable to failure during the 4-PBT: (a) the rupture mid-span with the smallest average cladding 

thickness and the highest oxidation level and (b) locations on either side of the rupture mid-span, between 

rupture edges and balloon necks, with sufficient oxidation and hydrogen pickup. As an aid to 

understanding phenomena relevant to the 4-PBT results presented in this section, it is useful to know the: 

(a) ballooning and rupture geometrical parameters, (b) effective oxidation hold time (the), (c) effective 

oxidation hold temperature (The), and (d) CP-ECR (values given in Table 30). These parameters are listed

in Table 31 for the 11 tests conducted under reference conditions, for one 4-PBT test conducted at RT, 

and for one reverse 4-PBT conducted with the rupture tips subjected to maximum compressive stress. 

The effective time (the) during which significant high-temperature oxidation occurs is longer than the 

programmed hold time because the controller slows down the temperature rate when the control 

temperature is within about 20°C of the target temperature. For a target temperature of 1200°C, the is the 

time for which the control TC reads ≥1180°C to the time corresponding to the end of the heating phase. 

The is the average temperature during the. 

Ballooning and rupture parameters of interest are the: (a) maximum circumferential strain (ΔCmR/Cmi

values are given in Table 30), (b) maximum width of the rupture opening (some δR values are given in 

Table 29), (c) axial length (LR) of the rupture opening, and (d) axial length (LB) of the balloon region. 
125

Table 31. Summary of LOCA integral post-quench parameters that may affect AF-ZIRLO cladding performance d

OCZL#

ΔCmR/Cmi, 

% 

δR, 

mm

LR, 

mm

LB, 

mm

the, 

s 

The, 

°C

CP-ECR,

% 

14 48 3.8 11 150 77 1194±5 18.3 Large cont

betwe

15 52 4.4 18 160 89 1199±19 18.4 R

17 45 5.7 19 170 30 1192±7 12.7 

18 43 5.4 18 160 30 1191±9 12.3 

19 25 3.3 12 140 141 1200±11 17.3 

21 28 4.9 13 160 22 1183±13 10.6 Low 

22 23 5.2 10 130 29 1216±11 11.6 

25 42 8.7 26 160 89 1202±11 16.1 2 relia

29 49 5.9 21 150 99 1197±7 17.3 

32 49 7.2 22 170 84 1181±11 15.9 RT 4-PBT

t

36 53 10.2 24 180 68 1215±18 18.3 Higher tha

37 46 5.2 19 160 172 1211±14 23.1 

43 50 6.1 19 160 74 1210±21 17.6 
126 

LB is defined as the axial span for which the average diametral stress is ≥5% (see Appendix B for 

diametral strain data). LR is significant because of the high oxygen and low hydrogen pickup for this 

region. Also, all failure locations given in this section are referenced to the mid-span of LR. LB is 

significant because the highest hydrogen pickup occurs between the edges of the rupture opening and the 

necks of the balloon region.

The results of the 4-PBTs are summarized in Table 32 for the 11 data-generation tests conducted under 

reference conditions, for the one reverse-bend test conducted with the rupture tips under axial 

compression, and for the one test conducted at RT. Results for the three metrics selected to characterize 4-

PBT performance are discussed in the following.

Table 32. Summary of results for LOCA integral tests terminated with quench at 800°C and postquench 4-PBTs with AF ZIRLO cladding. The 4-PBT displacement rate was decreased 

from 2 mm/s to 1 mm/s after OCZL#21 sample test.

OCZL#

Rupture

Strain, 

% 

CP-ECR, 

% 

4-PBT

Failure

Location, mm 

Mmax, 

N•m

Emax, 

J 

Offset

Displace., 

mm Comment

14 47 18.2 ≈0 5.7 0.24 0 

15 51 18.3 No Failure 9.0 >1.9 >10 Reverse 4-PBT

17 45 12.7 ≈0 8.4 0.65 >0.5 

18 43 12.3 ≈0 13.5 1.28 0 

19 25 17.3 -23, +23 5.7 0.23 0 

21 28 10.6 -29, +33 13.8 1.17 0 

22 23 11.6 -27, +25 11.1 0.83 0 

25 42 16.1 -26,+26 8.3 0.50 0 

29 49 17.3 ≈0 4.7 0.40 >9 <10% intact

32 49 15.9 ≈0 6.7 0.26 0 RT 4-PBT

36 53 18.3 -40, +32 9.5 0.56 0 

37 46 23.1 -40, +35 4.1 0.13 0 

43 50 17.6 ≈0 6.1 0.24 0 

Maximum Bending Moment (Mmax) 

Results for the three ramp-to-rupture scoping tests (see Table 28 and Figs. D.1 to D.3) indicate that AF 

ZIRLO with essentially no oxidation has an average Mmax value of about 20 N•m. Although none of these 

samples failed through 14-mm displacement, it appears that the bending moment leveled off at Mmax and 

began to decrease slowly. Mmax was relatively independent of rupture strain for values from 22% to 70%. 

As shown in Table 32, Mmax decreased significantly with increasing oxidation level (i.e., CP-ECR).
127 

The 4-PBT results indicate that samples failed either near the rupture node mid-span (0 mm) or at two 

locations between the rupture edges and the balloon necks. Mmax results are plotted separately in Fig. 73 

for these two cases. With the exception of the high Mmax value for the OCZL#36 sample, which appears to 

be an outlier, results for both cases appear to be essentially the same within expected data scatter. 

Excluding the OCZL#36 result, the best linear fit to the data for 10% to 19% CP-ECR is:

Mmax = 13.70 – 1.053 (CP-ECR – 10%), N•m (24) 

After about 19% (19.1% to be precise), the maximum bending moment appears to level off at 4.1 N•m on 

the basis of results at 23% CP-ECR. Although no data were generated in the oxidation range of >0% to 

10%, results from the ramp-to-rupture scoping tests (20 N•m) and the Eq. 24 value of about 14 N•m at 

10% CP-ECR can be used to estimate the effects of increasing oxidation level in this regime.

For the previous licensing limit (17% CP-ECR), Eq. 24 gives a Mmax value of 6.3 N•m. Data in the 

oxidation range of 17±1% CP-ECR give experimental values in the range of 5.8±1.4 N•m. It is interesting 

to note that the result (6.7 N•m) for the 4-PBT conducted at RT with a 17% CP-ECR sample is within this 

range. Although one RT data point does not constitute a trend, it appears that the maximum bending 

moment, which is the metric for maximum strength, is insensitive to 4-PBT temperature in the range of 

20°C to 135°C. The results for maximum strength are quite different from the RCT ductility results for 

which there was a very high sensitivity of ductility to RCT temperature.

The OCZL#12 sample (see Table 28) was cooled without quench prior to the 4-PBT. It is interesting to 

note that the sample failed at a cross section that included the mid-span of the rupture opening with a 

maximum bending moment of 8.8 N•m. Equation 24, which is based on data for post-quench samples 

predicts 9.1 N•m for Mmax. For this one sample, which failed in the high oxidation region with little 

hydrogen pickup, the absence of quench did not improve the maximum strength of the sample. 

The 4-PBT results for OCZL#17 and #18 are interesting to compare because the two samples had 

essentially the same ballooning and rupture parameters (see Table 31), the same effective hold 

temperature, and the same oxidation level (12.7% and 12.3% for OCZL#17 and #18, respectively). Also, 

both samples failed along cross sections that included the rupture opening. However, the maximum 

bending moments (8.4 N•m vs. 13.5 N•m) were quite different. The average of these experimental results 

(11 N•m) is in excellent agreement with the value predicted by Eq. 24 (11 N•m). The deviation of the data 

(about ±23%) is about the same as the ±24% for four tests conducted with samples that were oxidized to 

17±1% CP-ECR. 

For the reverse-bend-test sample from test OCZL#15 (18% CP-ECR), partial-wall cracking and 

crumbling were observed to occur in the compressive-stress region around the rupture opening. 

Concurrent with this degradation, the load decreased (see Figs. D.7a and D.7b) and the thicker regions of 

the cross section continued to deform plastically without cross-section failure (see Fig. E.5).

The 4-PBT strength results, based on Mmax, appear to be a meaningful measure of cladding performance 

in that the strength exhibits the expected decrease with increasing oxidation level. However, application 

of these results to LOCA acceptance criteria is not straightforward. In Section 5.6, post-quench bending 

strength results are compared to JAEA results for tests with a restraint-induced axial force during quench 

to assess the adequacy of the cladding strength determined from bend tests.
128 

Figure 73. Maximum 4-PBT bending moment (Mmax) vs. oxidation level (CP-ECR) for 

AF ZIRLO following ballooning, rupture, oxidation, and quench.

Significant insight into failure locations and modes can be gleaned from metallographic evaluation of 

layer thicknesses and from axial distributions of diametral strain, hydrogen content, and oxygen content. 

Detailed results are available for the OCZL#18 (12% CP-ECR) and OCZL#19 (17% CP-ECR) samples. 

The OCZL#18 sample severed along a cross section through the rupture node, while the OCZL#19 

sample severed at two locations that were between rupture edges and balloon necks. 

The OCZL#18 sample had large rupture strain (43%), thinner pre-oxidation cladding (0.40 mm average), 

and thinner metal wall (0.35 mm average) following oxidation. Outer- and inner-surface oxide layers 

were 29-μm thick. Figure 67 shows a low magnification image of the rupture-node cross section. Figures 

74a and 74b, respectively, show the thickest (0.44 mm) and thinnest (0.14 mm) metal wall thicknesses

after oxidation, quench, and 4-point bending. During the 4-PBT, there was an abrupt load drop from 540 

N to zero, indicating very rapid crack growth (see Fig. D.9). The sample failed near the axial center of the 

rupture opening. The results are interesting in that the local CP-ECR at the back wall was <10% and the 

hydrogen pickup was very small (see Fig. 75). If this segment had been tested in isolation, it would have 

behaved in a very ductile manner. However, it appears that this region, along with the whole cross 

section, did not have high enough toughness to blunt crack growth. Also, the cross section was weaker 

than regions between the rupture edges and the balloon necks, which had high hydrogen pickup (up to 

1800 wppm) and lower oxidation. 

The OCZL#19 sample had low rupture strain (25%) and thicker pre-oxidation cladding (0.46 mm 

average). The sample failed at locations that were 24 mm above and 24 mm below the rupture-opening 

mid-span (see Fig. 76). Both hydrogen and oxygen contents were measured along the sample (see Fig. 

