APPENDIX D

CATEGORY D 

SEPARATE EFFECT TESTING 

PHENOMENA DESCRIPTIONS AND RATIONALES FOR IMPORTANCE 

RANKING, APPLICABILITY, AND UNCERTAINTY 

This appendix provides a description for each phenomenon appearing in Table 

3-4, Separate Effect Testing PIRT. Entries in the Table D-i, columns 1 and 2, 

follow the same order as in Table 3-5. Table D-1, column 3, also documents the 

PIRT-panel developed rationales for three types of Panel findings. 

First, rationales are provided for the importance (High, Medium, or Low) 

assigned by the panel to each phenomenon. Because importance ranking was 

established by a vote of the panel members, a rationale is provided whenever one 

or more panel members voted a particular rank, i.e., High, Medium or Low. If 

there were no votes for a given importance rank, "No votes" is entered. 

Second, the PIRT panel considered the applicability of the baseline PIRT to a 

broader set of circumstances, e.g., different fuel arrays, cladding types, reactor 

types, and burnups to 75 GWd/t. The specific question addressed by the PIRT 

panel was as follows: "Could the importance ranking assigned for the given 

phenomenon in the baseline PIRT be for different for other fuel arrays, cladding 

types, reactor types, or burnups?" If this question is answered with a "no", the 

following entry appears in Table C-1: "Baseline PIRT importance rank is 

applicable." If this question is answered with a "yes", the rationale is entered. 

Additional details are presented in the footnotes to Table 3-5. 

Third, the PIRT panel considered the current state of knowledge or uncertainty 

regarding each phenomenon. The phenomenon is characterized as "known (K)" 

if approximately 75-100% of full knowledge and understanding of the 

phenomenon exists. The phenomenon is characterized as "partially known (PK)" 

if between 25-75% of full knowledge and understanding of the phenomenon 

exists. The phenomenon is characterized as "unknown (UK)"if less than 25% of 

full knowledge and understanding of the phenomenon exists. Because the 

uncertainty ranking was established by a vote of the panel members, a rationale is 

provided whenever one or more panel members voted a particular uncertainty, 

i.e., known, partially known, or unknown. If there were no votes for a given 

uncertainty level, "No votes" is entered 

There were several phenomena for which no importance rank was recorded. In 

such cases "No rationale recorded" is entered.

Appendix D-1
Subcategory (Test type)

Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing 

Phenomena (Parameter) I Definitinn and Raihnnf- 1- .I.. A 1 .1

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15)t 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Specimen selection: 

Alloy type

Composition or designation of the metal utilized in fuel-rod fabrication 

H(3) Data (B&W-10227) show that the 2nd layer develops differently on different 

alloys. Initial oxide layer may be different between alloys and thus behave 

differently. 

M(2) Similar to rationale for high but the oxide kinetics do not change and the other 

differences may not affect brittleness. 

L(O) No votes.

Fuel: N 

Clad: NA 

Reactor: N 

Burnup: N 

K(2): Data, judgement 

PK(2): Judgement, no data 

UK(1): Data, judgement

t The Separate Effect Testing category was developed by considering the types of separate effect experiments that might be conducted to develop needed data. The panel defined six test types and the phenomena associated with each. Prior to voting on the phenomena themselves, the panel voted on the importance of each 

test type.

Appendix D-2

mror axwe, Ep ca !!!L, anu uncertainty)
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15) 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Specimen selection: 

Thickness and 

morphology of pre

existing oxide

The total amount of oxide formed on the cladding and whether the oxidation is uniform 

or nodular, and whether there is extensive cracking and spalling. 

H(1) Thickness and morphology controls passage to the metal, even though the 

oxidation rates are the same. 

M(2) Hydrogen pickup during corrosion may affect oxide rates. Some early 

irregularities in oxide rate data exist and may be due to corrosion layer. 

L(2) Data (French and Japan) show that only the thin dense oxide layer controls 

the oxidation rate at high temperature and this layer is independent of the 

initial oxide. 

Fuel: N 

Clad: NA 

Reactor: N 

Bumup: N

K(2): 

PK(3): 

UK(O):

Data 

Data (incomplete) 

No votes

Appendix D-3
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15) 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Specimen selection: 

Bumup, including fluence

The amount of burnup to which the fuel rod used for the specimen was exposed. 

H(3) Bumup per se may not be so important but there could be effects such as 

precipitate dissolution for which testing is needed. 

M(1) Burnup per se may not be so important but there could be effects such as 

precipitate dissolution for which testing is needed, but the effect is not 

expected to be so pronounced. Discovery of unknown effects may occur if testing 

takes place. 

L(1) Irradiation damage is expected to be annealed out.

Fuel: 

Clad: 

Reactor: 

Bumup: 

K(1): 

PK(4): 

UK(0):

N 

N 

N 

NA 

Data 

Data, judgement 

No votes

Appendix D-4
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15) 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Specimen selection: 

Pre-existing hydrogen 

content and distribution

Amount and distribution of hydrogen associated with fuel rod clad segment. This 

hydrogen may be in solution in the metal or may exist combined with the metal as a 

discrete hydride phase. 

H(1) Affects oxygen repartition during oxidation and oxygen solubility in the beta 

phase. 

M(3) Initial amount of hydrogen has a slight impact on the cladding strain during 

the ballooning phase and a very low impact on oxidation at high temperature 

and behavior upon quench; the initial H distribution has no impact. High 

concentration of hydrogen stabilizes beta phase and is conducive to thicker 

layer of load-bearing prior beta phase. 

L(1) The available testing by the Japanese and French indicate a relatively minor 

effect of hydrogen content on nonirradiated or irradiated cladding high 

temperature oxidation.

Fuel: N 

Clad: NA 

Reactor: N 

Bumup: NA 

K(2): Data 

PK(2): Data, judgement 

UK(O): No votes

Appendix D-5
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15) 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Conduct of Test-During 

Oxygen potential

Related to partial pressure of oxygen in system. 

H(3) The oxygen availability directly controls the alpha and oxide layer 

developments and the hydrogen pickup. Condition of steam starvation must be 

avoided to have a valid test. The oxygen potential is the boundary condition 

that determines oxidation rate. 

M(1) The development of an appropriate environment is essential to obtaining 

meaningful results. However, the measurement of oxygen potential may not be 

necessary to ensure a prototypical environment. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(1): 

PK(2): UK(0):

N 

N 

N 

N 

Data. Judgement 

Data, Judgement 

No votes.

Appendix D-6
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15) 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Conduct of Test-During 

Temperature and time

Measurement of the time-varying temperature. 

H(5) May affect oxygen distribution and hydrogen pickup. Key parameter needed to 

analyze the tests. Since the corrosion process is such a strong function of 

temperature, accurate measurement of the temperature during the test is 

essential to providing meaningful results.

M(O) 

L(O)

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(2): 

PK(2): 

UK(O):

No votes. 

No votes.

N 

N 

N 

N 

Data, Judgement 

Data, Judgement

Appendix D-7
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Imivortance. Aronliabililtv and 1Jncprtalntv•

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15) 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Conduct of Test-During 

Total steam pressure

Total steam pressure 

H(1) Recent Korean testing on oxide at temperatures below 1800 F shows that the 

impact of pressure on oxidation rate could be important in small break LOCA 

M(3) Available experimental results do not show large effects for low burnup 

Zircaloy; to be confirmed at high burnup and for alternative cladding alloys. 

Not a key parameter but should be measured to check the experimental 

conditions are the same. Available testing information indicates measurable 

differences in high temperature cladding corrosion rate for different steam 

pressures. 

L(1) As long as steam starvation is prevented, total steam pressure is less important. 

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(3): 

PK(1): 

UK(0):

Data, Judgement 

Data, Judgement 

No votes.

Appendix D-8

S.. . , a.. . " . . . . . ." -rIL .. . . . . . . . . . . .
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15) 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Conduct of Test-During 

Weight gain

On-line measurement of the gain in mass of the test specimen. Used to measure the total 

amount of oxygen absorbed during the high temperature oxidation phase. 

H(4) Very important parameter to measure because it is directly linked with the 

equivalent cladding reacted (ECR) used for a LOCA criterion. Less accurate 

than direct 0 measurement (integrate Temperature distribution effects and end 

effects) but very useful in relative. 

M(1) Weight gain is a primary measure of corrosion, however, the more important 

measure in this case would be performed by metallography (differentiation 

between oxide, oxygen-stabilized alpha-phase, and remaining beta phase). 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(5): 

PK(O): 

UK(O):

N 

N 

N 

N 

Data, Experience, Judgement 

No votes. 

No votes.

Appendix D-9
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter)_] Definition and Rationale (Importance, Applicability, and Uncertainty)

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15) 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Conduct of Test-During 

Steam consumption

On-line measurement of the level of steam consumption during the LOCA transient

H(1) This is a primary indicator of oxidation.

M(2) Crosscheck measurement for the weight gain. Useful data but less precise than 

the post test measurement when the high temperature oxidation is non

uniform along the rod. 

,L(2) The use of steam consumption was suggested as an alternate, independent 

assessment of oxygen absorbed, however, the more direct measurement is 

performed with metallography and weight gain. Steam consumption is not 

easily measured and is not necessarily an accurate indication of the degree of 

total oxidation (because of steam condensation). 

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N

K(3): 

PK(2): 

UK(0):

Data, Experience, Judgement 

Judgement 

No votes.

Appendix D-10
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15) 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Conduct of Test-During 

One-sided vs. two-sided

Whether oxidation is allowed on both the interior and exterior surfaces of the tube to 

model both oxidation processes as will occur during an accident. 

H(2) Two-sided oxidation is important in order to reproduce the specific oxidation 

conditions inside the balloon (stagnant steam conditions inducing a higher 

hydrogen pickup). A true simulation of the actual condition is essential to 

avoid overlooking unanticipated effects; such as the role of inner surface fission 

products, presence of a zirconium liner, or other possible differences. 

M(3) Both kind of tests are doable. The analysis should take into account the test 

conditions. Two-sided oxidation may have slightly different kinetics; at least 

it is worth investigating. 

L(1) High-temperature oxidation is controlled primarily by the process of oxygen 

transport across the oxide layer. 

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N 

K(2): Data 

PK(3): Data, Judgement: Prior work has shown, for example, that the presence of a 

zirconium liner does not significantly affect the cladding high temperature 

corrosion behavior. However, all possible effects have not necessarily been 

quantified. 

UK(O): No votes.

Appendix D-11
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15) 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Conduct of Test-PTE 

Oxide thickness

Post test measurement of the oxide thickness that developed during the high 

temperature phase of the LOCA transient. 

H(5) One of the most pertinent measurements. Oxide thickness is the most important 

parameter of oxidation. Key parameter for the interpretation and analysis of 

the test. Posttest metallography is the primary and most reliable 

quantification, however the measurement is not necessarily just oxide 

thickness, also region of oxygen-stabilized alpha layer and remaining beta 

layer. Useful parameter to measure primary item studied. 

M(O) No votes. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(3): 

PK(1): UK(0):

N 

N 

N 

N 

Data, Experience 

Data, Judgement 

No votes.

Appendix D-12
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15) 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Conduct of Test-PTE 

Characteristic ct-0 

morphology

Measurement of the thickness and morphology of the different metallurgical layers that 

formed during the LOCA transient. 

H(5) Key parameter to analyze the test. Alpha and beta layer thicknesses and the 

degree of alpha "incursion" are important oxidation parameters that influence 

the mechanical properties of the cladding. Relative amounts of alpha and 

beta determine the behavior of oxidized clad. Note: Posttest metallography 

is the primary and most reliable quantification, however the measurement is 

not necessarily just oxide thickness, also region of oxygen-stabilized alpha 

layer and remaining beta layer. 

M(O) No votes. 

L(O) No votes.

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N 

K(3): Da 

PK(2): Da 

UK(O): Nc

ta, Experience, Judgement 

ta, Judgement 

votes.

Appendix D-13
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15) 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Conduct of Test-PTE 

Oxygen distribution

Distribution of oxygen either dissolved or existing as an oxide phase in the cladding. 

H(5) Oxygen content in the beta phase is important for embrittlement. Key 

parameter that governs the behavior upon quench of the cladding. Posttest 

metallography is the primary and most reliable quantification, however the 

measurement is not necessarily just oxide thickness, also region of oxygen

stabilized alpha layer and remaining beta layer. Distribution of oxygen in 

prior beta layers is an important oxidation parameter that strongly influences 

the mechanical properties. 

M(O) No votes. 

L(O) No votes. 

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(2): 

PK(2): UK(0):

Data, Experience, Judgement 

Data, Judgement 

No votes.

Appendix D-14
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategorv (Test type) I Phenomena (Parameter) - ---.- - r_ - -I LL _ it Z

I Definition and Ratinn1Tnnnrhni'9 Ann1it'2h1i4#, 2nA TT.4..4.A

Oxidation rate, oxygen 

distribution, effect of 

chemistry on solubility 

(vote=15) 

Separate effect test to 

measure the steam 

oxidation kinetics at 

high temperature in 

Zirconium alloys used for 

cladding.

Conduct of Test-PTE 

Hydrogen pickup and 

distribution

Amount and distribution of H absorbed by the cladding during the LOCA transient. 

H(4) Hydrogen pickup in the beta phase is important for brittleness. The amount of 

H usually picked-up is small. It can be locally very high in case of early 

failure of the cladding during the ballooning phase followed with steam 

ingress. Hydrogen content and distribution in the prior beta layer are 

important parameters that influence the mechanical properties of the 

cladding. 

M(2) Prior work by the Japanese and French have shown the pre- and post-test 

hydrogen content and distribution to be not so influential for practically 

achievable hydrogen levels, however, the amount of hydrogen absorbed during the test is an item of active interest and deserves characterization. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(3): 

PK(2): UK(0):

N 

N 

N 

Y(1): More important for high-bumup fuel 

Data, Judgement 

Data, Judgement 

No votes.

Appendix D-15
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Specimen selection: 

Hydrogen content and 

distribution

Amount and distribution of hydrogen associated with fuel rod clad segment. This 

hydrogen may be in solution in the metal or may exist combined with the metal as a 

discrete hydride phase. 

H(3) Affects oxygen solubility in the beta phase and post-quench ductility. Hydride 

dissolves during oxidation at high temperatures and most hydrogen atoms are 

concentrated in the beta phase. Hydrogen content and distribution in the 

transformed beta phase are important parameters that influence clad 

resistance to thermal-shock failure. 

M(2) Data show low impact of H on clad behavior upon quench. Prior testing has 

demorrstrated that quench behavior is not significantly affected by the amount 

of prior hydrogen (oxygen embrittlement is more important), distribution of 

prior hydrogen (homogenization occurs during the high temperature period), or 

hydrogen absorbed during the high temperature oxidation reaction (hydrogen 

absorption is minimal). However, this is an item of active interest and 

deserves a characterization. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(2): 

PK(2): 

UK(O):

N 

N 

N 

N 

Data 

Data, Judgement 

No votes.

Appendix D-16
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Specimen selection: 

Alloy type

Composition or designation of the metal utilized in fuel-rod fabrication 

H(2) May affect oxygen distribution and hydrogen pickup. Stability of beta phase 

and mechanical properties of the load-bearing prior beta layer are 

significantly influenced by the addition of Nb, therefore, thermal-shock 

resistance of M5 and Zirlo is expected to differ from that of Zircaloy. 

M(2) In general, differences among the characterized zirconium based materials 

have shown differences in high-temperature oxidation and quench behavior to 

be relatively minor. However, specific details on the newer materials are not 

available to the reviewer and so bets are hedged. 

L(1) Data show no significant impact of the Alloy type on the behavior upon quench

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(1): 

PK(2): UK(1):

Y(1): M5- or Zirlo-clad fuels 

N 

Y(1): PWR 

N 

Data 

Data 

Judgement

Appendix D-17
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Specimen selection: 

Thickness and 

morphology of pre

existing oxide

The total amount of oxide formed on the cladding during in-reactor operation and 

whether the oxidation is uniform or nodular, and whether there is extensive cracking and spalling. 

H(3) These parameters influence the degree of transient oxidation and hydrogen uptake, the two properties that strongly influence clad resistance to thermal 

shock. It is important to use representative cladding even though data shows 

little effect. 

M(3) Oxidation characteristics are less important than associated hydrogen pickup. However, non-prototypical fabrication conditions may artificially enhance its 

impact. For example, oxide layers produced under a gaseous mixture of noble 

gas and steam is dense and protective while oxide layer produced under 

irradiation is defective and non-protective. Data show that pre-existing oxide 

has no significant impact on the clad resistance upon quench. Nevertheless the 

clad thinning associated to a thick pre-existing oxide layer affects slightly the stress field in the cladding and the overall clad behavior during the LOCA 

transient. In general, differences among the characterized zirconium based 

materials have shown differences in high-temperature oxidation and quench behavior to be relatively minor. However, specific details on the newer 

materials are not available to the reviewer and so bets are hedged. 

L(O) No votes. 

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N

K(3): 

PK(1): 

UK(1):

Data 

Data 

No rationale provided

Appendix D-18
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) J Definition and Rationale (Imvortance. Annlicabilitv. and I lnoprtiinfv1

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Specimen selection: 

Bumup

The amount of burnup and cladding fluence to which the fuel rod used for the specimen 

was exposed. 

H(2) High burnup alters several important properties of cladding and the nature of 

pellet-cladding interface which influence the resistance to thermal-shock 

failure. If the fuel rod is full of fuel, the fuel morphology will influence the 

quenching. Full rod fuel morphology (fragmentation, rim characteristics, 

bonding, etc.) is important). 

M(2) The clad temperature during a LOCA transient is large enough to anneal all 

irradiation defects. At the time of quench there is no irradiation damages left 

in the cladding. Prior testing by the French and Japanese have shown a 

relatively minor if any effect of pre-existing oxide thickness. However, high bumup specimens should be selected to address the question of unknown or 

previously otherwise uncharacterized effects. Independently of the other 

degradation variables (0, H, etc) burnup may not be important but it's good to 

preserve prototypicality. 

L(O) 

Fuel: Y(1): MOX (agglomerates) in case of full rod 

Clad: N 

Reactor: N 

Burnup: N 

K(2): Data 

PK(2): Data, Judgement: Much is known (Data), but this testing is intended to also 

address the unknown. 

UK(1): No rationale recorded.

Appendix D-19

. .. . .~~~~~~~~~~I . ." - i . . . . L, r . . . . . . . . IL..I 
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Conduct of Test-During 

Axial constraints

Manner in which test specimen is constrained by fittings to simulate potential in-reactor 

axial constrains during the core reflood phase. 

H(6) Phebus 219 rod 18 shows that constraints can affect test outcome. Japanese test 

shows the restraint can affect brittleness results.

M(O) 

L(O)

No votes 

No votes

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(0): 

PK(5): 

UK(1):

No votes 

Data 

Data (incomplete), judgement

Appendix D-20
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Conduct of Test-During 

Azimuthal quenching

Simulation of azimuthal heat flux gradient during testing. 

H(1) This is the type of effect that should be tested using a limiting case so that it 

can be disposed of as a problem or investigated further. 

M(3) In rupture processes, any asymmetry in stress field enhances the rupture. 

Impact is expected to be of the second order. 

L(1) Azimuthally localized nonuniform partial quenching is less prototypic and 

should be avoided in the test. 

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N

K(0): 

PK(4): 

UK(1):

No votes. 

Data, Judgement 

No rationale provided

Appendix D-21
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Conduct of Test-During 

Empty/full

Conduct of test using specimens that have either had the fuel removed (empty) or retained the fuel 

(full). 

H(4) In the case of fuel relocated into the balloon, the clad is less susceptible to 

rewetting due to enhanced heat transfer to the cladding associated with the 

stored energy and the residual power of the fuel. The desire is to perform a 

separate effect test to understand the behavior of the cladding, as opposed to 

confusing interactions introduced by the presence of fuel pellets. The effect of 

the fuel pellets will be assessed during the integral testing. Intact fuel and the 

state of fuel-cladding gap or bonding influence important parameters such as 

clad ID oxidation, hydrogen uptake, and clad mechanical constraints. 