77). It is clear from Figs. 76−77 that the failure occurred in regions where the combination of hydrogen 

and oxygen was high enough to weaken the material relative to the highest-oxygen/lowest-hydrogen 

rupture-node cross section and the lower-oxygen/highest hydrogen regions closer to the balloon necks.

0

2

4

6

8

10

12

14

16

9 11 13 15 17 19 21 23 25

Mmax (N•m) 

CP-ECR (%)

Rupture Node Failure

Non-Rupture Node Failure
129 

(a) Rupture tip

(b) Back side of cross section

Figure 74. Metallographic images of the OCZL#18 severed cross section following oxidation to 

12% CP-ECR, quench, and 4-point bending: (a) rupture tip (0.14-mm metal wall) 

and (b) back side of cross section (0.44-mm metal wall).

Figure 75. Axial profile of hydrogen content in OCZL#18 sample after oxidation to 12% 

CP-ECR.
130 

Figure 76. 4-PBT failure locations for the OCZL#19 sample after oxidation to 17% CP-ECR.

Figure 77. Axial profiles for hydrogen and oxygen content along the OCZL#19 sample after 

oxidation to 17% CP-ECR and 4-point bending.
131 

Metallographic examination was performed for the failure cross section 24 mm below the rupture 

opening. The average outer- and inner-surface oxide layers were 50-μm and 35-μm thick, respectively, 

and the metal wall was 0.46-mm thick. The pre-oxidation metal thickness was calculated to be 0.51 mm, 

which is thicker than the 0.46-mm average value calculated for the rupture-node cross section. The 

differences in pre-oxidation wall thicknesses between the failure cross section and the rupture cross 

section would suggest that the rupture cross section would be weaker. However, high-hydrogen/loweroxygen contents at ±24 mm from the rupture-opening mid-span had a larger effect on weakening the 

cladding at the failure locations than the strengthening effect of a thicker wall.

Failure Energy

The failure energy or maximum energy (Emax) when failure does not occur is plotted vs. oxidation level in 

Fig. 78, with data for failures within the rupture node cross section indicated by solid circles and data for 

non-rupture-node failures indicated by hollow circles. The load-displacement results from ramp-torupture scoping 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 these samples

accumulated through 14-mm displacement. The maximum energy decreased sharply from about 8 J for 

0% CP-ECR to about 1.2 J for 10% CP-ECR. Figure 78 shows that Emax decreased by a factor of ≈10 

from 10% CP-ECR to 19% CP-ECR and then leveled off at 0.13 J from 19% to 23% CP-ECR.

Figure 78. Maximum (failure) 4-PBT energy (Emax) vs. oxidation level (CP-ECR) for AF 

ZIRLO following ballooning, rupture, oxidation, and quench. 

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

Emax (J) 

CP-ECR (%)

Rupture Node Failure

Non-Rupture Node Failure
132 

Within the scatter of the data, there was no significant difference between failure energies for samples 

that failed in the rupture-opening cross section and samples that failed between the rupture edges and the 

balloon necks. At about 12% CP-ECR, the failure energies for the OCZL#22 and the OCZL#18 samples 

were both about 1 J (0.83 J and 1.28 J, respectively). The best-fit linear correlation (excluding the

suspected outlier OCZL#36 data point of 0.56 J at 18.3% CP-ECR) for failure energy vs. CP-ECR in the 

range of 10 to 19% CP-ECR is given by:

Emax = 1.204 – 0.1211 (CP-ECR – 10%), J (25) 

The failure energy vs. CP-ECR shows the same trends as does the maximum bending moment vs. CPECR. Namely, it shows a significant decrease in failure energy with oxidation level up to about 19% 

(18.9% to be precise). For a triangular load-displacement curve with linear rise in load up to the failure 

load followed by a very abrupt load drop, Emax gives about the same failure characterization as Mmax. 

However, it gives more information for samples that do not sever completely at the maximum load (e.g., 

see Figs. D.15a and D.15b for the OCZL#29 test sample) and continue to deform plastically at a 

decreasing load.

Offset Displacement

Offset displacement is a measure of plastic deformation. For the three ramp-to-rupture scoping-test 

samples (0% CP-ECR), 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 

Figs. D.1 to D.3). Measured offset displacements were in the range of 7 to 8 mm for these samples. The 

loading stiffness values were reasonably close, but they appeared to increase somewhat with decreasing 

rupture strain: 120 N/mm for 70% rupture strain, 130 N/mm for 33% rupture strain, and 135 N/mm for 

22% rupture strain.

The load-displacement curves for oxidized and quenched 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 displacement had

hydrogen contents and oxidation levels at the failure locations that would qualify them as brittle based on 

RCT results (Eq. 17) for non-deformed cladding samples. Metallographic imaging and analysis were 

performed at one of two severed cross-section locations for the post-bend OCZL#19 sample, and 

hydrogen content was determined from rings sectioned close to the severed cross-section (see Fig 76). 

The three quenched samples that severed in the cross section containing the rupture-opening node with 

zero offset displacement had oxidation levels of 12% CP-ECR and 18% CP-ECR (2 samples). Counting 

the 16% CP-ECR sample tested at RT, four samples severed in the rupture-opening node. 

Figure 79 shows the post-bend appearance of the OCZL#15 sample that was tested with the rupture tips in 

compression. This sample had an oxidation level of 18% CP-ECR. It was subjected to loading, unloading, 

and re-loading. As indicated in Figs. D7a and D7b, there was evidence of some offset displacement prior 

to the gradual load drop. As Fig. 79 shows, the sample continued to deform with significant plasticity as 

the partial cross-section crack grew.
133 

Figure 79. Post-bend appearance of the OCZL#15 sample oxidized to 18% CP-ECR, 

quenched, and subjected to a reverse 4-PBT with the rupture tips in compression.

Figure 80 shows the pos-test image of the OCZL#17 sample (13% CP-ECR), which was subjected to 

reference 4-PBT conditions. The back region of the cladding remained intact after 10-mm displacement 

with an offset displacement of about 9 mm. Figure D.15a shows that the maximum load was 335 N and 

the load appeared to drop abruptly to zero at a total displacement of about 3.4 mm. The test was 

terminated at 4-mm displacement to allow for examination of the sample. It was observed that the sample 

was partially intact. The sample was then reloaded. The measured loads during the reloading were very 

low (3N to 13 N) as the ligament at the back of the cladding continued to deform plastically. This plastic

deformation region contributed about 10% to the maximum energy. The images shown in Fig. 80 were 

taken after the reloading test was completed.

(a) Side view

(b) Rupture view

Figure 80. Post-bend-test images of the OCZL#17 sample oxidized to 13% CP-ECR, quenched, 

and subjected to bending at 135°C: (a) side view and (b) rupture view. 
134 

The OCZL#29 sample (17% CP-ECR) was also subjected to the reference 4-PBT. This test was 

terminated after ≈3-mm displacement (see Fig. D.15a) and reloaded (see Fig. D.15b). Cracking began at a 

relatively low load (188 N) and bending moment (4.7 N•m). The load dropped 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 (see Fig. 81a). After reloading, the crack continued to propagate in a 

plastic manner leaving only the back region intact (see Fig. 81b). 

(a) Side view after 1st loading-unloading cycle 

(b) Rupture view after 2nd loading-unloading cycle 

Figure 81. Post-bend-test images of the OCZL#29 sample (17% CP-ECR) after 4-PBTs 

conducted under reference conditions: (a) side view after 1st loading-unloading cycle 

and (b) rupture view after 2nd loading-unloading cycle. 
135 

The use of offset displacement as a measure of plastic deformation was found to be meaningful only for 

samples that did not partially or fully sever following achievement of the maximum load. These samples 

included AF cladding alloys (no oxidation) used for benchmark testing and ramp-to-rupture scoping test 

samples, also at 0% CP-ECR. In most cases, the AF ZIRLO samples exhibited rapid crack growth and no 

offset displacement following oxidation and quench. For two of the samples tested under reference bend 

conditions, more than 50% of the cross section cracked during the first load drop. This was followed by 

crack extension through most of the cross section and plastic flow of a ligament of the cladding at very 

low load. Given that it can only be correlated with plastic flow of the structure prior to crack initiation, 

offset displacement was not a very meaningful metric for post-quench samples with CP-ECRs in the 

range of 11% to 23%. The use of offset displacement may be a meaningful metric for samples oxidized 

within the range of 0% to 10% CP-ECR, as there is clearly a transition from high (>8 mm) to low (≈0 

mm) offset displacements in this oxidation range.

5.5 Bend-Test Results for PH ZIRLO

The pre-hydriding process is described in Section 3.2.2 for shorter (75−100 mm long) samples. The same 

process was used for 318-mm long ZIRLO LOCA integral samples. However, it was well recognized that 

axial variations in hydrogen content would be larger for the longer samples. The target hydrogen content 

for the mid-section of the cladding was 250±50 wppm. A number of practice samples were pre-hydrided 

and weighed to estimate average hydrogen content. Some of these were sectioned near the ends and midregions to get some idea of axial distribution of hydrogen. For some samples, hydrogen contents near the 

ends of the samples were higher than near the mid-regions. For other samples, the reverse was true. Also, 

most measured peak hydrogen contents were significantly higher than average values determined from 

weight-gain measurements. Reproducibility was also poor as there was wide variability in average 

hydrogen content for samples subjected to the same conditions for the same pre-hydriding time.

Engineering judgment was used to select six pre-hydrided samples for LOCA integral testing and 4-PBTs. 

Hydrogen contents near the cross section containing the rupture mid-span could only be determined by 

measuring local hydrogen content after ballooning, rupture, oxidation, quench and bending. As discussed 

in Section 3.2.2, these post-test hydrogen contents were corrected for oxygen-induced weight gain to 

determine pre-test and pre-rupture hydrogen content. For two of the samples, the pre-test hydrogen 

content (220 wppm) in the rupture-node cross section was within the target range. For the other samples, 

the hydrogen content (390−700 wppm) was higher than the target range.

Table 33 lists the PH ZIRLO samples tested, the hydrogen contents, and other parameters of interest such 

as those listed in Table 30 for AF ZIRLO. Five of the six PH ZIRLO samples survived disassembly and 

were available for 4-PBTs. The OCZL#42 sample failure near the rupture-opening mid-span during 

disassembly. For four of these five samples, there was very good temperature control during the heating 

ramp, the hold time and the cooling ramp. For LOCA integral test OCZL#41, the control TC reading 

fluctuated significantly during the heating ramp and during the cooling ramp just before and just after the 

hold time. As the output of the control TC determined the furnace power, the other TCs fluctuated 

significantly during this same time period. The last test (OCZL#45) was the best one in terms of 

temperature control, absence of secondary ballooning region, and rupture-opening orientation and width. 