M(2) Data showing the importance of the presence of the fuel within the rod during 

the quenching are not available. Needs testing. 

L(O) No votes.

Fuel: N 

Clad: N 

Reactor: N 

Burnup: Y(1): More important at high burnup.

K(1): ANL and JAERI data, judgment 

PK(4): Data, Judgement: Some pellet effects have been hypothesized, but the point 

here is that this is to be a separate effects test - and so it should be a separate 

effects test (no pellets). 

UK(O): No votes.

Appendix D-22
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Conduct of Test-During 

One-sided vs. two-sided

The quench is applied either on the external side of the cladding or on both sides. The 

latter simulates the case of a rod that burst during the blow down phase. 

H(2) The magnitudes of total oxidation, hydrogen pickup, temperature gradient 

during quench, and thermal stress are influenced significantly by this choice. 

Internal oxidation in the vicinity of the rupture provides significant hydride 

production, which affects the results. 

M(4) In a real quench process, at the balloon height, the cladding suffers both 

mechanisms, i.e., two-sided near the burst opening and one-sided at the 

opposite azimuth. Data show low impact on the result. The thought is that 

the primary parameter affecting quench behavior is the remaining beta phase, 

and not so much where that is (ID in one-sided test or mid-thickness on two

sided tests). However, there should be at least a few tests of the actual 

condition (two-sided) to confirm that a significant difference in performance 

does not occur. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(1): 

PK(4): 

UK(O):

N 

N 

N 

N 

Data 

Data, Judgment 

No votes.

Appendix D-23
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Conduct of Test-During 

Cooldown before quench

The cladding needs to cool down naturally from the oxidation temperature to the 

quenching temperature. 

H(6) The important parameters are the relative values of the beta and alpha 

transition temperatures, which controls some mechanical properties, and the 

quench initiation temperature. The test should be as prototypical as possible to 

avoid undesirable artifacts. French testing has demonstrated a significant 

difference between a fast quench and a fast quench preceded by a slow cooldown 

phase that is more prototypical of the actual condition. Cooldown rate 

strongly influences 0 and H distributions and the microstructure and the 

mechanical properties of the prior beta layer. 

M(O) No votes. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Bumup: 

K(4): 

PK(2): UK(0):

N 

N 

N 

N 

Data 

Data, Judgement 

No votes.

Appendix D-24
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategorv (Test type) I Phenomena (Parameter) I Definition and Rationale (lmnnrtancp Annlieth1ilit ant TTn,',4t2;,i1,,

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Conduct of Test-During 

Clad temperature at 

time of the quench

Clad temperature at the time the quench occurs. It depends on the rod environment. It 

has an impact on the metallurgical morphology of the clad. 

H(5) Whether quench occurs before or after the completion of beta-to-alpha phase 

transformation is a major parameter that influences the magnitude of thermal 

stress and the properties of the prior-beta phase. 

M(1) Hobson's ring tests showed that even for the same level of oxidation, the 

mechanical properties were different for different high temperature levels. 

This difference may or may not affect the quench behavior, and may or may not 

be significant when all temperatures are below 2200 F, however, it suggests 

that a range of pre-quench temperatures be explored to confirm no unexpected differences in behavior exist. 

L(0) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(3): 

PK(4): 

UK(0):

N 

N 

N 

N 

Data, Judgement 

Data, Judgement 

No votes.

Appendix D-25

, v " r ," " .i i .. ... .. . .. . . ... .. .. . .. rL .. .. . . . . .. ILI j . . . .. . . . ,
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Conduct of Test-During 

Cycling of quenching

Repeated loading of the test segment via dryout followed by quenching. 

H(1) Several thermal shocks may result in a higher probability of failure. 

M(2) Before permanent reflood, a cladding may encounter several partial and 

temporary rewetting and dryout periods. This phenomenon may enhance the 

clad failure. low impact of the cycles are expected on the overall behavior of 

the cladding. Should be checked through a separate effect tests series (phase equilibria and transformation kinetics tests). 

L(3) The most severe thermal shock is produced at the first quenching. Not sure 

why this is being considered. It's not the expected case. 

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(2): 

PK(2): 

UK(3):

No votes. 

Data, Judgement 

No information on cyclic quenching, but also not convinced it will happen.

Appendix D-26
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Conduct of Test-During 

Temperature history

Measurement of the time-varying temperature throughout oxidation and quenching. 

H(4) Temperature history is the primary factor that determines oxidation, 

hydrogen uptake, and the mechanical properties of the load-bearing prior beta 

layer. 

M(2) The key parameter is the clad temperature before quench. To measure the local 

clad temperature during the quench might be useful for the analysis but the 

measurement has not to be intrusive (risk of experimental artifacts). 

Knowledge of the actual temperature prior to the slow cool down phase, and 

during the slow cool down phase are important to know that we attained the 

desired pre-quench conditions. However, measurement during the quench is less 

critical (you get what you get and will likely not be able to measure it 

precisely). 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(3): 

PK(3): 

UK(O):

N 

N 

N 

N 

Data, judgement 

Data, Judgement 

No votes.

Appendix D-27
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Quench tests, quench 

rate, Tquench, etc. 

(yote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Conduct of Test-During 

Pre-thinning of cladding

The specimen has experienced the ballooning phase of the LOCA transient before being quenched. 

H(4) The idea here was that the cladding tested should represent the as-thinned 

condition resulting from pre-transient oxidation, since the remaining metal 

thickness pre-test determines the remaining ductile ligament after high 

temperature oxidation (as opposed to using the a section of the fuel rod for the 

test that exhibits unusually low corrosion (like the bottom of the rod for 

PWRs). Since this pre-transient metal loss can be significant (perhaps 10%), this effect can significantly influence the quench test results by correspondingly reducing the remaining ductile region. Pre-thinning influences directly the 

thickness of the load-bearing prior-beat layer available for a given transient 

as well as the magnitude of thermal stress. 

M(2) The clad thinning of the ballooned area of the cladding can be taken into 

account through calculation (geometrical effect only). The residual stresses 

related to the clad straining are annealed during the high temperature 

oxidation phase. 

L(0) No votes.

Fuel: 

Clad: 

Reactor: 

Bumup: 

K(1): 

PK(5): UK(0):

N 

N 

N 

N 

Data 

Data, Judgement 

No votes.

Appendix D-28
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Conduct of Test-During 

Quench mass flow rate

Conduct of Test-PTE 

Equivalent cladding 

reacted (ECR) at location 

of failure

.1

Delivery rate of water used to quench the rod.

H(O) 

M(1) 

L(3)

No vote. 

Important for the thermal-hydraulic conditions. 

Low impact is expected. It is expected that a significant variation in mass flow 

rate can be permitted and will result in effectively equivalent quench 

characteristics.

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(4): 

PK(4): UK(0): Data, Calculation, Judgement

The percentage of Zr atoms that would have reacted with oxygen to form ZrO2 if all 

oxygen absorbed into the cladding were used to form ZrO2.

H(7) 

M(O) 

L(O)

Primary parameter used to understand, extend and evaluate results.

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(4): 

PK(2): 

UK(O):

Data 

Judgement based on data and calculations

Appendix D-29
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategorv (Test tvve) I Phenomena (Parameter) I Definition and Ratinnaip Jmnnw4an'n AnnhI.2h17 ,nA 

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Conduct of Test-PTE 

Metallography

-4 .

Conduct ot Test-PTE 

Fragment/non-fragment

Metallographic examination to determine the morphology of the material, and which 

can be related to phase formation during quench. 

H(6) Metallography is the primary means to determine the degree of oxidation, 

phase structure, and the microstructure of the prior beta layer.

M(O) 

L(O)

No votes. 

No votes.

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N

K(3): 

PK(2): UK(0):

Data, Judgement 

Data, Judgement 

No votes.

Determine if the cladding embrittlement has led to cladding fragmentation 

H(6) Fragmentation of the cladding means risk of fuel dispersal and subsequent 

coolability concern.

M(O) 

L(O)

Fuel: 

Clad: 

Reactor: 

Bumup: 

K(2): 

PK(3): 

UK(0):

No votes. 

No votes.

N 

N 

N 

N 

Data, judgement 

Data, Judgement 

No votes.

Appendix D-30

Phenomena (Parameter) a AL 

I efinition and Rationale (Im ortance lw2hiltf ý"A TT .4 . 4. K r- Z-1
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Quench tests, quench 

rate, Tquench, etc. 

(vote=14) 

Separate effect test to 

determine the thermal 

shock resistance of 

cladding when quenched 

after high temperature 

oxidation.

Conduct of Test-PTE 

Characterization of 

tubing integrity

Perform leak test on the cladding to detect a potential crack. 

H(3) The determination of cladding failure is the primary objective of test. Agree 

with the medium definition but give it more weight. 

M(5) To establish the limit of failure upon quench the cladding integrity has to be 

defined and checked. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(2): 

PK3): 

UK(O):

N 

N 

N 

N 

Data, Judgement 

Data, Judgement 

No votes.

Appendix D-31
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) - ALiLjJUALULLLt I)11L1UII1i3', aria uncertaintyI . fini.ini .. . nA. . -I---Ii ITL.U.La. A 1 tl-ca - and -l- --- uncertainty)

Phase equilibria and 

transformation kinetics

chemistry effects 

(vote=11) 

Measurement of phase 

equilibria and phase 

transformation kinetics 

that can provide 

fundamental data 

relevant to the cladding 

behavior during LOCA 

events.

Specimen selection: 

Hydrogen content and 

distribution

Amount of hydrogen in the sample and where it is located. 

H(4) Hydrogen affects the alpha, alpha plus beta, and beta boundaries. Data show that hydrogen content has an impact on phase transformation temperature and 

kinetics. Hydrogen is a strong beta stabilizer which influences phase equilibria, transformation kinetics, and the structure of prior beta layer. M(1) This type of information is useful in explaining, and possibly extrapolating the observations, but does not represent direct confirmation that criteria are met or 

not met. As such this whole category is rated at a lower level than some of the 

other more performance related characterizations. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Bumup: 

K(3): 

PK(2): 

UK(O):

N 

Y(1): More important for M5 and Zirlo. 

N 

Y(1): More important for high-burnup fuel. 

Data 

Data, Judgement 

No votes.

Appendix D-32
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Phase equilibria and 

transformation kinetics

chemistry effects 

(vote=11) 

Measurement of phase 

equilibria and phase 

transformation kinetics 

that can provide 

fundamental data 

relevant to the cladding 

behavior during LOCA 

events.

Specimen selection: 

Alloy type

Composition or designation of the metal utilized in fuel-rod fabrication 

H(4) Alloying elements such as Sn and Nb affect the alpha, alpha plus beta, and 

beta boundaries. Nb-bearing M5 and Zirlo behave significantly different from 

zircaloy. Affects the phase equilibria and the transformation kinetics. 

M(1) This type of information is useful in explaining, and possibly extrapolating the 

observations, but does not represent direct confirmation that criteria are met or 

not met. As such this whole category is rated at a lower level than some of the 

other more performance related characterizations. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(2): 

PK(3): 

UK(O):

N 

N 

Y(1): PWR 

N 

Data 

Data, Judgement 

No votes.

Appendix D-33
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Phase equilibria and 

transformation kinetics

chemistry effects 

(vote=11) 

Measurement of phase 

equilibria and phase 

transformation kinetics 

that can provide 

fundamental data 

relevant to the cladding 

behavior during LOCA 

events.

Specimen selection: 

Oxygen content

Amount of oxygen either dissolved or existing as an oxide phase in the cladding. 

H(3) Oxygen is a strong alpha stabilizer. Total amount of oxygen will affect phase 

equilibria. 

M(3) Data show low impact of initial oxide per se. Only Hydrogen related to the 

pre-transient oxide plays a role. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Bumup: 

K(4): 

PK(1): 

UK(O):

N 

N 

N 

N 

Data, Judgement 

Data, Judgement 

No votes.

Appendix D-34
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategorv (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Phase equilibria and 

transformation kinetics

chemistry effects 

(vote=11) 

Measurement of phase 

equilibria and phase 

transformation kinetics 

that can provide 

fundamental data 

relevant to the cladding 

behavior during LOCA 

events.

Specimen selection: 

Fluence

Fast neutron fluence experienced by the specimen

H(O) No votes. 

M(2) At 62 GWd/t, the major factor is hydrogen pickup; however, the effect of 

fluence is less certain at higher burnups. 

L(3) Irradiation damages are quickly annealed during the transient. Second-phase 

precipitates are amorphized by irradiation, irradiation damages are annealed 

out rapidly at >550'C. Do not see how this will influence the result of the test.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(2): 

PK(3): 

UK(O):

N 

Y(1): Not well known for M5 and Zirlo. 

N 

N 

Data, judgement 

Data, Judgement 

No votes.

Appendix D-35
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcator (Test type) I Phenomena ............. -.... I ...... ... .'* =AIJ-•VyL•an% tppllcaW1t1[y, and_ uncertainty)

Phase equilibria and 

transformation kinetics

chemistry effects 

(vote=11) 

Measurement of phase 

equilibria and phase 

transformation kinetics 

that can provide 

fundamental data 

relevant to the cladding 

behavior during LOCA 

events.

Determination of 

hydrogen and oxygen 

solubilities in ot and 0 

phases as a function of 

hydrogen, oxygen, and 

temperature for relevant 

alloys

Establishes the relationship between oxygen and hydrogen potential and phase compositions. 

H(4) These parameters are necessary to allow relevant modeling and analysis. M(1) Seems to be redundant with establishing the phase diagram. Data directly influence assessment of the validity of the current LOCA embrittlement 

criteria for high-burnup operation and for new types of alloy. L(O) No votes. 

Fuel: N 

Clad: Y(1): More important for M5 and Zirlo 

Reactor: N 

Burnup: Y(1): More important at high burnup 

K(1): Data, judgement 

PK(4): Calculations, Data, Judgement 

UK(O): No votes.

Appendix D-36
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Phase equilibria and 

transformation kinetics

chemistry effects 

(vote=11) 

Measurement of phase 

equilibria and phase 

transformation kinetics 

that can provide 

fundamental data 

relevant to the cladding 

behavior during LOCA 

events.

Determination of time 

constants for limiting 

mechanisms for phase 

transformation during 

heating as a function of 

hydrogen, heating rate 

and cooling rate

Phase transformations often require time to transform from one phase to the other, and 

kinetic transport mechanisms often control the reaction rates. 

H(3) Should be known for relevant calculation and analysis. Knowing these 

transport mechanisms may allow the determination of reaction kinetics. 

M(1) Seems to be redundant with establishing the phase diagram. 

L(1) Alpha-to-beta transformation kinetics during the heatup phase are very fast. 

Beta-to-alpha transformation during the cooldown phase is not limited by 

cooling rate; that is, depending on cooling rate, either diffusionless martensitic 

transformation or transformation via nucleation and growth are possible. 

However, transformation microstructure and the degree of 0 and H 

redistribution during the transformation are strongly influenced by the cooling 

rate.

Fuel: 

Clad: 

Reactor: 

Bumup: 

K(2): 

PK(3): UK(0):

N 

N 

N 

N 

Data, Judgement 

Data, Judgement 

No votes.

Appendix D-37
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Phase equilibria and 

transformation kinetics

chemistry effects 

(vote=11) 

Measurement of phase 

equilibria and phase 

transformation kinetics 

that can provide 

fundamental data 

relevant to the cladding 

behavior during LOCA 

events.

Phase equilibria and 

transformation kinetics

chemistry effects 

(vote=11) 

Measurement of phase 

equilibria and phase 

transformation kinetics 

that can provide 

fundamental data 

relevant to the cladding 

behavior during LOCA 

events.

Determination of 

diffusion coefficient of 

oxygen in individual 

phases

4 I

Determination of the 

retained 3 and 

transformed O-phase 

morphology and oxygen 

plus hydrogen 

redistribution during 

P - a transformations 

(cooling), including 

Niobium-rich alloys

Need definition

H(1) 

M(1) 

L(1)

Drives the level of clad embrittlement. 

Seems to be redundant with establishing the phase diagram. 

A significant database is available for Zr and Zircaloy.

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N

K(2): 

PK(1): UK(0):

Data 

Data, Judgement 

No votes.

Defined by title.

H(2)

M(O) 

L(O) 

Fuel: 

Clad: 

Reactor: 

Bumup: 

K(0): 

PK(2): UK(0):

Rated higher simply because this is the bottom line; other items are needed to 

develop an analytical representation. These factors play direct and very 

important roles, which determine the clad resistance to thermal shock and the 

post-quench mechanical properties. 

No votes. 

No votes. 

N 

N 

N 

N 

No votes. 

Data, Judgement 

No votes.

Appendix D-38
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicabilitvo and Uncertainty)

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Specimen selection: 

Pre-existing oxide

The total amount of oxide formed on the cladding during in-reactor operation and 

whether the oxidation is uniform or nodular, and whether there is extensive cracking and 

spalling. 

H(1) Directly influences burst and creep strengths. 

M(2) Data show that impact is limited to related Hydrogen content and clad 

thinning. 

L(1) The pre-existing oxide thickness determines the remaining metal thickness, 

which directly determines the creep and burst behavior. However, over the 

ranges of practical interest, the variations are not expected to be all that 

significant.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(4): 

PK(0): UK(0):

N 

N 

N 

N 

Data, Judgement 

No votes. 

No votes.

Appendix D-39
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategorv (Test type) I Phenomena (Parameter) I Definition and Rationale (Imoortance. AmDlicabilitv. and U/ncertainfv•

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Specimen selection: 

Alloy and initial 

thermomechanical 

treatment

Composition or designation of the metal utilized in fuel-rod fabrication and the 

thermomechanical treatment. 

H(5) Alloy type may affect the burst behavior. Beta favoring or alpha favoring 

may impact the mechanical behavior. Data are needed to do relevant 

analysis. Creep behavior is known to vary significantly with 

thermomechanical treatment and alloy type.

M(1) 

L(O)

Do not expect a big influence of cladding type at high temperature. 

No votes.

Fuel: N 

Clad: Y(1): More important for M5 and Zirlo. 

Reactor: N 

Burnup: N 

K(2): Data 

PK(2): Data, Judgement 

UK(1): No rationale provided.

Appendix D-40

Definition and Rationale (Imi3ortance A licabilit and ncertaint I
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Specimen selection: 

Hydrogen content

Total amount of pre-existing hydrogen in the material.

H(2) 

M(4) 

L(0) 

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(1): 

PK(4): 

UK(0):

Affects burst behavior (alpha to beta transformation). 

Data show some hydrogen impact on the mechanical properties at high 

temperatures. The French have reported an effect of cladding creep strength 

with hydrogen content. Hydrogen content influences phase stability and burst 

behavior. Creep failure at <550C during a LOCA is of less concern. Do not 

expect a large effect. 

No votes. 

N 

Y(1): Important for M5 and Zirlo 

N 

N 

Data 

Data, Judgement 

No votes.

Appendix D-41
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Specimen selection: 

Fluence (radiation 

damage)

Level of radiation damage in cladding after in-reactor irradiation. 

H(1) Irradiation hardening is known to affect the cladding creep behavior. 

Alternate effects, such as precipitate dissolution, are not as well 

characterized. 

M(1) At 62 GWd/t, the major factor is hydrogen pickup; the importance of fluence at 

higher burnups is less certain. 

L(3) Irradiation damages are annealed at the test temperature. Irradiation 

damages are virtually annealed out rapidly at >600'C.

Fuel: 

Clad: 

Reactor: 

Bumup: 

K(3): 

PK(2): UK(0):

N 

N 

N 

N 

Data 

Data, Judgement 

No votes.