Results of the 4-PBTs are listed in Table 34 for PH ZIRLO samples subjected to the reference LOCA 

integral test sequence, including quench. The results in Table 34 show a clear trend of decreasing Mmax

and Emax with increasing hydrogen content, as well as with increasing oxidation. This trend can be see 

more clearly in Figs. 82 (Mmax) and 83 (Emax), which show the comparison between results for PH and AF 

ZIRLO cladding samples.
136

Table 33. Summary of LOCA integral post-quench parameters that may affect PH-ZIRLO cladding performance d

OCZL#

(wppm)

ΔCmR/Cmi, 

% 

δR, 

mm

LR, 

mm

LB, 

mm

the, 

s 

The, 

°C

CP-ECR,

% 

39 

(660) 

57 1.9 7.3 120 20 1195±19 13.2 2 relia

Large sec

50 mm 

40 

(390) 

47 4.2 16 130 14 1196±14 11.9 3 relia

Regu

41 

(220) 

56 3.3 15 150 20 1212±52 15.6 Large cont

during he

Slanted ru

a

42 

(520) 

71 0.8 7.5 140 22 1202±39 14.4 Control TC 

Very s

Sample fai

44 

(700) 

56 8.2 18 130 23 1201±19 13.5 

Large ruptu

m

45 

(220) 

68 5.0 17 150 14 1188±9 13.2 
 

137 

Table 34. Summary of results for LOCA integral tests terminated with quench at 800°C and postquench 4-PBTs with PH ZIRLO cladding. 4-PBT displacement rate was 1 mm/s. 

OCZL#

Rupture

Strain, 

% 

CP-ECR, 

% 

4-PBT

Failure

Location, mm 

Mmax, 

N•m

Emax, 

J 

Offset

Displace., 

mm Comment

39 

(660) 

57 13.2 ≈0 2.8 0.05 0 ─

40 

(390) 

47 11.9 ≈0 4.9 0.16 0 ─

41 

(220) 

56 15.6 ≈0 3.2 0.07 ≈0 ─

44 

(700) 

56 13.5 ≈0 2.4 0.11 ≈0 Some offset 

displ. after 

>50% cracking

45 

(220) 

68 13.2 ≈0 6.6 0.28 0 ─

Figure 82. Maximum 4-PBT bending moment (Mmax) vs. oxidation level (CP-ECR) for AF 

and PH ZIRLO following ballooning, rupture, oxidation, and quench.

0

2

4

6

8

10

12

14

16

9 11 13 15 17 19 21 23 25

Mmax (N•m) 

CP-ECR (%)

AF ZIRLO

220-wppm ZIRLO

390-wppm ZIRLO

660-wppm ZIRLO

700-wppm ZIRLO
 

138 

Figure 83. Maximum (failure) 4-PBT energy (Emax) vs. oxidation level (CP-ECR) for AF 

and PH ZIRLO following ballooning, rupture, oxidation, and quench.

The Mmax and Emax data points for ZIRLO samples with 220 wppm are of particular interest in assessing 

whether or not such samples have sufficient values of Mmax and Emax if the oxidation level were limited. In 

Section 5.4, it is established that the correlation value (Eq. 24) for the Mmax of AF ZIRLO is 6.3 N•m at

17% CP-ECR. The corresponding value (Eq. 25) for Emax is 0.36 J at 17% CP-ECR. The 220-wppm 

sample at 13.2% CP-ECR had a maximum bending moment of 6.6 N•m, which is marginally above the 

17% CP-ECR for AF cladding. However, if the oxidation level were limited based on retention of 

ductility (Fig. 25 and Eq. 17), then the CP-ECR would be limited to 11%. On the basis of linear 

extrapolation, the 220-wppm ZIRLO sample would have Mmax and maximum Emax values of 9.6 N•m and 

0.46 J, both of which are greater than the corresponding values for AF ZIRLO cladding oxidized to 17% 

CP-ECR. Thus, it appears that limiting oxidation level based on ductility retention for non-deformed 

cladding would be adequate to ensure that cladding with pre-test hydrogen had Mmax and Emax levels at 

least as high as levels for AF cladding at 17% CP-ECR. The PH ZIRLO sample at 220-wppm and 13.2% 

CP-ECR was of high quality because temperatures were well controlled during the oxidation phase of the 

LOCA integral test. The sample at 220-wppm and 15.6% was of much lower quality because of 

significant temperature fluctuations (due to electrical noise) just before and just after the hold time. A 

more conservative approach for estimating Mmax and Emax at 11% CP-ECR for 220 wppm pre-test 

hydrogen is to use the slopes in Eqs. 24 and 25 for extrapolation. This approach leads to values of 8.9 

N•m Mmax and 0.55 J Emax, which are still above the 17% CP-ECR values for AF ZIRLO.

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

Emax (J) 

CP-ECR (%)

AF ZIRLO

220-wppm ZIRLO

390-wppm ZIRLO

660-wppm ZIRLO

700-wppm ZIRLO
 

139 

Considerable characterization work was conducted after the 4-PBTs to determine the hydrogen and 

oxygen content within the rupture-node cross sections near the failure region, as well as the thickness of 

the outer-surface oxide layer in this region. The oxygen content was used to determine weight gain, which 

was then used to calculate pre-test hydrogen content. 

Figure 84 shows results of hydrogen- and oxygen content measurements for the OCZL#39 sample 

following severing through a rupture-node cross section during the 4-PBT. Also shown in Fig. 84 is the 

secondary balloon that formed about 50 mm above (to the right in the figure) the primary balloon. The 

outer-surface oxide-layer thickness was measured for the metallographic-sample cross section containing 

the rupture opening near the bottom edge of this opening. The measured value (30±3 μm) was about 7% 

higher than the CP-predicted value (28±3 μm, based on two TCs). Agreement to within 10% is considered 

acceptable. Also, the circumferential variation in oxide layer thickness is consistent with the measured 

circumferential variation in the effective hold temperature (1195±19°C).

Figure 84. Sectioning diagram and characterization results for the OCZL#39 post-bend sample 

with 660-wppm pre-test hydrogen oxidized to 13.2% CP-ECR. 

Figure 85 shows the hydrogen and oxygen contents measured for rings close to the severed cross section 

of the OCZL#40 sample. Three TCs remained reliable through the temperature history for this test. 

Measured (26±4 μm) and CP-predicted (27±2 μm) values for outer-surface oxide layer thickness were 

very close and the circumferential temperature variation (1196±14°C) was relatively mild. 

Figure 85. Sectioning diagram and characterization results for the OCZL#40 post-bend sample 

with 390-wppm pre-test hydrogen oxidized to 11.9% CP-ECR.
 

140 

The sectioning diagram and characterization results for the post-bend OCZL#41 sample are shown in Fig. 

86 for regions close to the severed cross section. The measured (31±4 μm) and calculated (33±6 μm) 

outer-surface oxide layer thicknesses were closer than expected given the temperature fluctuations before 

and after the hold time. The effective hold temperature was 1212±52°C with significant uncertainty in the 

±52°C, which would have lead to larger variation in measured oxide layer thickness,

Figure 86. Sectioning diagram and characterization results for the OCZL#41 post-bend sample 

with 220-wppm pre-test hydrogen oxidized to 15.6% CP-ECR.

The OCZL#42 sample could not be subjected to a 4-PBT because the sample severed at a location within 

the rupture opening during disassembly. Hydrogen and oxygen measurements were made on rings 

sectioned from just above and just below the severed cross section to determine the pre-test hydrogen 

content (520 wppm).

Figure 87 shows the sectioning diagram and characterization results for the OCZL#44 test sample 

following the 4-PBT. The rupture opening was rather unique for this sample (see pre-bend photo in Fig. 

C.29) as it was bounded by a flap of metal resembling a partially opened door. 

Figure 87. Sectioning diagram and characterization results for the OCZL#44 post-bend sample 

with 700-wppm pre-test hydrogen oxidized to 13.5% CP-ECR.
 

141 

Figure 88 shows the sectioning diagram and characterization of the OCZL#45 sample after the 4-PBT. 

This was perhaps the most reliable PH ZIRLO sample because the four TCs remained intact during the 

thermal history and the circumferential temperature variation (1188±9°C) was low.

Figure 88. Sectioning diagram and characterization results for the OCZL#45 post-bend sample 

with 220-wppm pre-test hydrogen oxidized to 13.2% CP-ECR.

5.6 Discussion of Strength of Ballooned and Ruptured Cladding

Sections 5.4 and 5.5 establish maximum bending moments and failure energies for AF and PH ZIRLO 

cladding, respectively, as a function of oxidation level and pre-test hydrogen content. Both 4-PBT metrics 

decreased with increasing oxidation level (CP-ECR) and increasing pre-test hydrogen content. In order to 

determine oxidation-level and hydrogen-content limits to retain “adequate” post-quench strength, the term 

adequate needs to be quantified. Two approaches were considered for quantifying adequate strength. 

The first approach comes from observations made in Section 5.5. If oxidation is limited according to Eq. 

17 (see Fig. 25), then Mmax and Emax for PH ZIRLO are at least as high as for AF ZIRLO with oxidation 

limited to 17%. It was also shown in Section 3.4 that pre-hydrided cladding is a good surrogate for highexposure cladding with the same hydrogen content. It is further known that Mmax and Emax are measures of 

strength and toughness, respectively. Therefore, it can be concluded that the strength and toughness of 

high-exposure and ruptured cladding containing hydrogen is at least as high as that of fresh cladding at 

17% CP-ECR. Because the 17% limit has been used as a measure of adequacy for fresh cladding by the 

NRC for more than 40 years, it can be concluded that the same level of adequacy for high-exposure 

cladding would be maintained using Eq. 17 to limit the oxidation level based on the pre-LOCA-transient 

hydrogen content. 

The second approach, which is more quantitative, comes from considering axial loading during quench 

that results from partial restraint of axial contraction during quench cooling due to bonding of the 

cladding to the grid spacers. JAEA and others have estimated limits for the axial restraint load due to this 

phenomenon. In addition, JAEA has generated a large data set for failure/non-failure loads for AF Zry-4, 

PH Zry-4, and irradiated Zry-4 under conditions of full restraint (AF and PH) and load-limited restraint 

(PH and irradiated) [37−40]. For the load-limited restraint cases, loads of 735 N, 540 N, and 390 N were 

used. The load limit of 540 N was adopted as the reference case for testing irradiated material even 

though it was argued by JAEA that this was a very conservative upper bound on the restraint load that 

could be imposed by cladding bonded to grid spacers. In the JAEA tests, the cladding was restrained from 

contracting during slower cooling from the oxidation temperature to the quench temperature (700°C) and
 

142 

during rapid quench cooling until the load limits were reached. For the remainder of the quench to 100°C, 

the axial load on the cladding was maintained at the load limit.

In this section, the JAEA results for fully restrained AF 17×17 Zry-4 are compared to the Argonne 4-PBT 

results for AF 17×17 ZIRLO. In both types of tests, the cladding was exposed to axial stresses. The 

maximum stress for which Argonne samples did fail is compared to the maximum stress for which JAEA 

fully constrained cladding did not fail. Thus, adequate strength in this approach is defined as the strength 

needed to survive full axial constraint during quench cooling.

The description of the JAEA test parameters and results is documented in Refs. 37−40 and summarized in 

Ref. 5, which also includes the detailed analytical method for relating the maximum axial tensile stress in 

4-PBT samples to the maximum axial tensile stress in fully restrained samples during cooling. For the 

convenience of the reader, the analytical method and results are repeated below.

Traditional axial tension tests are conducted either in the displacement-controlled mode or the loadcontrolled mode. In the displacement-controlled mode, the displacement rate (and hence displacement) is 

the independent parameter and the load is measured as a function of displacement. The RCTs and 4-PBTs 

used in the current work were displacement-controlled tests. The JAEA LOCA integral samples were not 

tested in the traditional way. The cooling rate from the oxidation temperature (e.g., 1200°C) to the quench 

temperature (700°C) and end-of-quench cladding temperature (≤100°C) were controlled; the load was 

measured as a function of time. For fully constrained tests, the displacement was zero. 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.

The JAEA-measured loads were relatively low during cooling from the oxidation hold temperature to the 

700°C quench temperature due to the low yield strength and high plasticity of the cladding metal at 

elevated temperatures and the relatively low cooling rate. However, the load increased rapidly during 

quench cooling because the metal yield strength increased with decreasing temperature and the metal did 

not have time to flow plastically to relax the axial thermal stresses induced by the restraint.

ANL 4-PBT results can be compared to JAEA axial-restraint results as both tests induce axial stresses. 

The stress distributions are different: (a) for the JAEA tests, the stress acting on a cross section is tensile 

and (b) for the 4-PBT, the axial stress acting on a cross section 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 same maximum tensile 

stress can be determined. This determination can be made more easily for 4-PBT samples that failed

outside the rupture region. The 4-PBT results for the OCZL#19 test sample are used for this comparison 

because: (a) the temperature control during the oxidation phase was excellent (1200±11°C), (b) the CPECR was about 17% (17.3%), (c) the oxidation level and hydrogen content were well characterized at one 

of two failure locations, and (d) metallographic examination was used to determine oxide-layer and 

metal-wall thicknesses at one failure location. Figure 89a (repeat of Fig. 76) shows the failure locations 

and some of the characterization results for the failure cross sections that occurred 24-mm below the 

rupture mid-span. Figure 89b shows a low magnification image of the severed cross section, which is 

nearly circular in shape. Table 35 defines the geometric and material properties needed for analyses and 

lists properties and dimensions determined from diameter measurements, lower-magnification (100X) 

metallographic images, and higher-magnification (200X) metallographic images. 
 

143 

(a) Failure locations

(b) Severed cross section at 24-mm below (left) the rupture mid-span

Figure 89. Images for the post-bend OCZL#19 sample oxidized to 17.3% CP-ECR and 

quenched: (a) failure locations and (b) severed cross section at 24-mm below 

the rupture mid-span.
 

144 

Table 35. Dimensions, properties, and failure bending moment at severed location for the 

OCZL#19 test sample with 17% CP-ECR in the rupture node and 12% measured ECR 

at the severed location 24-mm below the rupture mid-span. 

Parameter Value Comment

Outer Diameter (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 (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 [41]

MATPRO [42]

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 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 (26) 

where Rmo is the outer radius (Rmo = Dmo/2) of the metal wall (5.285 mm in this case). This is the location 

of the outer surface of the oxygen-stabilized alpha layer, which has low strength and is more brittle than 

the tetragonal oxide layer and the beta layer.

The maximum tensile stress in the metal is thus, σmax = 5.7 N•m (5.285 mm)/248 mm4

 = 121.5 MPa. 

For the axial tensile test, the maximum load (Pmax) is related to the maximum tensile stress (σmax) in the 

metal according to: 

Pmax = σmax Aeq (27) 

Setting σmax to 121.5 MPa and Aeq to 19.3 mm2 gives Pmax = 2345 N. If the correlation value (see Eq. 24) 

of 6.3 N•m at 17% CP-ECR had been used, then Pmax would be equal to 2591 N. 
 

145 

For the JAEA fully restrained samples [37,38], the maximum load measured for AF 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% Baker-Just-ECR, the maximum load was 1200 to 2000 N for PH samples that survived quench. 

These results are for cladding with pre-test hydrogen contents ≥100 wppm. Thus, even under these 

conditions, it is evident that the OCZL#19 test sample would have enough strength to survive the 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 fullconstraint test with quench from 700°C to 135°C. However, based on one 4-PBT conducted at 20°C, the 

result for Mmax was not sensitive to test temperature. Thus, the Argonne 4-PBT results would apply to the 

end-of-quench temperature (≤100°C) achieved in the JAEA tests.

Equations 26 and 27 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) (28) 

In principle, Eq. 28 can be used for samples that severed in the rupture node (e.g., OCZL#29 at 17% CPECR). However, determination of the equivalent area, the location of the neutral axis, and the equivalent 

area moment of inertia would require numerical integration for cross sections containing a rupture node. 

The Em (86.5 GPa) used for cladding metal was calculated from the MATPRO correlation for asfabricated, recrystallized-annealed cladding at 135°C. This requires validation because the Young’s 

modulus (Emβ) in the prior beta-phase is 75.6 GPa, and it increases with oxygen content in the oxygenstabilized alpha layer from 95.6 GPa (alpha/beta boundary) to 115 GPa (alpha/oxide boundary). The 

analysis was repeated by assuming that the ID/OD alpha layer thicknesses were the same as the ID/OD 

oxide layer thicknesses and the Young’s modulus (Emα) was 105 MPa. Using 75.6 GPa for Emβ and 105 

GPa for Emα resulted in a change in Pmax of only <0.5%. Thus, the use of 86.5 GPa for Em is justified.

Argonne’s test samples with 220-wppm hydrogen would also have survived the JAEA fully restrained 

quench tests as long as the oxidation were limited to <17%. The Mmax data reported in Section 5.5 for 

these two samples are: 6.6 N•m at 13.2% CP-ECR and 3.2 N•m at 15.6% CP-ECR. It appears as if the 

sample with 220-wppm and 13.2% CP-ECR would have survived the JAEA fully restrained quench test 

whereas the sample with 220-wppm and 15.6% CP-ECR might not survive. However, if the oxidation 

limit for 220 wppm hydrogen were based on data shown in Fig. 25 and the Eq. 17 correlation value for 

ductility retention in non-deformed samples, this oxidation limit would be 11%. As shown in Section 5.5, 

extrapolation of the 220-wppm data to 11% CP-ECR gives Mmax in the range of 8.9 N•m to 9.6 N•m. 

Although clearly conservative, it appears that adequate strength of PH samples can be preserved if Eq. 17 

is used to limit the oxidation level. This result is fortuitous as Eq. 17 is based on retention of ductility for 

uniformly oxidized non-deformed cladding rings subjected to RCTs. 

For the other three PH samples tested, hydrogen contents were ≥390 wppm, which would limit oxidation 

levels to ≤7 % CP-ECR, as well as the peak oxidation temperature for >400 wppm. Although no data 

could be generated for CP-ECR values of >0% to 10% at 1200°C peak temperature, it is speculated that 

these samples would survive the JAEA fully restrained test if oxidation levels were limited by Eq. 17.

Upper-bound conservatism has been incorporated in the quantification of adequate strength. 

Quantification is based on the assumption that the cladding must be strong enough to survive full axial 

restraint during quench cooling. It is highly unlikely that the cladding/grid-spacer bond and the grid

spacer springs could support such high loads. 

 

147 

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24. J.P. Mardon and N. Waeckel, “Behavior of M5TM alloy under LOCA conditions,” ANS-Top-Fuel, 

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Simulated LOCA Conditions,” JNST, Vol. 43, No. 9 (2006) 1114-1119. 

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NUREG/CR-6150, Vol. 4, Rev. 2, ML010330424, January 2001. 

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Oct. 31, 2014. 
APPENDIX A: TEMPERATURE AND PRESSURE HISTORIES

A-1

Figure A.1: Temperature histories for the OCZL#5 thermal benchmark test with unpressurized ZIRLO claddin

pellets. Sample was subjected to outer-surface oxidation with 40 ± 2 μm predicted and measured oxi

the TC locations. 

900

800

700

600

500

400

300

200

1300

1200

1100

1000

300 400 500 600 700 

C) ° Temperature (

Time (s)

0°, 51 mm above (Ctl TC)

0°, mid-plane

0°, 51 mm below

180°: mid-plane

at 600s:

Tavg =1206±9°C

5°C/s

697±6°C

800±2°C 

908±2°C 
A-2

Figure A.2: Temperature and pressure histories for thermal benchmark test OCZL#6 pressurized to 1200 psig, o

ECR, and cooled without quench. The control TC orientation was about 180° from the rupture open

temperature recorded on the rupture side was confirmed by oxide-thickness measurements.