Appendix D-42
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) i Definition and Rationale (Imvortance. Annlicabilitv. and T~nc'prbiinfv)

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Conduct of Test-During 

Strain profile as a 

f(r, 0, z, t)

Measurement of local strain variation with strain gages and similar equipment, such 

that the full strain distribution is known. 

H(4) These data are used to validate the thermomechanical models. In the interest 

of creep deformation, an accurate history of the deformation behavior during the test is essential to characterize differences in behavior. Relative to burst 

behavior, characterization of the resulting burst strain may be sufficient 

without detailing temporal characterization. Difficult to obtain this data, but 

useful and with high importance if we can get it. 

M(1) Data are needed to develop models. 

L(0) No votes.

Fuel: 

Clad: 

Reactor: 

Bumup: 

K(3): 

PK(1): 

UK(0):

N 

N 

N 

N 

Data, Judgement 

Data, Judgement 

No votes.

Appendix D-43

...... ~~~~D fn to Rai n l .... ......... rtnc A- -n -,"1 .. .. •.. an . .....e tan 
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Conduct of Test-During 

Pressure as f(t)

Internal pressure in the rod as a function of time. 

H(5) To allow proper interpretation of the test, the internal pressure versus time 

should be measured. The creep deformation history cannot be interpreted 

without a corresponding temporal characterization of the driving force 

(pressure).

M(O) 

L(1)

No votes. 

A pressure history is needed to develop models.

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(3): 

PK(2): 

UK(0):

Data, Calculation 

Data, Calculation, Judgement 

No votes.

Appendix D-44
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Conduct of Test-During 

Temperature as f(t)

Measurement of the time-varying temperature. 

H(5) Impacts mechanical resistances of the specimen. Phase transformation and 

distribution and subsequent cladding mechanical properties depend on the clad 

temperature. To use the test results as a validation database for the 

calculation codes, the time-varying clad temperature should be measured. 

Creep is a very sensitive function of the cladding temperature and therefore 

the temperature must be known. Burst and creep behaviors are sensitive to clad 

maximum temperature. Obviously need to know temperature. 

M(O) No votes. 

L(O) No votes. 

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N

K(4): 

PK(1): 

UK(0):

Data, Calculation 

Data, Judgement 

No votes.

Appendix D-45
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Conduct of Test-During 

Temperature profile as 

f(6) and f(z)

Measurement of azimuthal and axial variations of temperature. 

H(4) Phase transformation and distribution, and subsequent cladding mechanical 

properties, depend on the clad temperature. To use the test results as a 

validation database for the calculation codes, the time-varying clad 

temperature should be measured. 

M(1) Burst strain is known to be sensitive to circumferential temperature differences. 

Measurement of the circumferential temperature distribution would be needed 

to best interpret the testing results. 

L(0) No votes. 

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N

K(3): 

PK(2): 

UK(0):

Data, Calculation 

Data, Judgement 

No votes.

Appendix D-46
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Conduct of Test-During 

Open (actively 

pressurized) or closed

The test is conducted either with closed ends tube (constant number of moles) or with open 

ends (controlled pressure). 

H(5) It is crucial to represent the actual pressure evolution of a full-length rod. 

Knowledge and control of the internal pressure is essential to obtaining useful 

characterizations of creep and burst behavior which probably leads to open 

(actively pressurized) tubes. The pressure in a closed tube would vary with 

heatup and with cladding deformation. Affects test. 

M(0) No votes. 

L(1) Closed burst is a better simulation. 

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(4): 

PK(1): 

UK(O):

Data, Judgement 

Data, Judgement 

No votes.

Appendix D-47
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Conduct of Test-During 

Biaxiality ratio

The state of stress experienced by the cladding during testing. 

H(5) There is preliminary information on the impact of axial stress on the cladding 

rupture; additional experimental results are needed. Burst at <830'C 

(deformation controlled by prism slip in the highly anisotropic alpha phase) is sensitive to biaxial ratio. Influences failure limit. 

M(2) The most directly useful testing would simulate the actual cladding stress 

state. 

L(0) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(2): 

PK(2): 

UK(0):

N 

N 

N 

N 

Data 

Data, Judgement 

No votes.

Appendix D-48
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Conduct of Test-During 

Load and displacements, 

i.e., a and e behavior

Determination of the stress-strain response of the material using uniaxial testing. 

H(5) Objective of the test. The ability of the cladding material to withstand 

mechanical loadings is directly related to the basic mechanical properties; 

primarily ductility but also strength. Development of the material stress

strain curve provides information necessary to understand possible performance 

differences in the cladding capability with increasing exposure. Therefore, 

accurate measurement of these quantities is essential. Important for code 

development. 

M(0) No data. 

L(O) No data. 

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N

K(3): 

PK(2): UK(0):

Data, Calculation, Judgement 

Data, Judgement 

No data.

Appendix D-49
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) L ---- I LL r--- L IDefinition and Rationale (Im ortance. Annlicabilitv and TUncrf-inhA

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Conduct of Test-During 

Strain rate

The rate at which the strain is applied.

H(2) Relevant parameter. Primary test parameter to be fixed and measured. 

M(1) Available testing has demonstrated that strain-rate effects exist. However, 

for the purpose of tensile property testing, the strain rate should be relatively 

quick (to best reflect true elastic-plastic behavior and avoid creep-induced 

inaccuracies), and within the practical range of achievable test strain rates 

relative to the target, spurious strain rate effects are not expected. 

L(0) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(2): 

PK(1): 

UK(O):

N 

N 

N 

N 

Data 

Data, Judgement 

No votes.

Appendix D-50
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Mechanical Properties at 

high temperature, e.g., > 

300 C (vote=10) 

Creep and burst tests 

Designed to investigate 

creep and burst behavior 

of cladding at high 

temperature

Conduct of Test-During 

Circumferential 

(hoop)/axial (not ring)

I. 1*

Conduct of Test-PTE 

Post-test strain

Whether the uniaxial test should be done by a ring test or by an axial test.

H(5) 

M(1) 

L(0)

Relevant for non-isotropic materials. Primary test parameters to be measured. 

Appropriate test should be used 

No votes.

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(2): 

PK(1): 

UK(0):

Data 

Data, Judgement 

No votes.

Perform strain measurements at the failure locations. 

H(8) Relevant parameter. A key output of the burst characterization is the rupture 

strain, as input to flow blockage assessment. The most important test objective; 

needed to develop models.

M(0) 

L(1)

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(3): 

PK(2): 

UK(0):

No votes 

Relevance of the data is low especially if the internal pressure evolution is not 

prototypical. 

N 

N 

N 

N 

Data, Judgement 

Data, Judgement 

No votes.

Appendix D-51
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Mechanical Properties at 

low temperature, e.g., _ 

300 C (vote=10) 

Post oxidation and 

quench ductility test

Post oxidation and 

quench ductility tests, 

e.g., 

1. Axial tensile 

2. Ring tensile 

3. Ring compression 

4. Impact 

5. Bending

To simulate a post LOCA seismic event two types of mechanical tests are suggested: four

point bending test (Note 1) and impact test to simulate the impact between the grid 

dimples and the rods (Note 2). 

H(4) The best testing to perform is four-point bend testing, supplemented by ring 

testing. Four-point bend testing addresses the post-LOCA performance limit of 

greatest interest, while ring testing provides supplementary insight into the 

fundamental mechanical properties as and aid in interpretation of the four

point bend testing results. The order of higher importance is--impact test 

followed by ring-compression test. 

M(1) Axial tensile test is a convenient test that can identify the most vulnerable 

spot along the cladding length, however, post-quench axial tensile loading is 

not expected to occur or to be insignificant in magnitude (Note 4). 

L(1) Ring-tensile stress in a burst cladding is either negligible or insignificant in 

post-quench phase. 

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(1): 

PK(3): UK(0):

Data, Judgement 

Data, Judgement 

No votes.

Appendix D-52
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty) 

Notes: 

1. The specimen should be a least 50-cm long (one span) and should contain its gap closed fuel pellet stack. The loading to be applied to the 

specimen should be limited to a given deflection (during a seismic event the maximum deflection for a 4-m fuel assembly is currently less 

than 25-mm). 

2. The specimen should be 5 to 10 cm long and should include its fuel pellet stack. The loading is a pulse whose magnitude and width are 

respectively 165 N and 2-3 ms 

3. In order to get immediately comparable data to those from which the criteria were deduced in 1973, perform the new test with a ring 

compression process. Ring tensile tests may be very severe (recent Halden results on dried-out rods). The relevance of the mechanical test is 

the key. The loading applied to the specimen should be prototypical. It is imperative that the right type of testing be performed to 

determine the most relevant characterization. These types of test are most applicable to post-quench modes of loading (due to various 

hydraulic, handling, and other mechanical forces) and deformation of the ballooned, burst, and oxidized cladding. Bending test is 

addressed in separate below (i.e., ability to withstand post-LOCA seismic events and aftershocks). 

4. Test results should be interpreted very carefully, because according to previous studies in ANL and JAERI, the most vulnerable spot produced 

in burst cladding is strongly influenced by the selected heating method, the degree of temperature nonuniformity near the burst region, and 

oxidation and hydrogen uptake from the clad inner surface.

Appendix D-53
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) i Definition and Rationale (Imnortance. Anrnlh'ibili~v_ anAl lTn,'r.~A,4~},

Seismic tests 

Test type: 4--point 

bending 

Separate effect test that 

addresses the ability of 

the fuel rod to withstand 

a post-LOCA seismic 

event without shattering

Specimen selection: 

Alloy type

Composition or designation of the metal utilized in fuel-rod fabrication 

H(4) May affect oxygen distribution and hydrogen pickup. Hydrogen pick-up 

fraction of the alloy during operation will be a primary parameter since very little hydrogen is absorbed during the LOCA transient. Very little, if any, of 

this type of testing has been performed to date and therefore a judgement of the 

relative importance of various parameters is difficult and necessarily 

speculative at this time. It is proposed for this category that all identified 

parameters be assigned High Importance but Unknown. The effects of 0, H, 

hydrides, and the microstructure of prior beta layer are strongly influenced by alloy type. 

M(O) No votes.

L(O) 

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(1): 

PK(1): 

UK(2):

No votes. 

N 

Y(1): More important for M5 and Zirlo 

N 

N 

Data 

Data, Judgement 

Lack of data

Appendix D-54

. .. .. . . .,t. . .. ~~~~~~~~~L 'ie.. . .. Z . . .. . ... L 
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategorv (Test type) I Phenomena (Parameter) I Definition and Ratinna1(!ninnw4un'p AtnUi'h1Ifr 2 tiA TI. 4.4.4.A . . . . . L I Z-1-ý.. .a, A'tL'

Seismic tests 

Test type: 4--point 

bending 

Separate effect test that 

addresses the ability of 

the fuel rod to withstand 

a post-LOCA seismic 

event without shattering

Specimen selection: 

Thickness and 

morphology of pre

existing and transient 

oxides

The total amount of oxide formed on the cladding and whether the oxidation is uniform 

or nodular, and whether there is extensive cracking and spalling. 

H(3) Oxide related clad thinning will slightly affect the mechanical behavior of 

the cladding. Very little, if any, of this type of testing has been performed to 

date and therefore a judgement of the relative importance of various 

parameters is difficult and necessarily speculative at this time. It is proposed for this category that all identified parameters be assigned High Importance but Unknown. These parameters strongly influence the thickness, 0 and H 

contents, and hydriding behavior of the load-bearing prior beta layer. 

M(1) Oxidation characteristics less important than hydrogen characteristics. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Bumup: 

K(3): 

PK(0): 

UK(1):

N 

N 

N 

N 

Data, Calculation 

No votes. 

Lack of data

Appendix D-55
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Seismic tests 

Test type: 4--point 

bending 

Separate effect test that 

addresses the ability of 

the fuel rod to withstand 

a post-LOCA seismic 

event without shattering

Specimen selection: 

Bumup

The amount of burnup and cladding fluence to which the fuel rod used for the specimen 

was exposed. 

H(2) Very little, if any, of this type of testing has been performed to date and 

therefore a judgement of the relative importance of various parameters is 

difficult and necessarily speculative at this time. It is proposed for this 

category that all identified parameters be assigned High Importance but 

Unknown. Major factor that influences the material structure and properties. 

M(1) At 62 GWd/t, the major factor is hydrogen pickup; however, the importance of 

fluence at higher burnups is unclear. 

L(1) The irradiation damages have been annealed during the LOCA transient and 

will not affect the post-LOCA mechanical behavior of the cladding.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(2): 

PK(1): 

UK(1):

N 

N 

N 

N 

Data, Judgement 

Data, Judgement 

Lack of data

Appendix D-56
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Seismic tests 

Test type: 4--point 

bending 

Separate effect test that 

addresses the ability of 

the fuel rod to withstand 

a post-LOCA seismic 

event without shattering

Specimen selection: 

Pre-existing and 

transient hydrogen 

content and distribution

Amount and distribution of pre-existing hydrogen associated with fuel rod clad segment. 

This hydrogen may be in solution in the metal or may exist combined with the metal as a 

discrete hydride phase. 

H(4) Affects oxygen solubility in the beta phase and post-quench ductility in the 

absence of burst. Since the cladding during the LOCA absorbs little Hydrogen, 

the initial Hydrogen content will play a key role during the post-LOCA 

mechanical tests. The initial Hydrogen distribution has no importance. Very 

little, if any, of this type of testing has been performed to date and therefore a 

judgement of the relative importance of various parameters is difficult and 

necessarily speculative at this time. It is proposed for this category that all 

identified parameters be assigned High Importance but Unknown. Post-quench 

hydride distribution is the major factor that influences the mechanical 

properties, especially at <200°C. 

M(O) No votes. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(2): 

PK(1): UK(1):

N 

N 

N 

N 

Data, Judgement 

Data, Judgement 

Lack of data

Appendix D-57
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Seismic tests 

Test type: 4--point 

bending 

Separate effect test that 

addresses the ability of 

the fuel rod to withstand 

a post-LOCA seismic 

event without shattering

Specimen selection: 

With or without 

ballooning

Determining whether the four-point bending test should be performed on a fuel rod 

section that includes a ballooned region. 

H(4) The clad geometry change (clad thinning and diameter increase) will affect 

the mechanical behavior of the rod. The unanswered question is whether the 

seismic loads will cause unacceptable failure elsewhere. Ballooning and burst 

influence strongly clad ID-side oxidation, hydrogen uptake, and hydriding, 

producing weak spots near the burst opening. 

M(O) No votes.

L(O) 

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(1): 

PK(2): UK(0):

No votes. 

N 

N 

N 

N 

Data, Judgement 

Data, Judgement 

No votes.

Appendix D-58
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategorv (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Seismic tests 

Test type: 4--point 

bending 

Separate effect test that 

addresses the ability of 

the fuel rod to withstand 

a post-LOCA seismic 

event without shattering

Conduct of Test-During 

Temperature

The temperature at which the test is conducted. 

H(3) Ductility may vary a lot with temperature. Variations in performance with 

test temperature have been demonstrated by Hobson's ring tests. The more 

critical issue is to define the relevant temperature range, and then maintenance 

of that temperature during the test should be achievable and enforced. With 

the extent of embrittlement anticipated in the test sample, the maintenance of 

the intended temperature range is critical for obtaining meaningful test results. 

Bending temperature (<200°C) is a major factor that influences test results. 

M(O) No votes. 

L(1) The post LOCA tests are isothermal.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(O): 

PK(4): 

UK(O):

N 

N 

N 

N 

No votes. 

Data, Experience, Judgement 

No votes.

Appendix D-59
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

SSubcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Seismic tests 

Test type: 4--point 

bending 

Separate effect test that 

addresses the ability of 

the fuel rod to withstand 

a post-LOCA seismic 

event without shattering

Conduct of Test-During 

Strain rate 

(displacement ratio)

Measurement of the strain versus time 

H(3) Relevant parameter. The objective of the test is to measure the response of the 

rod to an imposed deflection. Strain rate effects can be expected; prototypical strain rates are needed to ensure meaningful test results 

M(1) No rationale provided (ANL) 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(O): 

PK(3): 

UK(O):

N 

N 

N 

N 

Data 

Data, Judgement 

No votes.

Appendix D-60
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Seismic tests 

Test type: 4--point 

bending 

Separate effect test that 

addresses the ability of 

the fuel rod to withstand 

a post-LOCA seismic 

event without shattering

Conduct of Test-During 

ASTM specification

The shape of the contact points and the way to apply the loading should follow the 

ASTM specification to avoid undesirable experimental artifact (local stress 

concentration) 

H(2) Relevant specification to avoid non-prototypical loading of the rod. 

M(1) It may be difficult to apply it for ballooned, burst, and nonuniformly oxidized 

and hydrided cladding. 

L(O) No votes. 

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N

K(1): 

PK(3): UK(O):

Data 

Data, Calculation, Judgement 

No votes.

Appendix D-61
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) 1 Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Seismic tests 

Test type: 4--point 

bending 

Separate effect test that 

addresses the ability of 

the fuel rod to withstand 

a post-LOCA seismic 

event without shattering

Conduct of Test-During 

Appropriate bending 

moment

During a seismic event the fuel assembly is submitted to a prototypical bending moment. 

H(4) The tests have to be prototypical. The loading to be applied to the specimen should be limited to a given deflection (during a seismic event the maximum 

deflection for a 4-m fuel assembly is currently less than 25-mm). The 

application of an appropriate bending moment is essential to obtaining meaningful results. However, it is also anticipated that once the prototypical bending moment is demonstrated to be successful, the bending moment will be 

increased to determine margin to the critical bending moment for failure and 

whether the consequences of failure are truly unacceptable from the 

consideration for maintenance of coolable geometry.

M(O) 

L(O)

No votes. 

No votes.

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N

K(1): 

PK(3): 

UK(O):

Data, Judgement 

Data, Calculation, Judgement 

No votes.

Appendix D-62
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Seismic tests 

Test type: 4--point 

bending 

Separate effect test that 

addresses the ability of 

the fuel rod to withstand 

a post-LOCA seismic 

event without shattering

Conduct of Test-During 

Cycling

Apply a cyclic loading on the rod to simulate the seismic event. 

H(3) Cycling induces metal fatigue. By the very nature of the post-LOCA seismic 

event, cycling can be anticipated, although at a relatively low frequency. It is 

well-known from fatigue studies that the allowable strain amplitude 

decreases with increasing number of cycles; the quantification of this reduction 

would be needed if it appears that the expected loading approaches the level 

for unacceptable consequences with a single cycle. However, if considerable 

margin to fracture/unacceptable consequences exists then multiple cycle testing 

may not be necessary.

M(O) 

L(1)

No votes. 

The magnitude of the seismic load is low enough to avoid plastic deformation 

of the clad and the number of cycles is not large enough to create fatigue 

damage. No cumulative damage is expected.

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N

K(1): 

PK(3): UK(0):

Data, Judgement 

Data, Judgement 

No votes.

Appendix D-63
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Seismic tests 

Test type: 4--point 

bending 

Separate effect test that 

addresses the ability of 

the fuel rod to withstand 

a post-LOCA seismic 

event without shattering

Conduct of Test-PTE 

Characterize integrity

Define the level of fragmentation of the cladding after the post-LOCA mechanical test. 

H(4) The objective of the test is to define the LOCA limits (%ECR and max clad 

temperature) that provoke clad fragmentation and potential subsequent core 

coolability concern by allowing extended fuel dispersal. These limits are 

beyond those leading to a simple loss of clad integrity. Characterization of 

post-test geometry is critical to the determination of whether an acceptable 

geometry has been maintained (when demonstrating the prototypical bending load case), or whether a truly unacceptable condition is developed with 

fracture (when the test is extended to intentionally develop fracture). 

M(O) 

L(O)

Fuel: 

Clad: 

Reactor: 

Bumup: 

K(1): 

PK(1): 

UK(2):

N 

N 

N 

N 

Data, Judgement 

Data, Judgement 

Lack of data

Appendix D-64
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Seismic tests 

Test type: 4--point 

bending 

Separate effect test that 

addresses the ability of 

the fuel rod to withstand 

a post-LOCA seismic 

event without shattering

Conduct of Test-PTE 

Characterize local 

hydrogen

Amount and distribution of hydrogen associated with fuel rod clad segment. This 

hydrogen may be in solution in the metal or may exist combined with the metal as a 

discrete hydride phase. 

H(4) Primary parameter that could impair the clad resistance. Only the initial 

amount is important. The Hydrogen distribution has no impact since it becomes 

uniform during the LOCA transient. A characterization of the hydrogen and 

oxygen distribution would aid in the interpretation of the test results. The 

fracture characteristics and susceptibility are expected to be directly related to 

these embrittling agents. Hydride morphology, orientation, number density, 

and distribution influence test result significantly.

M(O) 

L(O)

No votes. 

No votes.

Fuel: N 

Clad: N 

Reactor: N 

Burnup: Y(1): More important at high burnup 

K(1): Data 

PK(2): Data, Judgement 

UK(O): No votes.

Appendix D-65
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter)

.efin it on an d R ation .... ale.g. (!mtn.... A. 1,- l. . .U. Aa

Simulation of fuel 

relocation Specimen selection: 

Bumup

The amount of burnup to which the fuel rod used for the specimen was exposed. 

H(4) Fuel morphology (fragmentation, rim characteristics, bonding, etc.) is 

important. The nature of the bonding between the pellet and the cladding changes with the burnup increase. It will affect the potential for fuel 

relocation. The segment burnup level can determine the extent of pellet cladding bonding and corresponding susceptibility to fuel relocation during ballooning and rupture. Fuel and rim-zone microstructure and the state of 

bonding with cladding are strongly influenced by fuel burnup. 

M(O) No votes. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(3): 

PK(3): UK(0):

Y(1): MOX agglomerates 

N 

N 

N 

Data 

Data, Judgement 

No votes.

Appendix D-66
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) i Definition and Rationale (Importance, Applicability, and Uncertainty) 

Simulation of fuel Specimen selection: Composition of the fuel, i.e., a specified MOX composition. 

relocation Fuel type (MOX) 

H(2) May affect the amount of fine grain material after relocation. Fuel structure 

and mechanical properties are influenced by fuel type. 

M(1) The consequence of fuel fragments relocation (higher local decay heat and 

higher cladding temperature) could be more effective with MOX fuel than 

with U02 fuel. Nevertheless the viscoplastic properties of the MOX should 

impair the fuel fragments relocation at high burnup. 

L(1) No significant differences in pellet-cladding bonding behavior or pellet 

cracking behavior are anticipated or have been observed with MOX fuel, and 

therefore no significant differences in relocation behavior are anticipated. 

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N 

K(1): Data 

PK(2): Data, Judgement 

UK(1): Judgement

Appendix D-67
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Simulation of fuel 

relocation Specimen selection: 

Alloy type

Composition or designation of the metal utilized in fuel-rod fabrication 

H(2) May affect burst (beta favoring or alpha favoring additions). Ductile burst and 

brittle failure by thermal shock and post-quench forces are influenced strongly by cladding alloy type. 

M(1) In general, compositional differences have not been observed to significantly affect cladding burst behavior. However, if significant differences in burst 

behavior occurred, the relocation characteristics could be similarly significantly altered. 

L(1) Data show no significant impact of alloy type on the balloon size that could 

influence the fuel fragments relocation.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(2): 

PK(1): 

UK(1):

N 

N 

N 

N 

Data 

Data, Judgement 

Judgement

Appendix D-68
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty) 

Simulation of fuel Specimen selection: Chemical and mechanical bonding between the fuel pellet and the cladding prior to the 

relocation Chemical and test. 

mechanical bonding 

H(4) Fuel morphology (bonding) is important. It will affect the potential for fuel 

fragmentation relocation. It is speculated that bonding could significantly 

affect the relocation characteristics by impeding pellet fragment movement. 

However, this effect has not been demonstrated. Major factor that influences 

fuel slumping and potential release of fuel particles upon burst and subsequent 

fragmentation. 

M(O) No votes. 

L(O) No votes. 

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N 

K(O): No votes. 

PK(3): Data, Judgement 

UK(1): Lack of data

Appendix D-69
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty) 

Simulation of fuel Specimen selection: Crack pattern and crack density of the fuel pellets prior to the test. 

relocation Cracking 

H(2) Controls the rubble bed characteristics after relocation. Degree of fuel cracking 

directly influences the potential for fuel relocation and release. 

M(O) No votes. 

L(2) Beyond a given burnup the number of cracks is stable. In general the macroscopic 

fuel pellet cracking pattern develops early in life and does not change significantly with elevated exposures. Therefore, this contribution to fuel 

relocation susceptibility is not expected to be a dominant parameter during this 

test series. 

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N 

K(1): Data 

PK(3): Data, Judgement 

UK(O): No votes.

Appendix D-70
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategorv (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Simulation of fuel 

relocation

Conduct of Test-During 

With or without 

blowdown

Determination of whether blowdown processes must be simulated in the test. 

H(O) No votes. 

M(1) During the blowdown phase of the LOCA transient, fuel stored energy is 

redistributed in the pellet and the clad. This redistribution produces a 

decrease of the pellet centerline temperature and increases the pellet rim and 

clad temperatures. Due to these temperature transients, the central part of the 

pellet will suffer a contraction while the rim and the clad will experience an 

expansion. These adverse effects could induce fuel mechanical stresses and 

fragmentation. The expansion and contraction inside the fuel pellet may affect 

bonding and fuel debris sizes. 

L(2) Vibration loads occurring during the blowdown phase may cause additional 

pellet fragment movement. In general, pellet fragments are relatively 

constrained within the fuel rod by the column geometry, as evidenced by 

characterization of fuel column geometry in hot cells. Therefore, this effect is 

not considered to significantly contribute to relocation susceptibility later 

during the cladding heatup and rupture phases. Fuel thermal contraction and 

cladding heatup during the blowdown phase increases the pellet-cladding gap 

and possibly facilitates pellet fragment relocation. Cladding heatup rate and 

temperature, either with or without a blowdown, are the primary factors that 

influence burst shape and dimensional changes.

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N

K(0): 

PK(1): UK(2):

No votes. 

Data, Experience, Judgement 

Judgement

Appendix D-71
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Simulation of fuel 

relocation

Conduct of Test-During 

Blowdown temperature 

transients for fuel and 

cladding

Simulation of the temperature response of the fuel and cladding during the blowdown 

phase of a large-break LOCA.

H(2) 

M(O) 

L(1)

Important parameters that influence cladding burst and dimensional changes. 

No votes. 

Pellet fragment movement. In general, pellet fragments are relatively constrained within the fuel rod by the column geometry, as evidenced by characterization of fuel column geometry in hot cells. Therefore, this effect is 

not considered to significantly contribute to relocation susceptibility later 

during the cladding heatup and rupture phases. Fuel thermal contraction and 

cladding heatup during the blowdown phase increases the pellet-cladding gap and possibly facilitates pellet fragment relocation. Cladding heatup rate and 

temperature, either with or without a blowdown, are the primary factors that 

influence burst shape and dimensional changes.

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(1): 

PK(2): UK(0):

Data, Judgement 

Data, Experience, Judgement 

No votes.

Appendix D-72
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter)

Definition .. and Rationale ".rtan Um ........ Ann • ... '...11........ 1 . t .. .. T •"A-4 -l.I

Simulation of fuel 

relocation

Conduct of Test-During 

Pre- and post-burst test 

phases (2)

Look at the impact of fuel fragment relocation on the cladding temperature during the 

high temperature oxidation phase and the quenching phase. 

H(1) Data of fuel relocation determines the impacted phases. 

M(3) Needs in pile test to be prototypical (heating source should come from the 

fuel). If the objective is as speculated above, this test would help to 

characterize at which point in time the bulk of the relocation occurs. 

However, most rods that balloon also burst and it is not clear that a separation in time would significantly affect the LOCA performance (i.e., whether 

relocation occurs instantaneously to fill the ballooned region as opposed to 

instantaneous relocation on burst). Burst shape and dimensional changes are 

influenced by clad phase at the time of ballooning and burst. 

L(O) No votes. 

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(1): 

PK(3): 

UK(0):

Data, Judgement 

Data, Calculation, Judgement 

No votes.

Appendix D-73
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Simulation of fuel 

relocation

Conduct of Test-During 

Internal pressure and 

moles of gas

The amount of gas in the rod upper plenum, for a given initial pressure in the test rod. 

H(3) Driving force for relocation, together with gravity. It is crucial to have a 

pressure evolution representative of a full-length rod. Internal gas pressure is 

the driving force for fuel fragments relocation. To be prototypical the amount of 

gas within the rod prior to the test has to be maintained constant. The 

internal pressure is a measured parameter, not an input data. Initial pressure is 

the primary factor that determines the burst temperature and shape and 

potential release of fuel particles from rim zone at burst. Plenum gas inventory 

is a secondary factor. 

M(1) If gas flow is the primary relocation mechanism, then an accurate simulation of 

that gas flow would be needed to obtain the most meaningful results. However, 

it is anticipated that similar relocation behavior would be obtained over a 

relatively wide range of gas flows. 

L(O) No votes. 

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(0): 

PK(4): UK(0):

No votes. 

Data, Calculations, Judgement 

No votes.

Appendix D-74
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty) 

Simulation of fuel Conduct of Test-During During ballooning and after burst, the fuel rod vibration induced by the flow can favour 

relocation Flow induced vibration crumbling of the fuel pellet stack. 

H(O) No votes. 

M(2) Fuel column axial gaps have been observed to form and continue during normal 

reactor operation. This results suggests that fuel column shakeout is not likely 

with normal flow-induced vibration even over very extended periods. It is 

further noted that with cladding perforation, steam ingress will promote fuel 

pellet oxidation that has been observed, with failed fuel during normal reactor 

operation, to cause effective blockage within the fuel rod to preclude fuel 

downward fuel pellet fragment motion, again overriding the effects of flow 

induced vibration. Secondary driving force. 

L(2) Potential impact of rod vibration is expected to be small. Ballooning and burst 

occur after blowdown, and steam-flow-induced vibration during and after 

blowdown would be insignificant. 

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N 

K(O): No votes. 

PK(2): Data, Judgement 

UK(1): Judgement

Appendix D-75
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Simulation of fuel 

relocation

Conduct of Test-During 

Exterior rod constraints

Manner in which test specimen is constrained by surrounding rods to simulate potential in-reactor behavior. 

H(1) Prior ballooning experiments have shown that coplanar ballooning is not 

likely, and therefore balloons may not be constrained by adjacent ballooned 

sections. However, the constraints provided by adjacent non-ballooned rods can 

still provide a significant restriction on the amount of cladding ballooning and 

corresponding fuel relocation. 

M(1) Rod constraints during ballooning may affect the fuel distribution at the 

relocation site. 

L(2) The purpose of these tests is to analyze the separate effect of fuel fragment 

relocation. Exterior constraints influence ballooning shape to some extent. 

Fuel: Y(1): Most modem BWR fuel designs use part-length fuel rods resulting in zones 

where there is a significant gap between adjacent rods (because rods in certain 

lattice locations terminate at a lower elevation). This design feature may 

permit greater ballooning and relocation at those elevations. However, the 

fuel rods at those peculiar locations would correspondingly experience a 

circumferential temperature gradient, which is known to reduce the resulting 

burst strain. 

Clad: N 

Reactor: N 

Burnup: N

K(0): 

PK(4): 

UK(O):

No votes. 

Data, Judgement 

No votes.

Appendix D-76
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Simulation of fuel 

relocation

Conduct of Test-During 

Balloon size and burst 

size

Determination of the dimensions of the ballooned area and the cladding breach during the test. 

H(4) Affects the amount of relocated fuel in the balloon. The balloon and burst size 

represents the maximum potential volume for relocation. Directly influence the 

potential for fuel relocation, slumping, and release at and after burst. 

M(O) No votes. 

L(1) Balloon size and burst size are measured after the test. No need to measure it 

on-line

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(2): 

PK(2): 

UK(O):

N 

N 

N 

N 

Judgement 

Data, Judgement 

No votes.

Appendix D-77
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty) 

Simulation of fuel Conduct of Test-During Longitudinal dimension of the fuel rod segment to be tested. 

relocation Length 

H(2) The driving force for fuel fragments relocation is the internal gas pressure in 

the plenum. For high burnup fuel rods the axial gas transport is significantly 

impaired. A short rod would favor the plenum gas participation The rod 

length has to be prototypical to avoid experimental bias. At the least, the 

length between two grids must be tested. 

M(1) The amount of fuel above the ballooned/burst section defines the potential fuel volume to be relocated. However, the size of the ballooned/burst region 

defines the maximum possible relocated fuel volume. Therefore, if the 

ballooned/burst location can be defined with reasonable certainty, sufficient 

length can be provided above that region to enable prototypic relocation. 

L(1) Length more than about 15 times of the pellet length (6 inches) is sufficient. 

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N 

K(1): Calculation 

PK(3): Data, Judgement 

UK(O): No votes.

Appendix D-78
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Simulation of fuel 

relocation

Conduct of Test-PTE 

Granularity of dispersed 

material

Granularity of dispersed fuel fragments is measured to get relevant information on the 

fuel density in the relocated fuel fragments zone. 

H(3) The equivalent fuel density of the relocated fragments allows codes to 

simulate the local overheating of the cladding. Major factor that influences 

the potential for fuel relocation and release. 

M(1) Smaller pellet fragments would be expected to result in easier fuel movement 

and possibly a higher density of relocated fuel. However, pellet cracking 

patterns are established early in life and do not vary greatly with increased 

exposure, so a widely varied granularity of material, prior to dispersal, is not 

expected. 

L(O) No votes.

Fuel: 

Clad: 

Reactor: 

Bumup: 

K(O): 

PK(4): 

UK(0):

N 

N 

N 

N 

No votes. 

Data, Judgement 

No votes.

Appendix D-79
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty) 

Simulation of fuel Conduct of Test-PTE Non-intrusive measurement of the temperature differences of the tested fuel rod. relocation Thermography 

H(1) Provides the fuel distribution in 3D. 

M(O) No votes. 

L(2) Low added value 

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N 

K(O): Data 

PK(2): Data, Judgement 

UK(1): Judgement

Appendix D-80
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Im ortance, Applicability, and Uncertainty)

Simulation of fuel 

relocation

Conduct of Test-PTE 

Thermal diffusivity of 

rubble bed

Self defined.

H(1) Output parameter. 

M(1) Probably difficult to do, but would be useful in quantifying the effective 

thermal properties of the rubble mass (This assumes that in the 

ballooned/burst region if the material is still there - it may be worthwhile to 

capture this just prior to burst although there may not be significant relocation 

at that time if gas flow is the primary relocation mechanism), otherwise this 

is best done analytically. 

L(1) No rationale provided.

Fuel: 

Clad: 

Reactor: 

Burnup: 

K(0): 

PK(2): UK(1):

N 

N 

N 

N 

No votes. 

Data, Judgement 

Judgement

Appendix D-81
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Simulation of fuel 

relocation

Conduct of Test-PTE 

Strain profile of 

cladding as f(0,z)

Measure the shape and the size of the ballooned area of the tested fuel rod. 

H(3) The purpose of this test is to assess the amount and characteristics of 

relocation. A determining aspect of that process is the amount of ballooning (free volume to which the fuel may relocate), and therefore this volume 

should be known in any assessment of relocation characteristics. Note that the 

circumferential variation of cladding strain should also be determined. Axial 

variation of clad circumferential strain is a parameter that directly influences 

the potential for fuel relocation and slumping. 

M(1) Will give some indications on potential impact of the balloon the shape (magnitude and extension) on the amount of relocated fuel. 

L(O) No Votes. 

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N

K(l): 

PK(3): 

UK(0):

Judgement, Calculation 

Data, Judgement 

No votes.

Appendix D-82
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued)

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty) 

Simulation of fuel Conduct of Test-PTE Size of the opening in the cladding after the burst. 

relocation Burst size 

H(3) This is taken to be the effective surface area of the bulged region that was 

removed as a result of the burst. Similar to the preceding item, the hole size 

will be a determining factor in the amount of relocated fuel retained within 

the ballooned region. Burst opening size and burst circumferential strain are 

parameters that directly influence the potential for fuel relocation and release 

at and after burst. 

M(O) No votes. 

L(1) If the internal pressure is not maintained (prototypal case) the opening is 

small. 

Fuel: N 

Clad: N 

Reactor: N 

Bumup: N 

K(2): Data 

PK(2): Data, Judgement 

UK(O): No votes.

Appendix D-83
Table D-1. PWR and BWR LOCA. Category D- Separate Effect Testing (continued) 

Subcategory (Test type) I Phenomena (Parameter) I Definition and Rationale (Importance, Applicability, and Uncertainty)

Simulation of fuel 

relocation

Conduct of Test-PTE 

Material balance (in-rod 

and dispersed)

Determination of the mass of fuel remaining within the tested fuel rod and the mass that left the fuel rod through the rupture. 

H(2) This is the primary result to be quantified in this test series, to be correlated with the ballooned region and burst size. It is the amount of lost material that is of interest as it could possibly contribute to such effects as flow blockage, etc.

IvikuJ 

L(2) This information is covered by the local measurement of the fuel density.

Fuel: N 

Clad: N 

Reactor: N 

Burnup: N

K(1): 

PK(2): 

UK(1):

Judgement 

Data, Judgement 

Judgement

Appendix D-84
APPENDIX E

EXPERIMENTAL DATABASES 

The experimental databases identified in Section 4 of this report are further 

discussed in this appendix. The author of each contribution is identified. The 

contributed documentation exhibits some style differences. References providing 

additional details for each test program are provided at the end of each contributed 

entry. 

E-1. Separate Effect Tests 

E-1.1. Cladding Tests (United States) 

The information regarding this test series was provided by panel member A. Motta 

of the Pennsylvania State University and M. Billone of Argonne National 

Laboratory. 

Argonne National Laboratory (ANL) and The Pennsylvania State University (PSU) 

are working together on a program to investigate cladding properties and to test 

loss-of-coolant accident (LOCA) acceptance criteria at high burnups.('"' 3) Although 

the main focus of the program is to investigate fuel behavior under LOCA 

conditions, related mechanical properties testing is being done under both LOCA 

conditions and rod ejection accident (REA) conditions. 

Ring Tensile Tests. The tests at relatively low temperatures and high strain rates 

appropriate for rod ejection accident conditions are being performed. The objectives 

are two-fold: to understand the degradation in cladding failure behavior at high 

bumup and to obtain stress-strain relationships that will serve as inputs to codes. 

High-burnup fuel rods (about 70 GWd/t) from the H. B. Robinson PWR are expected 

to be available for these tests along with related archival fresh tubing. Although the 

fuel has not arrived at the time of this writing, high-burnup specimens (about 50 

GWd/t) from TMI-1 are available and have been used for preliminary testing along 

with non-irradiated Zircaloy-4 tubing. Oxidation kinetics and phase transformation 

characteristics are also being measured on high-burnup specimens. 

Axial Tensile Tests. Similar testing will be done on axial tensile specimens 

electromachined from de-fueled portions of irradiated fuel rods and from non

irradiated tubing specimens. These tests will be performed over the same 

temperature range and strain-rate range as the ring-stretch tests mentioned above. 

The combination of the axial and the hoop stress-strain properties will allow 

validation and improvement of the models used in fuel rod codes for predicting the 

mechanical behavior of an anisotropic alloy such as Zircaloy. 

Biaxial Tube Burst Tests. Biaxial tube burst tests are the most informative and the 

most difficult to perform, and they consume the largest amount of specimen 

material, which is a significant consideration when testing irradiated fuel material. 