900

800

700

600

500

400

300

200

1300

1200

1100

1000

100 200 300 400 500 600 700 

C) ° Temperature (

Time (s)

0°, 51 mm above (Ctl TC)

0°, mid-plane

0°, 51 mm below

180°, mid-plane

Pressure, psig

Rupture at 750±30°C

1276°C 

1204±7°C 
A-3

Figure A.3: Temperature and pressure histories for thermal-benchmark/4-point-bend test OCZL#7 pressurized 

to 16% CP-ECR, and cooled without quench. The control TC orientation was about 180° from the ru

900

800

700

600

500

400

300

200

1300

1200

1100

1000

100 200 300 400 500 600 700 

C) ° Temperature (

Time (s)

0°, 51 mm above (Ctrl TC)

0°, mid-plane

0°, 51 mm below

180°, mid-plane

Pressure, psig

Rupture at 810±30°C 

1211±13°C 

1260°C 
A-4

Figure A.4: Temperature and pressure histories for ramp-to-rupture scoping test OCZL#8 pressurized to 600 ps

quench.

900

800

700

600

500

400

300

200

1000

300 350 400 450 500 

C) ° Temperature (

Time (s)

 0°, 51 mm above (Ctrl TC)

0°, mid-plane

0°, 51 mm below

180°, mid-plane

Pressure, psig

Rupture at 845±25°C 
A-5

Figure A.5: Temperature and pressure histories for ramp-to-rupture scoping test OCZL#9 pressurized to 400 ps

quench.

900

800

700

600

500

400

300

200

1000

350 370 390 410 430 450 470 490

C) ° Temperature (

Time (s)

0°, 51 mm above (Ctrl TC)

0°, mid-plane

0°, 51 mm below

180°, mid-plane

0°, top plenum region

Pressure, psig

Rupture at 875±15°C 
A-6

Figure A.6: Temperature and pressure histories for ramp-to-rupture scoping test OCZL#10 pressurized to 1600 

without quench.

800

600

400

200

1600

1400

1200

1000

400 425 450 475 500 525 

C) ° Temperature (

Time (s)

0°, 51 mm above (Ctrl-TC)

0°, mid-plane

0°, 51 mm below

180°, mid-plane

Pressure, psig

Rupture at 715±10°C 
A-7

Figure A.7: Temperature and pressure histories for ramp-to-rupture scoping test OCZL#11 pressurized to 1400 

without quench.

900

800

700

600

500

400

300

200

1500

1400

1300

1200

1100

1000

200 300 400 500 600 700 800 

C) ° Temperature (

Time (s)

0°, 51 mm above (Ctl TC)

0°, mid-plane

0°, 51 mm below

180°, mid-plane

Pressure, psig

Rupture at ≈750°C 
A-8

Figure A.8: Temperature and pressure histories for scoping test OCZL#12 pressurized to 1000 psig, oxidized at 

without quench.

900

800

700

600

500

400

300

200

1300

1200

1100

1000

200 300 400 500 600 700 

C) ° Temperature (

Time (s)

0°, 51 mm above (Ctl TC)

0°, mid-plane

0°, 51 mm below

180°, mid-plane

Pressure, psig

Ruptue at 805±20°C 

1196±16°C
A-9

Figure A.9: Temperature and pressure histories for scoping test OCZL#13 pressurized to 1200 psig, oxidized at ≈

ECR , and cooled without quench.

900

800

700

600

500

400

300

200

1300

1200

1100

1000

100 200 300 400 500 600 700 800 

C) ° Temperature (

Time (s)

0°, 51-mm above (Ctrl-TC)

180°, 51-mm above

90°, 51-mm below

270°, 51-mm below

Pressure, psig

Rupture at 741±15°C 
A-10

Figure A.10: Temperature and pressure histories for data-generation test OCZL#14 pressurized to 1200 psig, oxid

18% CP-ECR, and cooled with quench.

900

800

700

600

500

400

300

200

1300

1200

1100

1000

100 200 300 400 500 600 700 8

C) ° Temperature (

Time (s)

0°, 57-mm above (Ctrl-TC)

180°, 57-mm above

270°, 44-mm below

Pressure, psig

Rupture at 735±6°C 

1274±11°C 
A-11

Figure A.11: Temperature and pressure histories for data-generation test OCZL#15 pressurized to 1200 psig, oxid

18% CP-ECR, and cooled with quench.

900

800

700

600

500

400

300

200

1300

1200

1100

1000

300 400 500 600 700 

C) ° Temperature (

Time (s)

180°, 57-mm above (Ctrl-TC)

90°, 44-mm below

270°, 44-mm below

Pressure, psig

Rupture at 755±23°C 
A-12

Figure A.12: Temperature and pressure histories for data-generation test OCZL#16 pressurized to 1200 psig, oxid

14% CP-ECR, and cooled with quench.

900

800

700

600

500

400

300

200

1300

1200

1100

1000

300 400 500 600 700 

C) ° Temperature (

Time (s)

180°, 57-mm above

90°, 44-mm below

Pressure, psig

0°, 57-mm above (Ctrl-TC)

Rupture at 734±15°C 
A-13

Figure A.13: Temperature and pressure histories for data-generation test OCZL#17 pressurized to 1200 psig, oxid

to13% CP-ECR, and cooled with quench.

900

800

700

600

500

400

300

200

1300

1200

1100

1000

300 350 400 450 500 550 600 650 700 

C) ° Temperature (

Time (s)

180°, 57-mm above (Ctrl-TC)

90°, 44-mm below

270°, 44-mm below

Pressure, psig

0°, 57-mm above

Rupture at 750±17°C 
A-14

Figure A.14: Temperature and pressure histories for data-generation test OCZL#18 pressurized to 1200 psig, oxid

12% CP-ECR, and cooled with quench.

900

800

700

600

500

400

300

200

1300

1200

1100

1000

200 300 400 500 600 700 

C) ° Temperature (

Time (s)

180°, 57-mm above (Ctrl-TC

270°, 44-mm below

Pressure, psig

0°, 57-mm above

Rupture at 748±4°C 
A-15

Figure A.15: Temperature and pressure histories for data-generation test OCZL#19 pressurized to 600 psig, oxid

17% CP-ECR, and cooled with quench.

900

800

700

600

500

400

300

200

1300

1200

1100

1000

300 350 400 450 500 550 600 650 700 750 800 

C) ° Temperature (

Time (s)

180°, 57-mm above

90°, 44-mm below

270°, 44-mm below

Pressure, psig

0°, 57-mm above (Ctrl-TC)

Rupture at 840±12°C 1201±11°C 
A-16

Figure A.16: Temperature and pressure histories for data-generation test OCZL#21 pressurized to 600 psig, oxid

11% CP-ECR, and cooled with quench.

900

800

700

600

500

400

300

200

1300

1200

1100

1000

300 350 400 450 500 550 600 650 700 

C) ° Temperature (

Time (s)

180°, 57-mm above

90°, 44-mm below

270°, 44-mm below

Pressure, psig

0°, 57-mm above (Ctrl-TC)

Rupture at 850±10°C 

1201±11°C 
A-17

Figure A.17: Temperature and pressure histories for data-generation test OCZL#22 pressurized to 600 psig, oxid

12% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1300

1200

1100

1000

300 350 400 450 500 550 600 650 700 

C) ° Temperature (

Time (s)

180°, 57-mm above (Ctrl-Tc)

90°, 44-mm below

270°, 44-mm below

Pressure, psig

0°, 57-mm above

Rupture at 837±12°C 

1202±10°C 

5-s Hold
A-18

Figure A.18: Temperature and pressure histories for data-generation test OCZL#24 non-pressurized, one-sided o

17% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1300

1200

1100

1000

300 500 700 900 1100 1300 1500 

C) ° Temperature (

Time (s)

180°, 57-mm above

90°, 44-mm below

270°, 44-mm below

0°, 57-mm above (Ctrl-Tc)

 1186±14°C 
A-19

Figure A.19: Temperature and pressure histories for data-generation test OCZL#25 pressurized to 1200 psig, oxid

16% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1300

1200

1100

1000

350 400 450 500 550 600 650 700 750 

C) ° Temperature (

Time (s)

180°, 57-mm above (Ctrl-TC)

90°, 44-mm below

270°, 44-mm below

Pressure, psig

0°, 57-mm above

Rupture at 757±21°C 
A-20

Figure A.20: Temperature and pressure histories for data-generation test OCZL#29 pressurized to 1200 psig, oxid

17% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1300

1200

1100

1000

350 400 450 500 550 600 650 700 750 

C) ° Temperature (

Time (s)

0°, 57-mm above

180°, 57-mm above (Ctrl-TC)

90°, 44-mm below

270°, 44-mm below

Pressure, psig

Rupture at 746±19°C

1198±12°C 
A-21

Figure A.21: Temperature and pressure histories for data-generation test OCZL#32 pressurized to 1200 psig, oxid

16% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1300

1200

1100

1000

350 400 450 500 550 600 650 700 750 

C) ° Temperature (

Time (s)

0°, 57-mm above

180°, 57-mm above (Ctrl-TC)

90°, 44-mm below

270°, 44-mm below

Pressure, psig

Rupture at 745±8°C 

1181±18°C 
A-22

Figure A.22: Temperature and pressure histories for thermal-benchmark test OCZL#36 pressurized to 1200 psig

to 18% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1300

1200

1100

1000

150 250 350 450 550 650 750 8

C) ° Temperature (

Time (s)

TC1: 57-mm above, 0°

TC2 (ctrl TC): 57-mm above, 180°

TC3: 44-mm below, 90°

TC4: 44-mm below, 270°

Pressure, psig

Rupture at 750±30°C

1216±17°C 
A-23

Figure A.23: Temperature and pressure histories for thermal/4-point-bend benchmark test OCZL#37 pressurized 

at ≈1200°C to 23% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1300

1200

1100

1000

300 350 400 450 500 550 600 650 700 750 800 850 900 

C) ° Temperature (

Time (s)

TC1: 57-mm above, 0°

TC2 (ctrl TC): 57-mm above, 180°

TC3: 44-mm below, 90°

TC4: 44-mm below, 270°

Pressure, psig

Rupture at 755±30°C

1211±16°C 
A-24

Figure A.24: Temperature and pressure histories for data-generation test OCZL#43 pressurized to 1200 psig, oxid

18% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1400

1300

1200

1100

1000

300 350 400 450 500 550 600 650 700 750 8

C) ° Temperature (

Time (s)

TC1: 57-mm above, 0°

TC2 (ctrl TC): 57-mm above, 1

TC3: 44-mm below, 90°

TC4: 44-mm below, 270°

Pressure, psig

Rupture at 738±16°C

1211±25°C
A-25

Figure A.25: Temperature and pressure histories for data-generation test OCZL#39 with pre-hydrided (660 wppm

to 600 psig, oxidized at ≈1200°C to 13% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1300

1200

1100

1000

300 350 400 450 500 550 600 650 700 7

C) ° Temperature (

TC1: 57-mm above, 0°

TC2 (ctrl TC): 57-mm above, 180°

TC3: 44-mm below, 90°

TC4: 44-mm below, 270°

Pressure, psig

Time (s)

Rupture at 742±15°C

1196±27°C
A-26

Figure A.26: Temperature and pressure histories for data-generation test OCZL#40 with pre-hydrided (390 wppm

to 1200 psig, oxidized at ≈1200°C to 12% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1300

1200

1100

1000

300 350 400 450 500 550 600 650 700 750

C) ° Temperature (

Time (s)

TC1: 57-mm above, 0°

TC2 (ctrl TC): 57-mm above, 180°

TC3: 44-mm below, 90°

TC4: 44-mm below, 270°

Pressure, psig

Rupture at 690±10°C

1196±17°C
A-27

Figure A.27: Temperature and pressure histories for data-generation test OCZL#41 with pre-hydrided (220 wppm

to 1200 psig, oxidized at ≈1200°C to 16% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1400

1300

1200

1100

1000

300 350 400 450 500 550 600 650 700 750 

C) ° Temperature (

Time (s)

TC1: 57-mm above, 0°

TC2 (ctrl TC): 57-mm above, 180°

TC3: 44-mm below, 90°

TC4: 44-mm below, 270°

Pressure, psig

Rupture at 730±32°C

1212±32°C
A-28

Figure A.28: Temperature and pressure histories for data-generation test OCZL#42 with pre-hydrided (520 wppm

to 1200 psig, oxidized at ≈1200°C to 14% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1400

1300

1200

1100

1000

300 350 400 450 500 550 600 650 700 750 80

C) ° Temperature (

Time (s)

TC1: 57-mm above, 0°

TC2 (ctrl TC): 57-mm above, 180°

TC3: 44-mm below, 90°

TC4: 44-mm below, 270°

Pressure, psig

Rupture at 682±5°C

1202±39°C
A-29

Figure A.29: Temperature and pressure histories for data-generation test OCZL#44 with pre-hydrided (700 wppm

to 1200 psig, oxidized at ≈1200°C to 14% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1400

1300

1200

1100

1000

300 350 400 450 500 550 600 650 700 

C) ° Temperature (

Time (s)

TC1: 57-mm above, 0°

TC2 (ctrl TC): 57-mm above, 180°

TC3: 44-mm below, 90°

TC4: 44-mm below, 270°

Pressure, psig

Rupture at 673±2°C

1202±22°C
A-30

Figure A.30: Temperature and pressure histories for data-generation test OCZL#45 with pre-hydrided (220 wppm

to 1200 psig, oxidized at ≈1200°C to 13% CP-ECR, and cooled with quench. 

900

800

700

600

500

400

300

200

1400

1300

1200

1100

1000

300 350 400 450 500 550 600 650 700 750 

C) ° Temperature (

Time (s)

TC1: 57-mm above, 0°

TC2 (ctrl TC): 57-mm above, 180°

TC3: 44-mm below, 90°

TC4: 44-mm below, 270°

Pressure, psig

Rupture at 737±11°C

1190±10°C
APPENDIX B: AXIAL PROFILES OF POST-RUPTURE 

CLADDING STRAINS

B-1

Figure B.1: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following thermal benchma

which was pressurized to 1200 psig, oxidized to 18% CP-ECR at ≈1200°C, and cooled without quenc

5

0

60

55

50

45

40

35

30

25

20

15

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

52%

33%
B-2

Figure B.2: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following scoping test OCZL

pressurized to 800 psig, oxidized to 16% CP-ECR at ≈1200°C, and cooled without quench.

0

60

50

40

30

20

10

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

D/D

Δ Strain (

o in %)

Distance from the Bottom (in.)

90° from Rupture

0° from Rupture

43%

19%
B-3

Figure B.3: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following ramp-to-rupture 

which was pressurized to 600 psig and cooled without quench.

0

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the Bottom (in.)

90° from Rupture

0° from Rupture 41%

18%
B-4

Figure B.4: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following ramp-to-rupture 

which was pressurized to 400 psig and cooled without quench.

0

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the Bottom (in.)

90° from Rupture

0° from Rupture

22%

45%
B-5

Figure B.5: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following ramp-to-rupture 

which was pressurized to 1600 psig and cooled without quench.

0

90

80

70

60

50

40

30

20

10

100

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

D/Do in %) 

Δ Strain (

Distance from the Bottom (in.)

90° from Rupture

0° from Rupture 80%

69%
B-6

Figure B.6: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following ramp-to-rupture 

which was pressurized to 1400 psig and cooled without quench.

0

80

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

60%

42%
B-7

Figure B.7: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following test OCZL#12, wh

to 1000 psig, oxidized to 15% CP-ECR, and cooled without quench.

0

80

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

49%

28%
B-8

Figure B.8: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following test OCZL#13, wh

to 1200 psig, oxidized to 14% CP-ECR, and cooled without quench.

0

80

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

54%

34%
B-9

Figure B.9: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

was pressurized to 1200 psig, oxidized to 18% CP-ECR, and cooled with quench.

0

80

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

63%

40%
B-10

Figure B.10: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

was pressurized to 1200 psig, oxidized to 18% CP-ECR, and cooled with quench.

0

80

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

66%

48%
B-11

Figure B.11: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

was pressurized to 1200 psig, oxidized to 14% CP-ECR, and cooled with quench.

0

80

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

53%

29%
B-12

Figure B.12: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

was pressurized to 1200 psig, oxidized to 13% CP-ECR, and cooled with quench.

0

80

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

50%

66%
B-13

Figure B.13: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

was pressurized to 1200 psig, oxidized to 12% CP-ECR, and cooled with quench.

0

80

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture 66%

41%
B-14

Figure B.14: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

was pressurized to 600 psig, oxidized to 17% CP-ECR, and cooled with quench.

0

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

38%

17%
B-15

Figure B.15: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

was pressurized to 600 psig, oxidized to 11% CP-ECR, and cooled with quench.

0

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

23%

47%
B-16

Figure B.16: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

was pressurized to 600 psig, oxidized to 12% CP-ECR, and cooled with quench.

0

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

43%

19%
B-17

Figure B.17: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

was pressurized to 1200 psig, oxidized to 16% CP-ECR, and cooled with quench.

0

80

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

46%

69%
B-18

Figure B.18: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

was pressurized to 1200 psig, oxidized to 17% CP-ECR, and cooled with quench.

0

80

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

47%

65%
B-19

Figure B.19: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

was pressurized to 1200 psig, oxidized to 16% CP-ECR, and cooled with quench. 

0

80

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

69%

45%
B-20

Figure B.20: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

was pressurized to 1200 psig, oxidized to 23% CP-ECR, and cooled with quench.

0

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

63%

41%
B-21

Figure B.21: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

was pressurized to 1200 psig, oxidized to 18% CP-ECR, and cooled with quench.

0

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

38%

61%
B-22

Figure B.22: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

wppm pre-hydriding, which was pressurized to 600 psig, oxidized to 13% CP-ECR, and cooled with 

0

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture 57%

38%
B-23

Figure B.23: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

wppm pre-hydriding, which was pressurized to 1200 psig, oxidized to 12% CP-ECR, and cooled with 

0

70

60

50

40

30

20

10

0 1 2 3 4 5 6 7 8 9 10 

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

26%

59%
B-24

Figure B.24: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

wppm pre-hydriding, which was pressurized to 1200 psig, oxidized to 16% CP-ECR, and cooled with 

0

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

36%

64%
B-25

Figure B.25: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

wppm pre-hydriding, which was pressurized to 1200 psig, oxidized to 14% CP-ECR, and cooled with 

0

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

47%

67%
B-26

Figure B.26: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

wppm pre-hydriding, which was pressurized to 1200 psig, oxidized to 14% CP-ECR, and cooled with 

0

70

60

50

40

30

20

10

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

D/Do in %)

Δ Strain (

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

42%

61%
B-27

Figure B.27: Axial profiles for minimum (90°) and maximum (0°) cladding diameters following data-generation te

wppm pre-hydriding, which was pressurized to 1200 psig, oxidized to 13% CP-ECR, and cooled with 

0

80

70

60

50

40

30

20

10

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

Strain (ΔD/Do in %)

Distance from the bottom (in.)

90° from Rupture

0° from Rupture

49%

80%

APPENDIX C: PHOTOGRAPHS OF POST-OXIDATION LOCA 

INTEGRAL SAMPLES

C-1

(a) 

(b)

Figure C.1: Post-oxidation images of OCZL#6 sample, which was pressurized to 1200 psig, oxidized to 18% CP-E

cooled without quench: (a) rupture-opening view and (b) side view.
C-2

(a) 

(b)

Figure C.2: Post-oxidation images of OCZL#7 sample, which was pressurized to 800 psig, oxidized to 16% CP-E

cooled without quench: (a) rupture-opening view and (b) side view.
C-3

(a) 

(b)

Figure C.3: Post-oxidation images of OCZL#8 sample, which was pressurized to 600 psig, heated to rupture, and 

quench: (a) rupture-opening view and (b) side view.
C-4

(a) 

(b)

Figure C.4: Post-oxidation images of OCZL#9 sample, which was pressurized to 400 psig, heated to rupture, and 

quench: (a) rupture-opening view and (b) side view.
C-5

(a) 

(b)

Figure C.5: Post-oxidation images of OCZL#10 sample, which was pressurized to 1600 psig, heated to rupture, a

quench: (a) low magnification and (b) higher magnification.
C-6

(a) 

(b)

Figure C.6: Post-oxidation images of OCZL#11 sample, which was pressurized to 1400 psig, heated to rupture, a

quench: (a) low magnification and (b) higher magnification.
C-7

(a) 

(b)

Figure C.7: Post-oxidation images of OCZL#12 sample, which was pressurized to 1000 psig, heated to rupture to 

15% CP-ECR, and cooled without quench: (a) low magnification and (b) higher magnification.
C-8

(a) 

(b)

Figure C.8: Post-oxidation images of OCZL#13 sample, which was pressurized to 1200 psig, heated to 1200°C, ox

ECR, and cooled without quench: (a) low magnification and (b) higher magnification. 
C-9

(a) 

(b)

Figure C.9: Post-oxidation images of OCZL#14 sample, which was pressurized to 1200 psig, heated to 1200°C, ox