These tests will be done in a more limited temperature range of 300 'C-400 °C, but 

they will explore the effects on deformation and failure of stress biaxiality ratios

E-1
from 1:1 to 2:1 at high strain rate. In principle, the tests can be run with the fuel 

intact or with the fuel removed. Some tests will be run with the fuel removed to 

generate baseline data for code validation along with data that can be compared to 

other such studies on non-irradiated and medium-burnup cladding. 

Post Quench Ductility Tests. Four-point bending tests will be performed after 

quenching to determine remaining ductility in oxidized cladding. Other post

quench ductility tests are being considered at the time of this writing. 

Oxidation Kinetics. Oxidation kinetics are being measured on high-burnup 

specimens to determine the effect of pre-LOCA oxidation (corrosion) on the 

allowable oxidation for regulatory considerations. Phase transformation 

characteristics and their effect on high-temperature oxidation are also being studied. 

References for Cladding Tests (United States) 

1. A. B. Cohen, et al., "Modified Ring Stretch Tensile Testing of Zr-lNb Cladding," 

Proc. USNRC Water Reactor Safety Information Meeting, NUREG/CP-0162 2, 

133-149 (October 20-22, 1977). 

2. T. M. Link, D. A. Koss and A.T. Motta, "Failure of Zircaloy Cladding under 

Transverse Plane-strain Deformation," Nuclear Engineering and Design 186 379

394 (1998). 

3. D. W. Bates, et al., "Influence of Specimen Design on the Deformation and 

Failure of Zircaloy Cladding," Proc. ANS International Meeting on Light Water 

Reactor Fuel Performance, Park City, Utah, 1201-1210 (April 10-13, 2000). 

E-1.2. Cladding Mechanical Property Tests (Japan) 

The information regarding this test series was provided by panel member T. Fuketa 

of the Japan Atomic Energy Research Institute (JAERI). 

Mechanical property tests for fuel cladding have been carried out at JAERI, applying 

different testing method and specimen configuration according to the purpose.(') 

They are ring tensile tests, axial tensile tests, and tube burst tests. The most general 

and reliable method to quantitatively examine the mechanical property of materials 

is the uniaxial tensile test. Therefore, the test is applied to examine the mechanical 

property changes after temperature transients expected in LOCAs. The specimens are 

electro machined from artificially hydrided Zircaloy-4 cladding, annealed at 

temperatures ranging from 873 °K-1373 'K for 300 s in argon flow, and tensile tested 

at room temperature and 573 'K. Strength and ductility changes are obtained as a 

function of hydrogen concentration and annealed temperature. Recent test results 

showed that 9-quenched Zircaloy cladding with a high hydrogen concentration over 

500 wtppm exhibits very low ductility at room temperature. The data obtained will 

be used to evaluate results from thermal shock tests and be provided as input to 

codes. Similar tests are planned with oxidized and hydrided specimens and high 

burnup specimens. The modified ring tensile test with machined specimen 

currently being developed might be applied in those tests, because of an advantage 

on the necessary specimen volume.

E-2
References for Cladding Mechanical Property Tests (Japan) 

1. T. Fuketa, F. Nagase, T. Nakamura, H. Sasajima and H. Uetsuka, "JAERI 

Research on Fuel Rod Behavior during Accident Conditions," Proceedings of 

the 27th Water Reactor Safety Information Meeting, Bethesda, Maryland, U.S.A., 

October 25-27, 1999, NUREG/CP-0169, 341-354 (2000). 

E-1.3. LOCA Separate Effects Tests (France) 

EDGAR program. The separate phenomena which are covered by the EDGAR 

programs are: (1) the phase transformation kinetics during heat-up and cooling 

down, and (2) the mechanical properties of the cladding (creep strain rate, time to 

rupture, diametral strain at rupture). Based on the hypothesis that most of the 

irradiation defects are annealed during the heat-up, some tests are made on 

prehydrided specimens to simulate the impact of irradiation on both mechanical 

and oxidation behaviors. Tests on pre-hydrided samples are performed at different 

hydrogen content levels. 

The EDGAR program is complementary to the LOCA integral tests (such as the 

LOCA criteria test performed at ANL). The integral tests programs are devoted to the 

study of the whole behavior of cladding materials during a LOCA. The knowledge of 

the elementary behaviors of the claddings will allow the interpretation of these 

integral LOCA tests and, more specifically, the extrapolation of these tests results to 

other in-reactor LOCA transients. 

TAGCIS-TAGCIR-HYDRAZIR-CINOG Program Test Series. The French LOCA test 

program consisted of four different test series, with their specific purposes. All the 

tests involved well-defined double-sided oxidation and thermal shock tests on 

empty 17X17 Zr-4 cladding samples in a steam environment: 

- The TAGCIS test series (1991-1993) included more than 110 thermal shock tests 

on as-fabricated or pre-oxidized Zr-4 cladding samples. This series investigated 

the thermal shock behavior of non-irradiated cladding having an initial 

corrosion similar to the end-of-life corrosion of a high burn up fuel rod. 

- The TAGCIR test series (1993-1996) included 25 oxidation and thermal shock tests 

on irradiated Zr-4 cladding samples from high burnup fuel rods irradiated for 5 

cycles up to 60-63 GWd/t in a commercial EDF nuclear power plant (corrosion 

from 60 to 100-120 gm). 

- The HYDRAZIR test series (1996-1999) included oxidation and thermal shock 

tests (at different cooling rates) on non-irradiated as fabricated and pre-hydrided 

Zr-4 cladding samples (0 to 5000 ppm H). One of the objectives of this series was 

to investigate the impact of pre-acddent hydriding on transient oxidation and 

quenching behavior. The interest in hydrogen impact came from the TAGCIR 

test series that showed that the behavior of irradiated zircaloy under LOCA 

conditions could possibly be related to the hydrogen.

E-3
- The CINOG test series (1997- 2001) included oxidation and thermal shock tests on 

as-fabricated Zr-4 samples. This series was devoted mainly to the licensing of 

new or advanced alloy under LOCA conditions. 

The experimental process for assessing the ECR values for the TAGCIR test series 

was based on the metallographic measurements of the different metallurgical phases layers thickness and the subsequent calculation of the total amount of oxygen 

present in the cladding. In the more recent HYDRAZIR and CINOG test series, the 

same test facility provided, in terms of the measured local cladding temperature, both the high temperature oxidation kinetic through weight gain measurements 

and the final quenching behavior. 

In all cases, the cladding was assumed to have failed as soon as it could not 

withstand an internal over-pressure of 1 bar of argon and small bubbles were 

detected on the sample surface. This methodology was extremely conservative since 

such a threshold was well below the fragmentation limit. 

E-2. Integral Tests 

E-2.1. LOCA Tests (United States) 

The primary purpose of these tests is to evaluate the performance of high burnup 

fuel relative to the NRC cladding embrittlement criteria defined in 10CFR50.46. The 

criteria relevant to this research effort are: 

1. The calculated maximum fuel element cladding temperature shall not exceed 

2200 OF 

2. The calculated total oxidation of the cladding shall nowhere exceed 0.17 times 

the total cladding thickness before steam oxidation. 

Within the ANL test plan, the LOCA integral tests will be conducted on fuel rod 

segments (300-mm long) with the as-irradiated cladding outside diameter (OD) and 

inside diameter (ID) oxide layers and the fuel intact. In this way, the high burnup effects of the oxide layers, the associated hydrogen pickup due to waterside 

corrosion, and the fuel cladding contact and/or bonding will be present in the tests. 

The central 150-200 mm of the test sample will be uniformly (within 25 'Q heated. 

The specimen will be pressurized, stabilized at 300 'C, heated at 5 'C/s to 1204 °C, held 

at 1204'C at a time corresponding to a calculated equivalent cladding reacted (ECR) of 

'17%, slow cooled to 750 'C-800 0C, and water-quenched. The calculation of the ECR 

vs. time at 1204'C will be made using the ANL "A Model", with model parameters adjusted based on the results of the oxidation test results at 1204 °C A minimum of 

three tests will be run. The time for the first test will be set to yield ECR = 17% 

including the in-reactor-formed oxide layers. The second test will be run at a longer time corresponding to ECR>30% in an effort to produce thermal-shock failure of the 

cladding. Based on these two results, an intermediate time-ECR test will be run (e.g., 17% ECR excluding the in-reactor oxide layer) to help determine margin to failure. 

Additional tests (up to 3) may be run based on what is learned from the first three 

tests.

E-4
As the planned tests with high burnup fueled cladding are first-of-a-kind relative to 

previous tests that have been conducted, there are other important responses that 

will be studied to resolve the effects of high burnup operation on LOCA-relevant 

phenomena. During the 5 'C/s rise to 1204 'C, the cladding will balloon and burst. Of 

interest from the ANL tests are the circumferential magnitude and axial extent of 

the ballooning, the geometry of the burst, possible fuel particle relocation to the 

ballooned and burst region, and the effects of these phenomena on the 

circumferential and axial temperature profile. To the extent practical, these 

phenomena will be observed, described and quantified. In terms of post-test 

analyses, the ECR, the phase distribution, and the hydrogen content will be 

measured in the ballooned-and-burst region and either in the thermal-quench

failed region (if different from the ballooned-and-burst region) or in a 

non-ballooned, non-burst, non-failed axial location for the tests in which thermal

shock failure does not occur. The ECR values based on data will be compared to the 

calculated ECR values to determine the degree of conservatism associated with the 

models. 

The rods, grid spans, locations within the grid spans and the times (in terms of ECR 

values for Limerick BWR LOCA integral tests are identified in Table E-1. The results 

of the tests using the irradiated fueled cladding samples will be compared to results 

obtained with non-irradiated cladding samples. Another option that is available for 

isolating the effects of tight fuel-cladding bonding is to defuel samples, fill the 

cladding with alumina or zirconia pellets and run these samples through the same 

the same temperature history. However, because of the anticipated low cladding 

oxide thickness and hydrogen content, it is difficult to justify such tests. The tests on 

the BWR Zircaloy-2 archive (or near archive) cladding can be run outside the hot 

cell in the LOCA Criteria Mockup. 

Table 3-1. LOCA Integral Test Matrix for Limerick BWR Fuel Rods 

Material Fuel Grid Span Test Time in Terms of Post-Test 

Condition Rod ID (Location from Calculated ECR Examinations bottom in nun) II 

Irradiated F9 5 (70-370) =-17% M, 0, P 

6 (70-370) >30% M, 0, P, F 

J4 6 (70-370) 17%<ECR<30% M, 0, P 

5(70-370) TBD TBD 

J6 6(70-370) TBD TBD 

5(70-370 TBD TBD 

Archive --- --- 17% M, O, P 

17-30% M, 0, P 

>30% M, 0, P, F 

(All tests are to be run at a peak cladding temperature (PCT) of 1204 'C; TBD = to be determined, M 

metallography, 0 = oxygen analysis, H = hydrogen analysis, P = profilometry, F = fractography)

E-5
E-2.2. LOCA Tests (Japan) 

The information regarding this test series was provided by panel member T. Fuketa 

of the Japan Atomic Energy Research Institute (JAERI). 

This test series(') examines the rod-burst, oxidation and reflooding stages, similar to 

the tests performed at ANL. To systematically investigate the influence of burnup 

extension on failure-bearing capability of the cladding during quench, pre-hydrided, 

pre-oxidized, and/or irradiated claddings as well as high burnup claddings are tested. 

A 600-mm-long cladding with end plugs is pressurized to 5 MIPa with Argon gas and 

heated in a steam flow at a rate of 10 K/s to a target temperature ranging 800 °C to 

1250 'C. Annular alumina pellets are contained in the cladding to simulate the heat 

capacity of UO2 pellets. After the isothermal oxidation period at the target 

temperature, the cladding is slowly cooled to 700 'C and finally water-quenched by 

reflooding from the bottom end. Cladding axially expands during heat up and 

oxidation, and shrinks during cooling and quench. The shrinkage of the cladding 

may be restricted between spacer grids in the bundle geometry. Then, the cladding is 

quenched fixing both ends, for a conservative condition, to examine the influence of 

restricting the cladding shrinkage and consequent tensile loading during the quench 

in some part of the present tests. 

Oxide layer thickness, hydrogen content and increase of circumference by ballooning 

are measured in the post-test examination to characterize the cladding failure. 

Failure-bearing capability will be evaluated based on ECR calculated both from 

oxidation temperature-time and measured oxide layer thickness. 

Tests have now been performed with artificially hydrided cladding (non-irradiated) 

to examine the separate effect of hydrogen absorption during operation. The test 

results obtained thus far indicate that the restriction of the cladding shrinkage 

during quench has a large influence on the failure boundary for the oxidized 

condition and it is even larger in pre-hydrided claddings. Therefore, the tests with 

pre-hydrided claddings under controlled axial loading condition will be added. 

Preparation of high burnup cladding specimens is now in progress, and the tests 

with high burnup PWR fuel claddings (about 42 MWd/t) are to be started beginning 

in 2002. 

References for LOCA Tests (Japan) 

1. F. Nagase, M. Otomo, M. Tanimoto, and H. Uetsuka, "Experiments on High 

Burnup Fuel Behavior under LOCA Conditions at JAERI," Proc. ANS 

International Meeting on Light Water Reactor Fuel Performance, Park City, 

Utah (April 10-13, 2000). 

E-2.3. BWR Transient Dryout and Rewet Tests 

The anticipated transient without scram (ATWS) instability and the LOCA have 

been identified as key events for the evaluation of fuel performance for a BWR. In 

ATWS instability the BWR will be at low flow for natural circulation and 

experience power oscillations. During these oscillations, the high power fuel 

bundles may undergo periodic boiling transition and rewet following each power

E-6
pulse. As long as the PCT remains below the minimum film boiling temperature, 

rewet will occur and excessive fuel heat up is avoided. However, if the cladding 

temperature exceeds the minimum film boiling temperature (approximately 600 'C 

[1100 °FD) following a power pulse the fuel may not rewet and substantial fuel heat 

up can occur. The following material is based on information found in Refs. 1 

through 31. 

Data for transient dryout, post dryout heat transfer and transient rewet have been 

obtained since the mid 1960s. The data include simple geometry tests as well as full 

scale simulated fuel bundles. 

Simple geometry data', have typically been obtained in tubular and annular 

geometries and include steady state as well as transient tests. These tests typically 

give well defined thermal hydraulic data and are excellent for model qualification. 

They do not, however, provide information on the cross sectional variation of 

thermal hydraulic conditions in a rod bundle. The maximum PCT for these tests 

goes well beyond the minimum film boiling temperature, where rewet is not 

obtained. These tests therefore provide valuable information on boiling transition, 

film boiling heat transfer and rewetting. 

Similar tests have been obtained in simple rod bundles,4

'°' 4 typically 4x4 rod 

bundles. In these tests both steady state and transient tests have been performed. The 

steady state tests were used to obtain information on film boiling heat transfer, 

while the transient tests were used to obtain additional information on transient 

dryout and rewet. The transients were either simple power and flow transients 

where either the power was temporarily raised or the flow temporarily reduced to 

obtain a boiling transition, or they were simulation of a reactor turbine trip or 

recirculation pump trip. These tests also give PCTs beyond the minimum film 

boiling temperature and provide valuable information on boiling transition, film 

boiling heat transfer and rewetting. 

BWR fuel vendors perform extensive critical power tests for each new fuel product 

that is developed. Steady state critical power data over a range of parameters 

covering normal steady state operation as well as the expected range of parameters 

for operational transients is obtained. These data are used to develop a fuel-type

specific critical power correlation. In addition a few transient tests are usually 

performed in order to demonstrate the applicability of the correlation under 

transient conditions.'6','81 1 ,129 '

2 8 29

,'03 1 The transient tests are simulated turbine trip 

and recirculation pump trip events, and in one instance a reactor instability was 

simulated. Since the transient tests are intended to demonstrate the applicability of 

the critical power correlation under transient conditions, the PCT typically does not 

exceed the saturation temperature by more than 100 'C-200 0C and thus does not 

provide data beyond the minimum film boiling temperature. 

Numerous LOCA experiments have provided information on transient dryout, 

film boiling heat transfer and transient rewet. These tests include data from the 

BDHT,15 TLTA, 16 FIST, 17,18 FIX, 9, 0,21=,23TBL, 24 and ROSA-III 25

-

26 test facilitiJes, which 

are all scaled simulations of BWRs. The upper range of PCTs for these tests is 

approximately 870 °C (1600 'F). High temperature data up to 1150 0C (2100 °F) have

E-7
been obtained in GE's core spray heat transfer test facility and from the GOTA test 

facility and similar facilities at Hitachi and Toshiba. 20 

Finally in-pile tests have been performed, where nuclear fuel rods have been subject 

to boiling transition during power and flow transients. Even though the primary 

purpose of these tests was to evaluate the thermal and mechanical response of the 

fuel, these tests also provide valuable data on transient dryout and rewet. The early 

data in the van Houten report'3 were collected for exposures up to 20 GWd/t and 

PCTs up to 1700 'C. The later data from the Halden test reactor13 had exposures up to 

40 GWd/t and PCTs up to 950 °C. 

The transient dryout and rewet tests are summarized in the Table E-2. 

Table E-2. Transient Dryout and Rewet Tests.

Geometry Test Type PCT References 

Tubular and Annular Steady State and Transient 850 °C 1, 2, 3 

Simple Rod Bundles 

4X4 Rod Bundles Steady State Film Boiling 715 -C 4, 5, 14 

Flow and Power Transients 

Simulated Turbine and Pump 

Full Scale Rod Bundles 

Simulated Turbine and Pump 6, 7, 8, 11, 19, 

Trips for 8X8, 9X9 and 1OX10 27, 28, 29, 30, 

Rod Bundles 31 

LOCA 

Scaled Simulation of a BDHT, TLTA, FIST, FIX, 870 0C 9, 10, 15, 16, 

BWR. TBL, ROSA 17, 18, 21, 22, 

23, 24, 25, 26 

Core Spray Heat CSHT, GOTA. Toshiba, 1150 'C 20 

Transfer Hitachi 

rIn-Pile Data 

Flow and Power Transients 1700 -C 12, 13

Note: Minimum Film Boiling Temperature 600 0C. 

References for BWR Transient Dryout and Rewet Tests 

1. A. W. Bennet, et. al., "Studies of Burnout in Boiling Heat Transfer to 

Round Tubes with Non-Uniform Heating," Atomic Energy 

Establishment document AERE-R55076, Harwell (1966).

Water in 

Research

2. A. Era, et. al., "Heat Transfer in the Liquid Deficient Region for Steam Water 

Mixtures at 70 kg/cm2 Flowing in Tubular and Annular Conduits," Centro 

Infomazioni Studi Esperienze, SpA document CISE-R-184 (1966).

E-8
3. D. C. Groeneveld, "An Investigation of Heat Transfer in the Liquid Deficient 

Region," Atomic Energy Canada, Ltd. Document AECL3281, Ottawa (1966). 

4. G. L. Yoder, et. al., Dispersed Flow Boiling Heat Transfer Data Near Spacer Grids 

in a Rod Bundle, Nuclear Technology 60, 304-313 (1983). 

5. G. L. Yoder, D. G. Morris, and C. B. Mullins, "Rod Bundle Burnout Data and 

Correlation Comparisons," Nuclear Technology 68, 355-369 (1985). 

6. 0. Nylund, et. al., "Post Dryout in Connection with BWR Main Circulation 

Pump Trip," European Two-Phase Flow Group Meeting, Munich (June 1986). 

7. R. Persson, L. Nilsson, and H. Gustafson, "FD(-lI - Transient Dryout Tests 

Summary Report," Studsvik Report NP-86/35, Nykoping (1986). 

8. "General Electric BWR Thermal Analysis Basis (GETAB): Data, Correlation and 

Design Application," General Electric document NEDO-10959A (January 1977). 

9. L. Nilsson, R. Persson, and H. Wijkstrom, "FDX-II - LOCA Blowdown and Pump 

Trip Heat Transfer Experiments," Summary Report for the First Experimental 

Period," Studsvik Report NP-83/325 (1984). 

10. L. Nilsson and R. Persson, "FIX-Il - LOCA Blowdown and Pump Trip Heat 

Transfer Experiments," Summary Report for the Second Experimental Period, 

Studsvik Report NP-85/42, 1985. 

11. J. S. Nunez and B. Matzner, "Critical Power, pressure Drop and Transient Tests, 

81 Rod Bundle with KWU Spacers AH16 and AH17," General Electric document 

NEDC-21984 (1978). 

12. M. A. McGrath, B. C. Oberlander Mykiebust et. al., "Dryout Fuel Behaviour Tests 

in 1FA613, Summary of Inpile Results and PIE Data," EHPG Meeting, Loen (May 

1999). 

13. R. van Houten, "Fuel Rod Failure as a Consequence of Departure from Nucleate 

Boiling or Dryout," United States Nuclear Regulatory Commission document 

NUREG-0562 (June 1979). 

14. B. S. Shiralkar, et. al., "Transient Critical Heat Flux - Experimental results, 

GEAP-13295, AEC Research and Development Report (1972) 

15. K. V. Moore, et. al., "Analysis of the GE BWR Blowdown Heat Transfer Program, 

Test 4906 (AEC Standard Problem 4)," Electric Power Research Institute 

document EPRI-280 (January 1975). 

16. L. S. Lee, G. L. Sozzi, and S. A. Allison S. A., "BWR Large Break Simulation Tests 

- BWR BlowdownlEmergency Core Cooling Program, "General Electric 

document GEAP-24962-1, NUREG/CR-2229, EPRI NP-1783, Volumes 1 & 2 

(March 1981).

E-9
17. W. S. Hwang, Md. Alamgir, and W. A. Sutherland, "BWR FIST Phase I Test 

Results," United States Nuclear Regulatory Commission document NUREG/CR

371 1 (November 1983). 

18. W. A. Sutherland, Md. Alamgir J. A. Findlay, and W. S. Hwang, "BWR Full 

Integral Simulation Test (FIST) Phase II Test Results and TRAC-BWR Model 

Qualification," United States Nuclear Regulatory Commission document 

NUREG/CR-4128, GEAP-30876, EPRI NP-3988 (June 1985). 

19. X. M. Chen, J. G. M. Andersen, L. Klebanov, and T. Anegawa, "A Transient 

Subchannel Analysis Method for BWR Fuel Bundles," NUTHOS-5 (1997) 

20. F. D. Shum, et. al., "SAFER Model for Evaluation of Loss-of-Coolant Accidents 

for Jet Pump and Non-Jet Pump Plants, Volume 1, SAFER - Long Term 

Inventory Models for BWR Loss-of-Coolant Analysis," General Electric 

document NEDO-30996A (October 1987). 

21. L. Nilsson and P. A. Gustafsson, "FIX-II - LOCA Blowdown and Pump Trip Heat 

Transfer Experiments, Summary Report for Phase 2: Description of Experimental 

Equipment Part 1. Main Text," Studsvik Report NR-83/238 Part 1 of 3, (February 

1983). 

22. L. Nilsson L., et al., "FIX-II LOCA Blowdown and Pump Trip Heat Transfer 

Experiments, Experimental Results from LOCA Test No. 5052," Studsvik Report 

NR-83/323, (March 1984). 

23. 0. Sandervag, and D. Wennerberg, "FIX-II - Experimental Results of Test 3025 

(ISP-15)," Studsvik Report NR-83/283 (July 1983). 

24. M. Naitoh, M. Murase, and R. Tsutsumi, "Large Break Integral Test with TBL- 1 

(Hitachi BWR Integral Facility)," 9'th Water Reactor Safety Information Meeting, 

National Bureau of Standards, Gaithersburg, Maryland (October 1981). 

25. Y. Anoda,et. al., "ROSA-III System Description Report Fuel Assembly No. 4," 

JAERI-M 9363 (February 1981). 

26. Y. Anoda et. al., "Experimental data of the ROSA-III Integral test Run 912 (5% Slit 

Break Test without HPCS Actuation," Japan Atomic Energy Research Institute 

document JAERI-M - -3 (July 1982). 

27. G. Preusche and Z, Stosic,. "Post-Dryout Methodology for BWRs," Siemens 

AG/KWU, Fachtagung der KTG-Fachgruppe "Brennelemente und 

Kernbauteile", Karlsruhe, Mats (2000). 

28. M. Kitamura, T. Mitsutake, K. Kamimura, N. Abe, S. Morooka, and J. Kimura, 

"BWR 9X9 Rod Assembly Thermal Hydraulic Test (4), Cladding Temperature 

Behavior During Power Increase," 1998 Fall Meeting, Atomic Energy Society of 

Japan (1998). 

29. M. Kitamura, T. Mitsutake, K. Kamimura, N. Abe, S. Morooka, and J. Kimura, 

"BWR 9X9 Rod Assembly Thermal Hydraulic Test (5), Cladding Temperature 

Behavior During Flow Coast Down Event," 1998 Fall Meeting, Atomic Energy 

Society of Japan (1998).

E-10
30. N. Abe, S. Mimura, S. Ebata, S. Morooka and T. Anegawa, "Post-BT Test Analysis 

by Best Estimate Thermal-Hydraulic Code TRACG," ICONE (1999). 

31. J. G. M. Andersen, et. al., "TRACG Qualification," General Electric document 

NEDE-32177P Revision 2. (January 2000). 

E-2.4. Dryout Effects on High Burnup Fuel (Halden Reactor Project-Norway) 

The information regarding this test series was provided by panel-member 

W. Wiesenack. 

Background. The objective of the dry-out test series was to provide information on 

the consequences for fuel of short-term dry-out incidents in a BWR. The 

experimental method employed was, on an individual basis, to expose fuel rod with 

different burnups to single or multiple dry-out events; to follow this by either 

unloading or continued operation in the reactor; and to finish with post irradiation 

examination and testing with emphasis on fuel clad properties. The test series was 

co-sponsored by the Halden Project's joint program and TEPCO (Japan). 

Testing program. The test series comprised three loadings of IFA-613. Each rod was 

contained in a stainless steel channel within the rig so that the coolant conditions 

for each rod could be controlled individually. In this way separate dry-out scenarios 

were effected for each rod. Thermocouples attached to the surface of the test rods 

were used to monitor clad surface temperature and clad elongation was monitored 

by way of an extensometer. The first and second loading operated for a month after 

dry-out while the rods in the last loading were unloaded directly after the dry-out 

procedure. In neither case did any fuel failures develop. 

The in-pile dry-out experiments with the third (and last) set of fuel rods in IFA-613 

were completed in January '98 (HWR-552, HP-1036) and the post irradiation 

examination (PIE) on all eight rods in the three test series were finished in 

September '98 (Kjeller hot cell). 

Summary of results. In total, 2 rods with fresh Zr-2 and Zr-4 and 6 rods with clad 

pre-irradiated to 22-40 MWd/kg (Zr-2, Zr-2 with liner and Zr-4) were individually 

exposed to reduced or no-flow conditions in a heated light water loop within the 

Halden reactor. Dry-out occurred over the upper region of each rod, with 6 rods 

developing PCTs in the range 950 'C-1200 'C. 

An overview of the condition of the rods in terms of clad surface condition, rod 

dimensions and hydriding was achieved using non-destructive PIE techniques such 

as profilometry and neutron radiography. Clad and fuel microstructure and clad 

mechanical properties were investigated with destructive PIE techniques including 

ceramography, metallography, microhardness and ring tensile testing. It was 

observed that while dry-out had not affected the fuel microstructure, significant 

changes had been induced in the clad. These included high temperature corrosion 

resulting in moderate growth of the outer surface oxide layer and H 2 pick-up 

(hydriding formation). Some of the rods also exhibited uniform and localised clad 

creep-down into pellet-pellet interfaces and in the most severely tested rods that clad

E-11
had undergone the a to 1 phase transformation. This material exhibited reduced 

ultimate tensile strength and brittle fracture. However, significant improvements of 

ductility were observed in clad that had been exposed to less severe in-pile transients 

where a small a-phase grain structure was retained and hydrogen pick-up was 

minimal. None of the rods failed, during either the dry-out phase or the subsequent 

steady-state normal operation. 

Applications. The data obtained will be used to assess and modify existing 

rules/regulations in member countries on the continued operation with fuel 

elements subjected to short-term dry-out transients in boiling water reactors.

E-12
APPENDIX F

DESCRIPTION OF PWR FUEL AND CLADDING STATE AT HIGH BURNUP÷ 

The extended operational exposure that accompanies high bumup causes changes to 

the fuel and cladding that may affect the fuel rod's ability to withstand the accident 

without losing its integrity (Fig. F-i). These changes, which occur gradually over the 

life of the fuel rod, can be considered as initial conditions for the accident. 

There are many changes that occur to the fuel and cladding as a result of prolonged 

exposure to the irradiation field present in a reactor core, and to the corroding 

environment and high temperature. The combination of high temperature, 

radiation damage, transmutation, mechanical stresses and chemical reactions causes 

the microstructure of cladding and fuel to evolve considerably during reactor 

exposure. These changes in microstructure, microchemistry, and macroscopic 

characteristics of pellet and cladding are responsible for the changes in material 

behavior observed at high bumup. These changes are very complex and difficult to 

predict in a mechanistic fashion. Of the many changes to the fuel and cladding, it is 

important to discern which are of greatest importance to determining fuel rod 

behavior during an accident. We list some of the more important material 

degradation phenomena below, recognizing that the list may not be inclusive. The 

changes to the fuel and cladding are important to both pressurized water reactor 

(PWR) and boiling water reactor (BWR) fuel types. A discussion specific to the 

changes to BWR fuel and cladding is presented in Appendix G. 

F.1. Cladding Changes 

The main degradation mechanisms to Zircaloy-4 cladding include uniform 

waterside corrosion, hydriding, and radiation damage. 

Uniform waterside corrosion occurs throughout the reactor exposure. The corrosion 

rates depend on many factors including alloy chemistry and thermomechanical 

treatment, coolant chemistry, radiation-induced changes to cladding 

microchemistry, and irradiation temperature. For cladding with burnups in excess 

of 50 GWd/t , the oxide thickness can exceed 100 gm depending on fuel duty, i.e., 

power and temperature versus time and burnup. The burnup level at which any 

given oxide thickness is reached for a given alloy is dependent on the fuel duty. The 

more modern alloys such as ZIRLO and M5, can have lower corrosion rates than 

standard Zr-4 and low-Sn Zr-4 at similar burnup. All of the zirconium alloys 

examined to date show a change in corrosion rate when the 6xide exceeds a certain 

thickness (20-30 gm in thickness), which indicates a change in corrosion regime, 

termed breakaway corrosion. Therefore, it is likely that even the new modern alloys 

such as ZIRLO will eventually experience breakaway corrosion. The question with 

'See Acknowledgments for authorship information.

F-1
the new modem alloys is the bumup level at which breakaway corrosion will be 

observed.

Oxi de spa 11 at i( 

"* Extent of spa 

"• Hydridepre 

"* Blisterform 

Wa tersi decorrosion 

"• Corrosi on thickness 

"• Hydrogenpickup 

"* Non-rii formuit ies / 

inoxidethickness

I la tion 

4cipit at ion

ation" HydrogeninZirca loy 

- Hpickupfraction 

(hydrogen level in clad) 

• Hydrideorientation 

*Hydride radialI distribution 

(hydride rimth ickness 

Claddinginternal 

S~Loose corrosion (SSC) fuel pieces 

Large scale pell etcracking from 

ronna 1 operat ion 

"" Width of cracks- missingfuel "• Crack spacing 

Fuel-cla ddirg chemical bonding 

"• Claddingcreepdown& fuel swell ing 

"* Compound fonmation

Radi at ion damage tocladdirg 

"• Formationofdislocationloops 

"• Precipitate evoluti on 

"• Irradiat ioncreepandgrowth 

(wallthirnirg)

Development of rim region 

• Grain size 

SBubble size distributi on 

• Bibble pressure 

• Total gasinrim 

• Puccntent & sol id fission products

Higher hydrogen 

content in between 

Pellet

Conist it utit ve 

propert ies change 

"• Strainhardenirg 

"• Stra inra te hardening 

"• Deformation localization 

(dislocat ion channel forma tion)

Fissiongasi nfuel matrix 

"* F ission gas amount insoluti on 

"* Intergranular bubbles 

"* Intragranularbubbles 

"• Conduct iv ity degradat ion 

"• F issicn gasreleased durirg 

normal operation 

"• Porosity gas cantent

Fig. F-1. Fuel state at high burnup

F-2
For example, fuel that experiences a high fuel duty will experience breakaway 

corrosion at a lower burnup level than fuel with a lower fuel duty. One of the 

concerns with large oxide thicknesses is the higher probability of oxide 

delamination, whereby portions of the oxide layer are detached from the adherent 

oxide creating an oxide region with worse heat conduction characteristics. 

Ultimately the detached oxide can break off (oxide spalling), creating a thinner oxide. 

The associated temperature gradients created by spalling have been shown to 

influence hydride blister formation in the spalled region." The hydride blister is 

brittle, and its presence has been shown to affect overall cladding ductility. 

The main concerns associated with the uniform corrosion process are the potential 

for oxide spalling resulting in hydride blisters, which affect the overall cladding 

ductility; loss of thermal conductivity; non-uniform wall thinning (non-uniform 

oxide); and overall wall thinning. 

Hydriding occurs as hydrogen is absorbed into the cladding as a result of uniform 

corrosion of the cladding (roughly 15%-20% of the hydrogen generated by the 

corrosion reaction is absorbed into the alloy). This hydrogen precipitates as hydrides 

throughout the cladding thickness at corrosion thicknesses greater than 50 microns 

(ptm). When the overall hydrogen level is high enough (>1000 parts per million 

[ppm]), the cladding is brittle when tested at reactor temperature. It is possible that 

lower levels of hydrogen (600-800 ppm) can affect cladding ductility, especially at 

lower temperature. 

However, lower levels of hydrogen, can also degrade the overall cladding ductility 

depending on the hydride distribution. Hydrogen has high mobility but low solubility in 

Zircaloy is very low, so hydrogen will tend to precipitate out in any cold spot formed 

in the material. For example, there is a much greater hydride concentration near the 

surface of the cladding creating a hydride rim with local hydrogen levels higher 

than 1000 ppm. In addition to being radially localized, the axial distribution of 

hydrogen is also non-homogeneous, with greater concentration in the region in

between the fuel pellets due to the slightly lower heat fluxes and lower temperatures 

at pellet interfaces. 

The main concerns associated with hydriding are: (a) lower ductility and/or 

embrittlement resulting from an overall change in constitutive properties, and (b) 

creation of weak spots in cladding resulting from the formation of a hydride rim, 

and/or hydride blisters. 

Radiation damage. When irradiated to 30 GWd/t (corresponding to a fast fluence of 

~10' n/cm2, E>1 MeV) the cladding suffers an amount of damage calculated at about 

20 dpa (displacements per atom)F-2. The dpa level is roughly proportional to the 

fluence or bumup, so that 60GWd/t corresponds to about 40 dpa and 75 GWd/t to 50 

dpa. This very high level of displacements is translated mostly into radiation

induced dislocation loops, both <a> and <c> type that form from the agglomeration 

of point defects. Although the overall <a> dislocation density saturates after about 

one month of reactor irradiation at a level comparable to that found in cold-worked,

F-3
stress-relieved (CWSR) cladding, the <c> type dislocations evolve over a more 

extended period of time. In addition there are microchemical changes in the alloy related to irradiation induced intermetallic precipitate amorphization and 

dissolution, which can change corrosion resistance and hydrogen pickup. 

The constitutive response of the cladding is also affected by the radiation damage, in 

particular the dislocation loop microstructure formed under irradiation. The yield stress increases, and the uniform strain decreases, i.e. the material undergoes hardening and ductility decrease. The increase in dislocation loop density decreases 

the strain hardening coefficient of the material. At the microscopic level,, these loops can also influence deformation localization at the microscopic level 

(dislocation channeling); the effects of these microscopic processes on macroscopic deformation and failure are not clear at the moment. There is also cladding creepdown, which can cause the gap to be dosed, creating the conditions for fuel

clad chemical bonding to develop. 

The main concerns relating to radiation damage are (a) radiation hardening and 

possible embrittlement, (b) change of corrosion resistance through microchemical 

changes, (c) mechanical property changes and (d) deformation localization (e.g., dislocation channeling, possibly leading to easier axial crack propagation). 

F.2. Fuel Changes 

During normal operation fission gas is formed inside the U0 2 fuel, and distributes 

itself largely into five inventories: (i) gas dissolved in the U0 2 matrix, (ii) gas in intragranular (matrix) bubbles, (iii) gas in intergranular (on grain boundaries) bubbles (iv) gas released to the rod void volume and (v) gas in fuel porosity. The 

amount of gas dissolved in the U0 2 matrix is small, as the solubility of fission gases in U0 2 is low. Contributions (ii) and (iii) result in fuel swelling with consequent pellet-cladding mechanical interaction (PCM) and contribution (iv) is the result of 

fission gas release (FGR), which increases the internal rod pressure and results in 

hoop stress on the cladding. The exact partitioning of these gases among the three 

inventories are dependent on the power history, temperature, fuel microstructure, 

etc. 

Rim Formation. Because of Uranium-238 resonance neutron capture at the U0 2 pellet surface, the amount of plutonium formed in the fuel is greater at the edge of the 

pellet than in the center. This causes the fission rate at the pellet surface to slowly increase with burnup while the fission rate in the bulk of the pellet decreases. The 

ratio of fission at the edge of the pellet to the center may be as high as 3 at high burnups. Such a region is called the rim region and its thickness is approximately 100-300 ttm. The rim region is formed when the local burnup at the rim exceeds 

approximately 60 GWd/t (40-45 GWd/t radial averaged). The rim region has a 

characteristic microstructure that consists of sub-micron size grains with bubbles 

under high gas pressures and has high porosity (20-30%). Some of these bubbles 

may be in non-equilibrium with the matrix because there are large strain fields

F-4
around the smaller bubbles and there is further evidence that they exist within the 

interior of the pellet as well as on the rim if the irradiation temperatures are low. 

The main concerns with the formation of the rim region concern its effects on (a) 

the amount of fission gas loading and (b) the lubrication (by shearing during 

deformation, the rim could reduce the friction coefficient between cladding and 

fuel). 

Fuel restructuring and large cracking. These phenomena occur at low burnups 

when a significant fuel-cladding gap exists. The fuel-cladding gap is either very 

small or non-existent (as evidenced by chemical bonding) in high burnup fuel even 

when the fuel is at hot zero power (reactor coolant is still hot). Therefore, these 

phenomena are not likely to occur in high burnup fuel. 

Microcracking. The mechanical stresses and thermal stresses present in the fuel 

during the RIA transient can cause microcracking to occur at the grain boundaries 

weakened by gas bubbles. The microcracking and its extent can affect both fission 

gas swelling and deformation. 

Pellet-cladding Interface. As burnup increases, a metallurgical or chemical bond 

starts to form between the cladding and the fuel, so that fuel-cladding bonding 

occurs. Clearly the development of this bond depends on the establishment of clad

fuel contact resulting from creepdown and fuel swelling. At intermediate stages, the 

friction coefficient will increase but without perfect bonding. It is important to 

determine the friction coefficient so that we can determine the stress state and 

failure mode of the cladding during pellet-cladding PCMI. 

F.3. References 

F-1. M. Garde, G. P. Smith, and R. C. Pirek, "Effects of Hydride Precipitate 

Localization and Neutron Fluence on the Ductility of Irradiated Zircaloy-4," in 

11th International Symposium on Zr in the Nuclear Industry, vol. STP 1295. 

Garmisch-Partenkirchen: ASTM, 1996, pp. 407. 