ECR, and cooled with quench: (a) low magnification and (b) higher magnification. 
C-10

(a) 

(b)

Figure C.10: Post-oxidation images of OCZL#15 sample, which was pressurized to 1200 psig, heated to 1200°C, ox

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-11

(a) 

(b)

Figure C.11: Post-oxidation images of OCZL#17 sample, which was pressurized to 1200 psig, heated to 1200°C, ox

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-12

(a) 

(b)

Figure C.12: Post-oxidation images of OCZL#18 sample, which was pressurized to 1200 psig, heated to 1200°C, ox

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-13

(a) 

(b)

Figure C.13: Post-oxidation images of OCZL#19 sample, which was pressurized to 600 psig, heated to 1200°C, oxi

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-14

(a) 

(b)

Figure C.14: Post-oxidation images of OCZL#21 sample, which was pressurized to 600 psig, heated to 1200°C, oxi

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-15

(a) 

(b)

Figure C.15: Post-oxidation images of OCZL#22 sample, which was pressurized to 600 psig, heated to 1200°C, oxi

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-16

(a) 

(b)

Figure C.16: Post-oxidation images of OCZL#24 sample, which was non-pressurized, heated to 1200°C, one-sided 

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-17

(a) 

(b)

Figure C.17: Post-oxidation images of OCZL#25 sample, which was pressurized to 1200 psig, heated to 1200°C, ox

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-18

(a)

(b)

Figure C.18: Post-oxidation images of OCZL#26 sample, which was pressurized to 1200 psig, heated to 1200°C, ox

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-19

(a) 

(b)

Figure C.19: Post-oxidation images of OCZL#27 sample, which was pressurized to 1200 psig, heated to 1200°C, ox

ECR, and cooled with quench: (a) low magnification and (b) higher magnification. 
C-20

(a) 

(b)

Figure C.20: Post-oxidation images of OCZL#29 sample, which was pressurized to 1200 psig, heated to 1200°C, ox

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-21

(a) 

(b)

Figure C.21: Post-oxidation images of OCZL#32 sample, which was pressurized to 1200 psig, heated to 1200°C, ox

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-22

(a) 

(b)

Figure C.22: Post-oxidation images of OCZL#36 sample, which was pressurized to 1200 psig, heated to 1200°C, ox

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-23

(a) 

(b)

Figure C.23: Post-oxidation images of OCZL#37 sample, which was pressurized to 1200 psig, heated to 1200°C, ox

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-24

(a) 

(b)

Figure C.24: Post-oxidation images of OCZL#43 sample, which was pressurized to 1200 psig, heated to 1200°C, ox

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-25

(a)

(b)

Figure C.25: Post-oxidation images of OCZL#39PH sample (660 wppm H), which was pressurized to 600 psig, hea

oxidized to 13% CP-ECR, and cooled with quench: (a) low magnification and (b) higher magnificati
C-26

(a)

(b)

Figure C.26: Post-oxidation images of OCZL#40 sample (390 wppm H), which was pressurized to 1200 psig, heate

to 12% CP-ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-27

(a) 

(b)

Figure C.27: Post-oxidation images of OCZL#41 sample (220 wppm H), which was pressurized to 1200 psig, heate

to 16% CP-ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-28

(a) 

(b)

Figure C.28: Post-oxidation images of OCZL#42 sample (520 wppm H), which was pressurized to 1200 psig, heate

to 14% CP-ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-29

(a) 

(b)

Figure C.29: Post-oxidation images of OCZL#44 sample (700 wppm H), which was pressurized to 1200 psig, heate

to 14% CP-ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
C-30

(a) 

(b)

Figure C.30: Post-oxidation images of OCZL#45 sample (220 wppm H), which was pressurized to 1200 psig, heate

to 13% CP-ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
APPENDIX D: 4-PBT LOAD-DISPLACEMENT CURVES

D-1

Figure D.1: 4-PBT load-displacement curve for OCZL#8 test sample, which was pressurized to 600 psig, heated 

without quench.

0

900

800

700

600

500

400

300

200

100

1000

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

Load (N)

Displacement (mm)

K = 136 N/mm

7.7 mm

Mmax = 20.9 N•m

Emax = 8.4 J

Pmax = 835 
D-2

Figure D.2: 4-PBT load-displacement curve for OCZL#9 test sample, which was pressurized to 400 psig, heated 

without quench.

0

900

800

700

600

500

400

300

200

100

1000

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

Load (N)

Displacement (mm)

K = 133 N/mm

7.8 mm

Mmax = 20.6 N•m

Emax = 8.3 J

Pmax = 82
D-3

Figure D.3: 4-PBT load-displacement curve for OCZL#10 test sample, which was pressurized to 400 psig, heated 

cooled without quench.

0

900

800

700

600

500

400

300

200

100

1000

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

Load (N)

Displacement (mm)

K = 120 N/mm

7.1 mm

Mmax = 19.7 N•m

Emax = 7.7 J

Pmax = 788 
D-4

Figure D.4: 4-PBT load-displacement curve for OCZL#12 test sample, which was pressurized to 1000 psig, heate

to 15% CP-ECR, and cooled without quench. 

0

500

400

300

200

100

0 1 2 3 

Load (N)

Displacement (mm)

K = 141 N/mm

0.37 mm

Mmax = 10.5 N•m

Emax = 0.78 J

Pmax = 42

Corrected Offset Displacement = 0.33 mm
D-5

Figure D.5: 4-PBT load-displacement curve for OCZL#13 test sample, which was pressurized to 1200 psig, heate

to 14% CP-ECR, and cooled without quench. 

0

500

400

300

200

100

0 1 2 3 

Load (N)

Displacement (mm)

K = 123 N/mm

0.19 mm

Mmax = 8.8 N•m

Emax = 0.58 J

Pmax = 352 N

Corrected Offset Displacement = 0.13 mm
D-6

Figure D.6: 4-PBT load-displacement curve for OCZL#14 test sample, which was pressurized to 1200 psig, heate

to 18% CP-ECR, and cooled with quench. 

0

50

300

250

200

150

100

0.0 0.5 1.0 1.5 2.0 2.5 

Load (N)

Displacement (mm)

K = 115 N/mm

Mmax = 5.7 N•m

Emax = 0.24 J

Pmax = 227 N
D-7

Figure D.7a: 4-PBT load-displacement curve for OCZL#15 test sample, which was pressurized to 1200 psig, heate

to 18% CP-ECR, and cooled with quench. Sample was subjected to reverse bending with rupture re

compression. Test was interrupted following non-linear displacement to examine sample.

0

400

300

200

100

0 1 2 3 4 

Load (N)

Displacement (mm)

K = 124 N/mm

0.95 mm

Mmax = 9.0 N•m

E1 = 0.39 J

Pmax = 360 
D-8

Figure D.7b: Continuation of 4-PBT load-displacement curve for OCZL#15 test sample, which was pressurized to 

1200°C, oxidized to 18% CP-ECR, and cooled with quench. Sample was subjected to reverse bendin

under maximum compression. Results are for reloading sample.

0

400

300

200

100

0 2 4 6 8 10 12 

Load (N)

Displacement (mm)

K = 95 N/mm

9.1 mm

Mmax = 8.9 N•m

E2 = 1.46 J

E1 = 0.39 J

Emax > 1.9 J

Pmax = 357 N

Total Offset Displacement >10 
D-9

Figure D.8: 4-PBT load-displacement curve for OCZL#17 test sample, which was pressurized to 1200 psig, heate

to 13% CP-ECR, and cooled with quench. 

0

400

300

200

100

0 1 2 3 

Load (N)

Displacement (mm)

K = 130 N/mm

Mmax = 8.4 N•m

Emax = 0.65 J

Pmax = 335 N
D-10

Figure D.9: 4-PBT load-displacement curve for OCZL#18 test sample, which was pressurized to 1200 psig, heate

to 12% CP-ECR, and cooled with quench. 

0

600

500

400

300

200

100

0 1 2 3 4 5 

Load (N)

Displacement (mm)

 KL1 = 130 N/mm

KL2 = 126 N/mm

 

Mmax = 13.5 N•m

Emax = 1.28 J

Pmax = 54
D-11

Figure D.10: 4-PBT load-displacement curve for OCZL#19 test sample, which was pressurized to 600 psig, heated 

to 17% CP-ECR, and cooled with quench. 

0

50

300

250

200

150

100

0 1 2 

Load (N)

Displacement (mm)

K = 124 N/mm

Mmax = 5.68 N•m

Emax = 0.23 J

Pmax = 227 N
D-12

Figure D.11: 4-PBT load-displacement curve for OCZL#21 test sample, which was pressurized to 600 psig, heated 

to 11% CP-ECR, and cooled with quench. 

0

600

500

400

300

200

100

0 1 2 3 4 5 

Load (N)

Displacement (mm)

K = 129 N/mm

Mmax = 13.8 N•m

Emax = 1.17 J

Pmax = 552 N
D-13

Figure D.12: 4-PBT load-displacement curve for OCZL#22 test sample, which was pressurized to 600 psig, heated 

to 12% CP-ECR, and cooled with quench. 

0

500

400

300

200

100

0 1 2 3 4 

Load (N)

Displacement (mm)

K = 120 N/mm

Mmax = 11.0 N•m

Emax = 0.83 J

Pmax = 440 N
D-14

Figure D.13: 4-PBT load-displacement curve for OCZL#24 test sample, which was non-pressurized, heated to ≈12

oxidized to 17% CP-ECR, and cooled with quench. 

0

900

800

700

600

500

400

300

200

100

1100

1000

0 1 2 3 4 5 6 7 

Load (N)

Displacement (mm)

 

Mmax = 24.1 N•m

Emax = 4.15 J

0.43 mm

K = 122 N/mm

Pmax = 962 N
D-15

Figure D.14: 4-PBT load-displacement curve for OCZL#25 test sample, which was pressurized to 1200 psig, heate

to 16% CP-ECR, and cooled with quench. 

0

50

400

350

300

250

200

150

100

0 1 2 3 

Load (N)

Displacement (mm)

K = 111 N/mm

Mmax = 8.32 N•m

Emax = 0.50 J

Pmax = 333 N
D-16

Figure D.15a: 4-PBT load-displacement curve for OCZL#29 test sample, which was pressurized to 1200 psig, heate

to 17% CP-ECR, and cooled with quench. Test was interrupted to allow examination of sample.