F-2. "Standard Practice for Neutron Radiation Damage Simulation by Charged

Particle Irradiation," American Society for Testing and Materials Standard 

Practice E521-96 (1996).

F-5
APPENDIX G 

DESCRIPTION OF BWR FUEL AND CLADDING STATE AT HIGH BURNUPt 

During irradiation, the fuel and cladding experience changes in geometry, material 

macrostructure and microstructure, mechanical properties, and other physical and 

performance characteristics. It is considered that some of these changes could 

possibly affect the fuel rod's ability to maintain its integrity when subjected to an 

accident. Figure G-1 presents a qualitative characterization of some of these fuel and 

cladding changes. These changes, which occur generally gradually over the life of 

the fuel rod, can represent initial conditions for the accident. 

Of the many changes experienced by the fuel and cladding, it is important to discern 

which of these are of greatest importance in determining fuel rod behavior during 

the power oscillations. Some of the more important phenomena are presented and 

discussed below, recognizing that the list may not be inclusive. The changes to the 

fuel and cladding indicated in Figure G-1 are possible, and have been observed, in 

both pressurized water reactor (PWR) and boiling water reactor (BWR) fuel types, 

although to varying extents. Recognizing that the power oscillations are a B WR 

event, the following discussion will attempt to clarify the applicability of the various 

phenomena as currently recognized in modern commercial BWR fuel. 

G.1. Cladding Changes 

The cladding material applied in BWRs is Zircaloy-2, most predominately in the 

annealed, fully recrystallized condition with a zirconium-based inner liner, 

although cold-worked stress relieved material and non-liner applications also exist. 

The zirconium liner can contain varying amounts of alloy additions, intended for 

post-defect corrosion resistance. The primary change mechanisms identified for the 

cladding are waterside corrosion, hydriding, and radiation damage. 

Cladding corrosion occurs through direct exposure of the cladding outer surface to a 

high temperature, highly oxidizing environment enhanced by the radiation field. 

The effects of cladding corrosion are wall thinning, increased heat transfer 

resistance, and cladding hydrogen absorption. In general, the BWR suppliers have 

progressively refined the cladding material processing to minimize the occurrence 

of nodular corrosion, thereby resulting in a generally uniform corrosion 

morphology. Where cladding corrosion distributions are typically peaked at the 

higher elevations in PWRs, the corrosion distributions are generally flatter along 

the fuel rod length in a BWR, with possible peaking at the lower elevations. 

Circumferential variations in cladding oxide layer thickness are observed in BWRs, 

but are generally minor in magnitude. Where cladding corrosion thicknesses of up 

to or greater than 100 ptm has been observed in PWRs, BWR dadding corrosion is 

significantly less, typically less than 50 gim at exposures up to ~62 GWd/MTU peak 

rod average exposure, as observed to date. 

* See Acknowledgments for authorship information.

G-1
Oxide spallation 

"• Extent of spallation 

"* Hydride precipitation 

"* Blister formation 

Waterside corrosion 

"* Corrosion thickness 

"* Hydrogen pickup 

"* Non-uniformities 

in oxide thickness '/.

Loose

Hydrogen in Zircaloy 

0 H pickup fraction 

(hydrogen level in clad) 

0 Hydride orientation 

Hydride radial distribution 

(hydride rim thickness 

Cladding internal 

corrosion 

fuel pieces 

"Large scale pellet cracking from 

normal operation 

"* Width of cracks - missing fuel 

"* Crack spacing

Fuel-cladding chemical bonding 

"* Cladding creepdown & fuel swelling 

"* Compound formation

-Radiation damage to cladding 

"* Formation of dislocation loops 

"* Precipitate evolution 

"* Irradiation creep and growth

Development of rim region 

* Grain size 

* Bubble size distribution 

• Bubble pressure 

* Total gas in rim 

* Pu content & solid fission products

7 Higher hydrogen 

content in between 

Pellet

Constitutive 

properties change 

"* Strain hardening 

"* Strain rate hardening 

* Deformation localization 

(dislocation channel formation)

Fission gas in fuel matrix 

"* Fission gas amount in solution 

"* Intergranular bubbles 

"* Intragranular bubbles 

"* Conductivity degradation 

"* Fission gas released during 

normal operation 

"* Porosity gas content

Fig. G-1. Fuel and Cladding State 

An important consideration is oxide layer cracking, delamination, and spalling. Oxide layer cracking and delamination can lead to an acceleration in the oxide layer growth rate. Spalled oxide regions result in a cooler cladding metal temperature during operation than exists under the adjacent unspalled oxide regions. The

G-2
presence of such "cold spots" can promote redistribution of any hydrogen absorbed 

from the cladding outer surface corrosion process, thereby leading to hydride localizations and even bulk hydride formation (observable as bulges or blisters) in 

the outer region of the cladding. Such bulk hydride formation regions are highly 

embrittled and are often accompanied with partial cladding cracks even in the 

absence of applied loading by the fuel pellets (caused simply by the volume 

expansion associated with the conversion of zirconium to zirconium hydride). 

Where significant accelerated corrosion and oxide layer spalling has been observed 

in PWRs, similar conditions are typically not observed in BWRs with modern 

cladding materials. 

Corrosion localizations have been observed at fuel assembly spacer locations, 

adjacent to Inconel components (typically referred to as "shadow corrosion"). 

Although accelerated localized corrosion, leading to fuel rod failure, has occurred at 

one BWR with an earlier cladding material type, in general, the available 

characterizations indicate that this localization develops relatively quickly, but then 

remains relatively stable, at least to exposure levels characterized to date (-62 

GWd/MTU peak rod average exposure). 

BWRs operate with several water chemistry options: Hydrogen Water Chemistry, 

Zinc Injection, and Nobel Metal Chemical Addition. To date, no unacceptable 

changes in the cladding corrosion performance have been observed under these 

water chemistry options. 

In summary, in BWRs with modern cladding, the primary effects of interest from 

the corrosion process are (1) wall thinning, (2) increased heat transfer resistance, and 

(3) the effects of corresponding hydrogen pickup. 

Hydriding occurs as hydrogen, liberated by the dadding outer surface corrosion 

process, is absorbed into the cladding. Typically, less than 20% of the hydrogen generated by the corrosion reaction is absorbed by the cladding. This absorbed 

hydrogen generally precipitates as circumferentially oriented zirconium hydride 

stringers when the amount of absorbed hydrogen exceeds the solubility level. 

Available testing has demonstrated no adverse influence of hydrogen on elevated 

temperature irradiated Zircaloy ductility (total elongation) for hydrogen contents up 

to at least 850 ppm(c"-). At higher hydrogen levels, something in excess of 1000 ppm, the cladding ductility can be reduced at operating temperatures. Most typically, 

BWR cladding hydrogen content is <200 ppm, as characterized at -50 GWd/MTU 

rod average exposure for modem BWR cladding materials. Although higher levels 

(less than 600 ppm) have been observed in older cladding types at elevated 

exposures (up to -65 GWd/MTU rod average exposure), even this level is below 

that required to significantly affect the cladding mechanical properties. 

At high hydrogen levels (in excess of 1000 ppm) a dense hydride rim can form near 

the cladding outer surface, primarily as observed in PWR fuel applications. Hydride localizations have also been observed at "cold spots" occurring at pellet-pellet 

interfaces (adjacent to pellet chamfers), and more significantly, at spalled oxide

G-3
locations as discussed previously. With the generally lower hydrogen concentration 

observed in BWR fuel, dense hydride rims or extreme localizations at pellet-pellet interfaces have not typically been observed, although the general tendency of 

hydride accumulations toward the cladding outer surface or near pellet-pellet 

interfaces has been observed. 

Another consideration, although not typically observed in either PWR or B W R 

applications is the development of radially oriented hydrides, which, in significant 

concentration, could affect the cladding ductility. 

In summary, in BWRs with modern cladding, the primary considerations with 

cladding hydrogen content are (1) the impact, if any, on the cladding mechanical 

properties, and (2) the effect of hydride localizations to form weak, damage susceptible regions. In general, these considerations have not been found to be 

significant for the hydrogen contents observed in modern BWR cladding to date. 

Radiation Damage to the cladding material occurs as a direct consequence of 

exposure to fast neutrons. This radiation damage is manifested as radiation-induced 

dislocation loops, both <a> and <c> type that form from the agglomeration of point defects. Although the overall <a> dislocation density saturates very early in life, the 

<c> type dislocations evolve over a more extended period of time. The effect of this 

damage is a strengthening of the material, with a corresponding reduction in 

ductility, and increased irradiation-induced stress-free growth (occurs in the absence 

of an applied stress). Additionally, microchemical changes occur as the irradiation 

induces intermetallic precipitate amorphization and dissolution, which can alter 

the mechanical properties, corrosion resistance and possibly also the hydrogen 

pickup of the cladding material. 

In addition to irradiation-induced growth of the cladding material, irradiation also 

induces cladding creep in response to the applied fuel rod internal-external pressure 

difference and pellet expansion loadings. 

In summary, the primary considerations relative to cladding radiation damage are 

(1) radiation hardening and the corresponding mechanical properties impact, and (2) 

deformation caused by irradiation-induced growth and creep. 

G.2- Fuel Changes 

Fission Products. During normal operation, solid and gaseous fission products are 

generated within the U0 2 fuel pellet. Whereas the solid fission products generally remain at the birthsite, the gaseous fission products are more mobile and distribute 

largely into five separate inventories: (i) gas dissolved in the U0 2 matrix, (ii) gas in 

intragranular (matrix) bubbles, (iii) gas in intergranular (on grain boundaries) 

bubbles (iv) gas released to the fuel rod void volume and (v) gas in fuel porosity. The amount of gas dissolved in the U0 2 matrix is limited by the solubility in U0 2. Solid fission products result in a progressive swelling of the fuel material with 

irradiation exposure. Gaseous fission product inventories (iii), and to a lesser extent

G-4
(ii) and (v), under high temperature low restraint conditions, can also result in fuel 

swelling with consequent pellet-cladding contact. Inventory (iv) is referred to as 

fission gas release (FGR) and produces an increase in the fuel rod internal pressure and corresponding cladding loading. The exact partitioning of the fission gases 

among the identified inventories is dependent primarily on the fuel pellet 

microstructure and thermal operating history. 

Rim Formation. As a result of Uranium-238 resonance neutron capture at the U0 2 

pellet periphery, the amount of plutonium formed in the fuel pellet is greater at the 

pellet periphery than in the center. This plutonium buildup causes a significant increase in the fission rate at the pellet periphery, relative to the fission rate in the 

bulk of the pellet. At elevated exposures, the result of this elevated fission rate is to 

produce a highly porous, fine grained structure. This altered structure region is 

called the rim region. The size of the rim region increases relatively progressively 

with increased exposure above -40-45 GWd/MTU pellet average exposure. The 

primary considerations with the formation of the rim region are (1) possible increased fission gas release, (2) possible increased resistance to heat transfer, and (3) 

possible increased gaseous swelling under high rim temperature conditions. It is 

noted that the pellet rim may provide a cushion, or lubricating, effect that may 

reduce the consequences of pellet-cladding mechanical interaction. 

Fuel restructuring and macrocracking. During the initial rise to power, the thermal 

stresses caused by the pellet radial temperature gradient cause the pellet to crack 

(primarily radially). With the release of strain energy, the cracked pellet segments 

relocate outwards toward the cladding (called fuel relocation or restructuring). With 

continued irradiation, additional outward movement of the pellet segments can 

occur. At -mid-life exposures, the combined effects of pellet relocation, fuel 

irradiation swelling, and cladding creepdown result in a closed pellet-cladding gap. 

From this point, (1) a reduction in the fuel pellet expansion (such as caused by a 

power decrease) can result in partial gap opening, and (2) additional fuel expansion (by progressive fuel swelling or as a result of a power increase) can cause pellet radial 

cracks to (partially) close, thereby increasing the effective pellet stiffness, and 

imposing loading and deformation of the cladding. No particular change in this 

behavior is expected at elevated exposures. 

Microcracking. During a rapid reactivity pulse where the pellet rim can experience 

significant heatup, and in the absence of significant constraint provided by the 

cladding, gas bubble expansion at the grain boundaries (most notably at the pellet rim) could lead to grain boundary cracking (decohesion). The result would be a 

release of fission gases to the fuel rod void volume with an increase in the fuel rod 

internal pressure and applied cladding pressure loading, with a subsequent 

reduction in the local pellet expansion. In the presence of significant cladding 

constraint, gas bubble expansion would be suppressed with a corresponding reactive 

increased loading of the cladding, likely with no significant fission gas release until 

release of the applied hydrostatic stress such as would occur on cooling. Additional 

pellet cracking can also occur on cooling, resulting in additional fission gas release,

G-5
but correspondingly also reducing the gaseous swelling potential for the next heatup 

cycle. 

Pellet-Cladding Interface. With the onset of pellet-cladding contact, a bond layer develops between the fuel pellet and the cladding. At elevated exposure, the 

magnitude (bond layer thickness) and extent (circumferential and axial surface 

coverage) increases. The development of this bond layer affects the ability of the 

pellet and cladding to move independently (effective friction), and thereby affects 

load transfer from the pellet to the cladding and the subsequent cladding stress state. 

The bond layer can fracture during cooldown or power reductions, leading to an 

intermediate state. 

G.3. References 

G-1. S. B. Wisner, R. B. Adamson, "Combined Effects of Radiation Damage and 

Hydrides on the Ductility of Zircaloy-2", Nuclear Engineering and Design, 185 

(1998), pages 33-49.

G-6
APPENDIX H

MEMBERS OF THE HIGH BURNUP FUEL PIRT PANEL 

Carl A. Alexander 

Carl Alexander is Chief Scientist of Battelle's government sectors operation. He has 

a B.S. in Mathematics from Ohio University, a M.S. in Physics from the same 

institution, and a Ph.D. in Ceramic Engineering received in 1961 from The Ohio 

State University. From 1962 to 1985 he was a member of the engineering and 

graduate faculty of The Ohio State University, with joint appointments as Adjunct 

Professor of Nuclear Engineering as well as Ceramics and Materials Engineering. He 

has also served as Adjunct Professor at the University of Maryland and 

Southampton University in the U.K. His specialty is nuclear fuels and 

thermodynamics. He performed some of the first loss-of-coolant simulations in the 

late 1950s early 1960s. He contributed to Wash-1400 in which he showed the 

importance of cesium iodide as a transport medium in a LOCA. He performed 

several studies of fission'product release with real fuels at very high temperatures 

and has evaluated a number of complexes involving urania and Zircalloy at very 

high temperatures. 

Jens G. M. Andersen 

Jens G. Munthe Andersen is a principal engineer at Global Nuclear Fuel. He has a 

M.S. in Nuclear Engineering for the Technical University in Denmark and obtained 

a Ph.D. in Nuclear Engineering from the same institution in 1974. From 1971 to 1978 

he was employed by Riso National Laboratory in Denmark. From 1978 Dr. Andersen 

has been employed by General Electric Nuclear Energy and since January 2000 by 

Global Nuclear Fuel (a joint venture of GE, Toshiba and Hitachi). He is currently 

leader of the Methods and Process Development team at Global Nuclear Fuel. Dr. 

Andersen has 29 years experience in the nuclear field. He has been primarily 

engaged in developing computer programs for boiling water reactor transient and 

safety analysis. He has participated in numerous PIRT panels and the application of 

the CSAU methodology to BWR. 

John A. Blaisdell 

John Blaisdell is a Senior Consulting Engineer at Westinghouse Electric Company 

(CE Nuclear Power, LLC). He received his BS degree in Mechanical Engineering 

from Clarkson University in 1961 and a Ph.D. in Mechanical and Aerospace 

Engineering from North Carolina State University in 1969. Dr. Blaisdell has worked 

in the nuclear industry for that past 29 years. His experience includes work on the 

PWR FLECHT-SEASET experimental program, including test specification 

development and review and correlation of results; development of best-estimate 

small break LOCA analytical methods; and supervising the development of 

mathematical models of fuel behavior during a LOCA. He was the Manager of 

Safety Analysis at Northeast Utilities where he managed the development of plantH-1
specific probabilistic safety analyses (PSAs) for four nuclear units and the in-house 

development of transient and LOCA analysis capability. He was also a project 

manager at Yankee Atomic Electric Company where he managed the development 

of a best-estimate containment analysis for the Maine Yankee power plant. This 

work included facilitating a PIRT panel and managing the development of the 

methodology to statistically combine the results of the mass and energy calculations 

with the containment response calculations. Dr. Blaisdell is currently involved in 

the LOCA and non-LOCA safety analyses for both PWRs and BWRs. 

Brent E. Boyack 

Brent E. Boyack is the facilitator for the High Burnup Fuel PIRT Panel. He is a 

registered professional engineer. He obtained his B. S. and M. S. in Mechanical 

Engineering from Brigham Young University. He obtained his Ph.D. in Mechanical 

Engineering from Arizona State University in 1969. Dr. Boyack has been on the staff 

of the Los Alamos National Laboratory for 19 years; he is currently the leader of the 

software development team, continuing the development, validation, and 

application of the Transient Reactor Analysis Code (TRAC). Dr. Boyack has over 30 

years experience in the nuclear field. He has been extensively engaged in accident 

analysis efforts, including design basis and severe accident analyses of light water, 

gas-cooled, and heavy-water reactors; reactor safety code assessments and 

applications; safety assessments; preparation of safety analysis reports; and 

independent safety reviews. He chaired the MELCOR and CONTAIN independent 

peer reviews and was a member of the Code Scaling, Applicability and Uncertainty 

or CSAU technical program group. He has participated in numerous PIRT panels. 

He has over 70 journal and conference publications and is an active member of the 

American Nuclear Society. 

Bert M. Dunn 

Bert M. Dunn obtained his B. S. in Physics from Washington State University in 

1968 and his M. S. in Physics from Lynchburg College in 1973. Mr. Dunn has worked 

in LOCA and Safety Analysis for the Babcock and Wilcox Company (B&W) and 

Framatome Technologies (FTI) for 28 years. Mr. Dunn has served as the lead 

technically for the development of the B&W and FTI LOCA evaluation models for 

once through and recirculating steam generator plants. He has worked with both 

deterministic and best estimate LOCA evaluation techniques. He has also been 

technical lead for method development and application of boron dilution accident 

methods and pressurized thermal shock evaluation methods. He is currently 

employed as an Advisory Engineer with responsibility for. the development of 

LOCA and Safety Analysis techniques for evaluation of advanced cladding 

materials. This includes test specification development, review and correlation of 

results, and the incorporation of results into requisite analytical methods. Mr. 

Dunn has been primary author on several company topical reports covering both 

methods development and accident analysis.

H-2
Derek B. Ebeling-Koning

Derek B. Ebeling-Koning is Manager of BWR Fuel and Engineering Analysis at 

Westinghouse Electric Company. He received his B.S. degree in Nuclear 

Engineering from Rensselaer Polytechnic Institute in 1977. He received his M.S. and 

Ph.D.degrees in Nuclear Engineering in 1979 and 1983, from Massachusetts Institute 

of Technology. He has worked for the last 17 years in BWR and PWR safety 

analysis, initially for Westinghouse Electric, then ABB Combustion Engineering 

starting in 1991 (now a subsidy of Westinghouse Electric.) His expertise includes 

methods and modeling of two-phase flow; BWR LOCA analysis, PWR operational 

transients and containment analysis; and a lead role in developing and licensing of 

ABB BWR reload analysis methodology in the U.S. Dr. Ebeling-Koning is a member 

of the ANS and ASME, and technical reviewer for several industry journals. 