0

50

200

150

100

0 1 2 3 

Load (N)

Displacement (mm)

K = 118 N/mm

Mmax = 4.69 N•m

E1 = 0.27 J

E2 = 0.13 J

Pmax = 188 N
D-17

Figure D.15b: 4-PBT load-displacement curve for OCZL#29 test sample, which was pressurized to 1200 psig, heate

to 17% CP-ECR, and cooled with quench. Sample was reloaded after visual examination.

0

50

200

150

100

0 1 2 3 4 5 6 7 8 9 1

Load (N)

Displacement (mm)

E2 = 0.13 J
D-18

Figure D.16: 4-PBT load-displacement curve for OCZL#32 test sample, which was pressurized to 1200 psig, heate

to 16% CP-ECR, and cooled with quench. 

0

50

300

250

200

150

100

0 1 2 

Load (N)

Displacement (mm)

Pmax = 267 N

K = 137 N/mm

Mmax = 6.67 N•m 

Emax = 0.26 J
D-19

Figure D.17: 4-PBT load-displacement curve for OCZL#36 test sample, which was pressurized to 1200 psig, heate

to 18% CP-ECR, and cooled with quench. 

0

500

400

300

200

100

0 1 2 3 

Load (N)

Displacement (mm)

K = 134 N/mm

Mmax = 9.47 N•m

Emax = 0.56 J

Pmax = 379 N
D-20

Figure D.18: 4-PBT load-displacement curve for OCZL#37 test sample, which was pressurized to 1200 psig, heate

to 23% CP-ECR, and cooled with quench. 

0

50

200

150

100

0.0 0.5 1.0 1.5 

Load (N)

Displacement (mm)

K = 111 N/mm

Mmax = 4.14 N•m 

Emax = 0.13 J

Pmax = 166 N 
D-21

Figure D.19: 4-PBT load-displacement curve for OCZL#43 test sample, which was pressurized to 1200 psig, heate

to 18% CP-ECR, and cooled with quench. 

0

50

300

250

200

150

100

0 1 2 

Load (N)

Displacement (mm)

Emax = 0.24 J 

Mmax = 6.12 N•m 

KL1 = 132 N/mm

KL2 = 128 N/mm

Pmax = 245 N 
D-22

Figure D.20: 4-PBT load-displacement curve for OCZL#39 (660 wppm H) test sample, which was pressurized to 6

≈1200°C, oxidized to 13% CP-ECR, and cooled with quench. 

0

50

200

150

100

0.0 0.5 1.0 1.5 

Load (N)

Displacement (mm)

Mmax = 2.80 N•m 

Emax = 0.05 J

KL1 = 133 N/mm

KL2 = 138 N/mm

Pmax = 112 N
D-23

Figure D.21: 4-PBT load-displacement curve for OCZL#40 (390 wppm H) test sample, which was pressurized to 1

≈1200°C, oxidized to 12% CP-ECR, and cooled with quench. 

0

50

300

250

200

150

100

0.0 0.5 1.0 1.5 

Load (N)

Displacement (mm)

KL1 = 114 N/mm

KL2 = 128 N/mm

Mmax = 4.89 N•m

Emax = 0.16 J

Pmax = 199 N
D-24

Figure D.22: 4-PBT load-displacement curve for OCZL#41 (220 wppm H) test sample, which was pressurized to 1

≈1200°C, oxidized to 16% CP-ECR, and cooled with quench. 

0

50

200

150

100

0.0 0.5 1.0 1.5 

Load (N)

Displacement (mm)

K = 128 N/mm

Mmax = 3.21 N•m

Emax = 0.07 J

Pmax = 128 N
D-25

Figure D.23a: 4-PBT load-displacement curve for OCZL#44 (700 wppm H) test sample, which was pressurized to 1

≈1200°C, oxidized to 14% CP-ECR, and cooled with quench. Test was interrupted to allow for visua

0

75

50

25

100

0.0 0.5 1.0 1.5 

Load (N)

Displacement (mm)

KL1 = 110 N/mm

KL2 = 110 N/mm

Mmax = 2.35 N•m

E1 = 0.062 J

E2 = 0.049 J

Emax = 0.11 J

Pmax = 94 N
D-26

Figure D.23b: 4-PBT load-displacement curve for OCZL#44 (700 wppm H) test sample, which was pressurized to 1

≈1200°C, oxidized to 14% CP-ECR, and cooled with quench. Sample was reloaded. 

0

75

50

25

100

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 

Load (N)

Displacement (mm)

E2 = 0.049 J
D-27

Figure D.24: 4-PBT load-displacement curve for OCZL#45 (220 wppm H) test sample, which was pressurized to ≈

≈1200°C, oxidized to 13% CP-ECR, and cooled with quench.

0

50

300

250

200

150

100

0 1 2 

Load (N)

Displacement (mm)

KL1 = 131 N/mm

Mmax = 6.59 N•m

Emax = 0.28 J

KL2 = 124 N/mm

Pmax = 263 N

APPENDIX E: PHOTOGRAPHS OF POST-BEND LOCA 

INTEGRAL SAMPLES

E-1

Figure E.1: Post-4-PBT images of ramp-to-rupture samples from OCZL#8, #9, and #10 test samples, which were 

to 600, 400, and 1600 psig, respectively, heated to rupture, and cooled without quench. 
E-2

(a)

(b)

Figure E.2: Post-4-PBT images of OCZL#12 test sample, which was pressurized to 1000 psig, heated to ≈1200°C

ECR, and cooled without quench: (a) low magnification and (b) higher magnification. 
E-3

Figure E.3: Post-4-PBT image of OCZL#13 test sample, which was pressurized to 1200 psig, heated to ≈1200°C, 

ECR, and cooled without quench.

Figure E.4: Post-4-PBT images of OCZL#14 test sample, which was pressurized to 1200 psig, heated to ≈1200°C

ECR, and cooled with quench. 
E-4

(a)

(b)

Figure E.5: Post-4-PBT images for the OCZL#15 test sample, which was pressurized to 1200 psig, heated to ≈120

CP-ECR, cooled with quench, and subjected to reverse bending: (a) low magnification and (b) highe
E-5

(a)

(b)

Figure E.6: Post-4-PBT images of OCZL#17 test sample, which was pressurized to 1200 psig, heated to ≈1200°C

ECR, and cooled with quench: (a) side view and (b) rupture view.
E-6

(a)

(b)

Figure E.7: Post-4-PBT images of OCZL#18 test sample, which was pressurized to 1200 psig, heated to ≈1200°C

ECR, and cooled with quench: (a) side view and (b) rupture view.
E-7

(a)

(b)

Figure E.8: Post-4-PBT images of OCZL#19 test sample, which was pressurized to 600 psig, heated to ≈1200°C, 

ECR, and cooled with quench: (a) low magnification and (b) higher magnification. 
E-8

(a)

(b)

Figure E.9: Post-4-PBT images of OCZL#21 test sample, which was pressurized to 600 psig, heated to ≈1200°C, 

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
E-9

(a)

(b)

Figure E.10: Post-4-PBT images of OCZL#22 test sample, which was pressurized to 600 psig, heated to ≈1200°C, 

ECR, and cooled with quench: (a) low magnification and (b) high magnification. 
E-10

(a)

(b)

Figure E.11: Post-4-PBT images of OCZL#24 test sample, which was non-pressurized, heated to ≈1200°C, one-sid

CP-ECR, and cooled with quench: (a) low magnification and (b) higher magnification. 
E-11

 

(a)

(b)

Figure E.12: Post-4-PBT images of OCZL#25 test sample, which was pressurized to 1200 psig, heated to ≈1200°C

ECR, and cooled with quench: (a) low magnification and (b) higher magnification.
E-12

(a)

(b)

Figure E.13a: Post-4-PBT images of OCZL#29 test sample, which was pressurized to 1200 psig, heated to ≈1200°C

ECR, and cooled with quench. Images were taken after first loading: (a) side A and (b) side B.
E-13

Figure E.13b: Post-4-PBT image of OCZL#29 test sample, which was pressurized to 1200 psig, heated to ≈1200°C, 

ECR, and cooled with quench. Image was taken after second loading. 
E-14

(a)

(b)

Figure E.14: Post-4-PBT Images of OCZL#32 test sample, which was pressurized to 1200 psig, heated to ≈1200°C

ECR, and cooled with quench: (a) low magnification and (b) higher magnification. 
E-15

(a)

(b)

Figure E.15: Post-4-PBT images of OCZL#36 test sample, which was pressurized to 1200 psig, heated to ≈1200°C

ECR, and cooled with quench: (a) low magnification and (b) high magnification. 
E-16

(a)

(b)

Figure E.16: Post-4-PBT images of OCZL#37 test sample, which was pressurized to 1200 psig, heated to ≈1200°C

ECR, and cooled with quench: (a) low magnification and (b) high magnification. 
E-17

(a)

(b)

Figure E.17: Post-4-PBT images of OCZL#43 test sample, which was pressurized to 1200 psig, heated to 1200°C, 

ECR, and cooled with quench: (a) low magnification and (b) higher magnification. 
E-18

(a)

(b)

Figure E.18: Post-4-PBT images of OCZL#39 (660 wppm H) test sample, which was pressurized to 600 psig, heate

oxidized to 13% CP-ECR, and cooled with quench: (a) low magnification and (b) higher magnificati
E-19

(a)

(b)

Figure E.19: Post-4-PBT images of OCZL#40 (390 wppm H) test sample, which was pressurized to 1200 psig, hea

oxidized to 12% CP-ECR, and cooled with quench: (a) low magnification and (b) higher magnificati
E-20

(a)

(b)

Figure E.20: Post-4-PBT images of OCZL#41 (220 wppm H) test sample, which was pressurized to 1200 psig, hea

oxidized to 16% CP-ECR, and cooled with quench: (a) low magnification and (b) higher magnificati
E-21

(a)

(b)

Figure E.21: Post-4-PBT images of OCZL#44 (700 wppm H) test sample, which was pressurized to 1200 psig, hea

oxidized to 14% CP-ECR, and cooled with quench: (a) low magnification and (b) higher magnificati
E-22

(a)

(b)

Figure E.22: Post-4-PBT images of OCZL#45 (220 wppm H) test sample, which was pressurized to 1200 psig, hea

oxidized to 13% CP-ECR, and cooled with quench: (a) low magnification and (b) higher magnificati




 

NUREG/CR-7219 Cladding Behavior during Postulated Loss-of-Coolant Accidents