Toyoshi Fuketa 

Toyoshi Fuketa is a Principal Engineer in the Fuel Safety Research Laboratory at the 

Japan Atomic Energy Research Institute (JAERI). He obtained his B. S., M. S. and 

Ph.D. in Mechanical Engineering Science from Tokyo Institute of Technology, Japan, 

in 1982, 1984 and 1987, respectively. Dr. Fuketa has been involved in the Nuclear 

Safety Research Reactor (NSRR) project to study behavior of LWR and research 

reactor fuels under reactivity accident and severe accident conditions and to 

evaluate the thresholds, modes, and consequences of fuel failure in terms of the fuel 

enthalpy, fuel burnup, coolant conditions, and fuel design. His research interests 

include fuel-coolant interactions, fuel failure mechanisms and transient fission gas 

behavior. He was engaged in small-scale steam explosion experiments at Sandia 

National Laboratories, Albuquerque, from 1988 to 1990, as a visiting scientist. 

Georges Hache 

Georges HACHE is a PIRT expert on LOCA fuel behavior from the French Nuclear 

Safety and Protection Institute (IPSN). He graduated from the French Ecole 

Polytechnique in 1973 and Ecole des Mines de Paris in 1976 (colleges of university 

level, admission to which is by strict examination). After ten years in the French 

nuclear inspectorate, he joined the IPSN Safety Research Division at Cadarache in 

1986. He was involved in the development and assessment of safety computer 

codes, describing fuel and fission products behavior during severe and design basis 

accidents of light water and fast breeder reactors (SCANAIR, ICARE2, ESCADRE...), 

including definition and interpretation of safety tests in the CABRI and PHEBUS 

reactors. After having led the safety software development and assessment team for 

a period of seven years, he is now senior scientific adviser to the head of the 

Division. He has been engaged in international cooperative efforts including the 

OECD / CSNI Rasplav program and a task force on fuel safety criteria.

H-3
Lawrence E. Hochreiter

L.E. (Larry) Hochreiter is a professor of Nuclear and Mechanical Engineering at the 

Pennsylvania State University and does research and teaching in the areas of two

phase flow and heat transfer, reactor thermal-hydraulics, fuel rod design, and nuclear reactor safety. He recieved a BS degree in Mechanical Engineering from the 

Universtiy of Buffalo and a MS and Ph.D degrees in Nuclear Engineering from Purdue University. While at Pennsylivania State University, Dr. Hochreiter has 

developed a detailed reflood heat transfer PIRT to guide the design and instrumentation of the NRC Rod Bundle Heat Transfer program, located at Penn 

State. Before joining the Penn State University in 1997, Dr. Hochreiter was a 

Consulting Engineer at the Westinghouse Electric Corporation for nearly 26 years and was responsible for the development, testing validation, and licensing of Westinghouse safety analysis methods. He developed the large-break Loss Of 

Coolant Accident (LOCA) PIRT for the Westinghouse Best-Estimate Methodology. He also participated in and helped develop the Westinghouse small-break LOCA 

PIRT. Dr. Hochreiter also developed several PIRTs for the Westinghouse advanced 

AP600 design for the accident analysis methods and presented these PIRTs to the 

NRC and the ACRS. 

S. E. "Gene" Jensen 

S. E. "Gene" Jensen is an experienced engineer with 38 years of involvement in 

nuclear safety applications and development. Mr. Jensen obtained his BS degree in 1961 from Montana State and his MS degree from the University of Idaho both in 

Chemical Engineering. He is a registered professional engineer. He spent the first 14 years of his career at the Idaho National Engineering Laboratory working in various phases of the nuclear safety test engineering program. He became involved 

with the early regulatory support activities including the development of the Water 

Reactor Evaluation Model (WREM) to perform LOCA analyses according to the NRC ECCS Rule which came out in 1975. In 1975, Mr. Jensen joined Exxon Nuclear 

Company (since purchased by Siemens), where he was instrumental in developing LOCA ECCS Evaluation Models for both PWRs and BWRs and obtaining NRC 

approval of these models. Mr. Jensen has spent 24 years with what is now Siemens 

Power Corporation all of which were involved with safety analysis applications and 

development including both transient and LOCA analysis. His most recent involvement has been as the technical lead in developing a PWR Realistic LOCA 

methodology for Large Break LOCA. Mr. Jensen is a member of both ANS and 

AIChE. 

Siegfried Langenbuch 

Siegfried Langenbuch is head of the reactor dynamics division of GRS. He obtained 

his Diploma in Physics from the University of Munich in 1969. The objective of his 

Dr. degree work was the development of an efficient spatial- and time-dependent 3D-neutronics model for studying reactivity initiated accidents. His research interests were code development for neutron dynamics and thermo-fluid dynamics

H-4
of the reactor core, including the coupling of 3D-neutronics models with plant 

system codes. In addition, he has experience in safety review of nuclear design, 

thermal design, and accident analysis of BWRs and PWRs as well as of VVERs and 

RBMKs of Russian design. He is a member of national and international working 

groups for the requirements of nuclear design. He has numerous publications in 

the field of reactor core dynamics. 

Frederick 1. Moody 

Frederick J. Moody is a Consulting Engineer in Thermal-Hydraulics, who has 

participated in numerous NRC - sponsored peer review groups and Technical 

Program Groups, involving the analysis of postulated nuclear reactor accidents. He 

received his Ph.D. in Mechanical Engineering from Stanford University in 1971. He 

completed 41 years of reactor and containment safety analyses at the General Electric 

Nuclear Energy Division, where he developed various industry-standard analytical 

models for studies involving pipe and component rupture blowdown of high 

pressure steam and water mixtures, containment pressure and jet impingement 

loads, waterhammer forces associated with pipe flow accelerations, dynamic and 

thermal response of nuclear reactor core components during accident conditions, 

and fluid-structure interaction of submerged structures. He has taught numerous 

engineering courses as an adjunct professor for 28 years at San Jose State University, 

as an in-plant instructor at General Electric, and more recently as an instructor for 

professional development courses sponsored by the American Society of Mechanical 

Engineers. He has authored numerous journal papers, written an engineering 

textbook, Introduction to Unsteady Thermo-Fluid Mechanics (Wiley Interscience, 

1990), and co-authored The Thermal-Hydraulics of a Boiling Water Nuclear Reactor, 

2nd Ed., ANS Press, 1993. 

Arthur T. Motta 

Arthur T. Motta has worked in the area of radiation damage to materials with 

specific emphasis in Zr alloys for the last fifteen years. He received a B.Sc. in 

Mechanical Engineering and an M.Sc. in Nuclear Engineering from the Federal 

University of Rio de Janeiro, Brazil, and a Ph.D. in Nuclear Engineering from the 

University of California, Berkeley. He worked as a research associate for the CEA at 

the Centre for Nuclear Studies in Grenoble, France for two years and as a post

doctoral fellow for AECL at Chalk River Laboratories, Canada, before joining Penn 

State in 1992. The research programs he developed at Penn State include mechanical 

behavior of Zr alloys, advanced techniques for characterization of Zr alloys, and its 

oxides, defects in intermetallic compounds and phase transformation under 

irradiation. He has expertise in transmission electron microscopy, charged particle 

irradiation, mechanical testing, positron annihilation spectroscopy and theoretical 

expertise on phase transformations under irradiation and microstructural evolution 

under irradiation. He has recently authored review articles on amorphization under 

irradiation and on zirconium alloys in the nuclear industry. He was recently guest 

editor for a special issue of the Journal of Nuclear Materials, and was a member of a

H-5
DOE panel to evaluate research needs on radiation effects on ceramics for 

radioactive waste disposal. 

Mitchell E. Nissley 

Mitchell E. Nissley obtained his B. S. and M. Eng. degrees in Nuclear Engineering from Rensselaer Polytechnic Institute. Mr. Nissley has been on the staff of the 

Westinghouse Electric Company for 18 years; he is currently the leader of the team 

responsible for the development, licensing and application of the various realistic 

large break LOCA analysis codes and methodologies employed by Westinghouse. His contributions to the nuclear industry include the development and licensing of 

critical heat flux correlations for advanced PWR and VVER fuel designs, and the 

development and licensing of realistic large break LOCA evaluation models for 

Westinghouse PWR designs (cold leg injection, upper plenum injection and 

AP600). He has numerous journal and conference publications. 

Katsuhiro Ohkawa 

Katsuhiro Ohkawa is an advanced technical engineer at Westinghouse Electric 

Company. He received his BS in Physics from Sophia University in Tokyo, MS 

(1978) and PhD (1983) in Nuclear Science and Engineering from Rensselaer 

Polytechnic Institute. His experience at Westinghouse includes the development of 

Liquid Metal Advanced Nuclear Plants, BWR and PWR safety methods. Since 1990, he has been involved in the development of CSAU based Best Estimate Large Break. 

Currently he is involved in the development of Best Estimate Small Break LOCA 

Methodology. 

Kenneth L. Peddicord 

Kenneth L. Peddicord is Associate Vice Chancellor and Professor of Nuclear 

Engineering at Texas A&M University. He received his B.S. degree in Mechanical 

Engineering from the University of Notre Dame in 1965. He obtained his M.S. 

degree in 1967 and his Ph.D. degree in 1972, both in Nuclear Engineering from the 

University of Illinois at Urbana-Champaign. From 1972 to 1975, Dr. Peddicord was a 

Research Nuclear Engineer at the Swiss Federal Institute for Reactor Research (now 

the Paul Scherrer Institute) where he worked in the plutonium fuels program. From 1975 to 1981, Dr. Peddicord was Assistant and Associate Professor in the 

Department of Nuclear Engineering at Oregon State University. From 1981 to 1982, 

he was a Visiting Scientist at the EURATOM Joint Research Centre in Ispra, Italy where he was involved in the Super Sara Severe Fuel Failure. Programme. In 1983, 

Dr. Peddicord joined Texas A&M University as Professor of Nuclear Engineering. He has served as Head of the Department of Nuclear Engineering (1985-88), Associate Dean for Research (1988-91), Interim Dean of Engineering (1991-93), and 

Director of the Texas Engineering Experiment Station (1991-93). Since 1994, he has 

been Associate Vice Chancellor of the Texas A&M University System. Dr. Peddicord 

serves as the representative of the A&M System to the Governing Board of the 

Amarillo National Resource Center for Plutonium. Dr. Peddicord's research

H-6
interests are in the performance and modeling of advanced nuclear fuels. Since 

1995, he has been a participant in joint DOE-Minatom activities on excess 

plutonium disposition and nuclear materials safety. Dr. Peddicord has 120 

publications in technical journals and conferences. He is a registered professional 

engineer in the state of Texas and has been a member of the American Nuclear 

Society since 1975. 

Gerald Potts 

Mr. Potts of Global Nuclear Fuel received a Bachelor of Science degree in 

Mechanical Engineering from the University of California, and a Master of Science 

degree in Mechanical Engineering from Santa Clara University. Mr. Potts has 

accumulated 28 years experience in the commercial nuclear power industry within 

the General Electric Nuclear Energy division. Mr. Potts' responsibilities and 

experience include fuel rod thermal-mechanical design, fuel rod thermal

mechanical performance and licensing basis analytical model development, and 

fuel integrity assessment under normal steady-state operation, anticipated 

operational transient, and accident conditions. 

Joe Rashid 

Joe Rashid is a Fellow of the ASME and a registered Nuclear Engineer. His general 

field of expertise is computational thermo-mechanics, structural mechanics and 

material constitutive modeling. He acquired his graduate and undergraduate 

education in mechanics at the University of California Berkeley, receiving the PhD 

degree in 1965. Having received his education at the birth place of the Finite 

Element Method in the early sixties, Dr. Rashid was among the pioneering 

contributors to its development, in particular three-dimensional computations. Dr. 

Rashid's three and a half decades career in the nuclear industry began with the gas

cooled reactor technology at General Atomics in San Diego, followed by an eight

year career in BWR technology at General Electric in San Jose, and finally at 

ANATECH Corp. which he founded in 1978. At General Atomics, his work in the 

mechanics of concrete reactor vessels and nuclear fuel particles led to the 

development of the smeared-crack model, which was adopted in finite element 

codes as the basic model for the cracking analysis of brittle materials. At GE, he was 

responsible for the development of the industry's first two-dimensional fuel rod 

behavior code for the analysis of the then-emerging pellet-clad interaction (PCI) 

problem. At ANATECH, Dr. Rashid undertook the development of the transient 

fuel analysis code FREY for the Electric Power Research Institute (EPRI). In the 

aftermath of the Three Mile Island accident, EPRI's collaboration with Sandia in 

reactor containment research, with Dr. Rashid as the principal investigator for EPRI, 

led to the institutionalization of the leak-before-break concept for reactor 

containment structures, thereby profoundly affecting risk assessment of loss of 

coolant accidents. He participated in severe accident work with Sandia and EPRI, 

which included the development of constitutive models and analysis methods for 

the creep rupture of pressure vessel lower head under loss of coolant accident. He 

participated in the expert review process for NUREG-1150, and was nominated by

H-7
NRC to chair an international expert panel for OECD'sVessel Investigation Project. Dr. Rashid's publications in the various fields of activity in which he had primary contributions exceed 100, which include journal articles, reports and white papers. 

Richard 1. Rohrer 

Richard J. Rohrer serves as a member of the High Bum-up Fuel Phenomena 

Identification and Ranking Table (PIRT) Panel. He is a registered professional engineer in the state of Minnesota. He obtained a B.S. in Nuclear Engineering from 

the University of Illinois, and an M.S. in Nuclear Engineering from the University of Wisconsin in 1983. He also holds an M.S. in Management from Cardinal Stritch 

College, and a Senior Reactor Operator Certification for the Monticello Nuclear 

Generating Plant. Mr. Rohrer has over 16 years experience supporting operations of 

nuclear power reactors, including licensing, reactor engineering, probabilistic safety 

assessment, core design, accident analysis, and transient analysis. He currently 

manages projects for the Monticello Nuclear Generating Plant in the Nuclear 

Analysis and Design group with Nuclear Management Company. Mr. Rohrer is a 

member of the American Nuclear Society and has published five technical papers 

on probabilistic safety assessment and Boiling Water Reactor stability. In addition, 

he is an active participant in the Electric Power Research Institute's Robust Fuel 

Program. 

James S. Tulenko 

James S. Tulenko is Chairman of the Nuclear and Radiological Engineering Department and a Professor of Nuclear Engineering at the University of Florida. He 

received his B.A. with honors in Engineering Physics from Harvard College and his 

M.A. in Engineering Physics from Harvard University in 1960. After military 

service in the Corps of Engineering, he obtained a M.S. in Nuclear Engineering from 

the Massachusetts Institute of Technology in 1963. In 1980 he obtained a M.E.A. 

from George Washington University. Professor Tulenko's professional activities 

have included all aspects of the nudear fuel cycle. He has over 35 years of 

experience in fuel design, fuel operation and fuel performance. Professor Tulenko 

was Manager of Nuclear Development at United Nuclear Corporation where he 

patented the water hole thermalization concept now utilized in all boiling water 

reactors. He also was project engineer for one of the first Plutonium reloads in a 

commercial reactor. He served as Manager of Physics for Nuclear Materials and 

Equipment (NUMEC) Corporation where he headed up nuclear physics activities. 

He later served as Manager of Physics and Manager of Nuclear Fuel Engineering for 

the Nuclear Power Division of Babcock and Wilcox. In 1979 he was made a Fellow 

of the American Nuclear Society (ANS) for his contributions to the fuel cycle. In 

1980 he received the Silver Anniversary Exceptional Service Award of the ANS for 

his outstanding contributions to the Nuclear Fuel Cycle in the first 25 year of the 

ANS. In 1997 he received the Mishima Award of the ANS given for outstanding 

contributions to Nuclear Material Research. He also was awarded the Glenn 

Murphy Award of the American Society of Engineering Education given to the 

Outstanding Nuclear Engineering Educator. He is a Board Member of the National

H-8
Nuclear Accrediting Board of the Institute of Nuclear Power Operations and a Board 

Member of the American Nuclear Society. He is also a Commissioner of the 

Engineering Accreditation Commission. He has over 100 journal and conference 

publications and has consulted for a variety of government agencies and 

commercial companies. 

Keijo Valtonen 

Keijo Valtonen is a Chief Inspector with the Radiation and Nuclear Safety 

Authority of Finland. He obtained his degree from the University of Helsinki 

where he majored in reactor physics and thermal hydraulics. His primary duties 

since 1975 have been fuel, nuclear and thermal-hydraulic design of reactor cores; 

transient and accident analysis for Loviisa (VVER-440 type PWR) and Olkiluoto 

(ABB-Atom type BWR); and operator qualification, including oral licensing 

examinations and review of operator instructions. He has reviewed plant feasibility 

studies, including those for the VVER-1000, ABB-Atom BWR 90, Siemens PWR, 

and SECURE and PIUS. He has reviewed numerous feasibility studies for new fuel 

designs, including VVER Zr 1% Nb, BNFL-VVEF fuel, ABB 8x8, SVEA 64, SVEA 

100, Siemens 9x9, GE12 and Siemens ATRIUM 10. He has participated in safety 

reviews for the RBMK. He has engaged in research work on the transient behavior 

of BWR and PWR reactor cores, BWR stability analysis, validation of TRACB and 

RAMONA computer codes, PWR boron dilution, and several fuel transient 

behavior studies for VVER and BWR reactors. He has been engaged in 

international cooperative efforts including IAEA and OECD development of safety 

criteria for future nuclear reactors, regulatory approaches to severe accident issues 

for the OECD/CNRA, a state-of-the-art report on BWR stability, the European 

Union's safety RBMK safety review, the OECD/CSNI task force of fuel safety criteria. 

Nicolas Waeckel 

Nicolas Waeckel is a visiting senior engineer from EDF to EPRI. During his stay at 

EPRI he managed the Working Group 2 (response in transient) within the Robust 

Fuel Program. Nicolas Waeckel was the Technical Leader and Manager of the 

Nuclear Fuel design and Survey Group at Electricit6 de France (EDF) Septen. At EDF 

he developed and managed fuel R&D programs including fuel R&D programs 

addressing RIA and LOCA issues. He developed and managed activities in areas of 

nuclear fuel rod and nuclear fuel assembly design, design methodologies and fuel 

rod performance code development (normal operation conditions and accidental 

conditions). He represented EDF in interacting with the French Safety Authorities 

on many key issues (RIA, LOCA, fuel assembly distortion, bum-up extension, etc...). 

From 1984 to 1990, he was in charge of the FBR and the LWR Structural Mechanics 

Design Group thereafter. He has developed several design methodologies (buckling 

of thin shells, creep-fatigue and progressive deformations) and participated to the 

writing of the RCCMR (design rules for the FBRs). He managed 15 PhD students and 

20 contracts with Universities, CEA and Novatome. He has authored papers and 

reports in areas of mechanical design of thin structures (buckling, creep-fatigue, 

ratchetting and fracture mechanics) and nuclear fuel design and performance (PCMI,

H-9
High burn-up properties, RIA and LOCA). He obtained an MS in Civil Engineering from the National Institute of Applied Sciences in Lyon (France) in 1978, a PhD in 

1981 and a Sciences Doctorate in 1983 from the same Institute. The topic of his 

Sciences Doctorate Thesis was the impact of initial geometrical defect on buckling of 

FBR related thin structures. 

Wolfgang Wiesenack 

Wolfgang Wiesenack is the acting general manager of the OECD Halden Reactor 

Project. He obtained an MS in nuclear engineering from the University of Hanover, 

Germany, in 1976 and a PhD in nuclear engineering and LWR fuel behavior 

modeling from the same university in 1983. Dr. Wiesenack had a research assistant 

position at the University of Hanover, working on LOCA analysis (RELAP 4) and 

modeling of LWR fuel behavior in normal operating conditions. He joined the 

OECD Halden Reactor Project in 1984. As senior reactor physicist he was responsible for the core physics calculations of the Halden reactor, including nuclear design studies of experimental rigs, core loadings and updating of the reactor's safety report. He was also responsible for the data acquisition of the reactor and implemented a 

completely renewed system. As the head of the Data Acquisition & Evaluation 

division, he was in direct contact with many aspects of fuels and materials 

behaviour under steady state and ramping and transient conditions. He was 

actively engaged in the execution of the IAEA code comparison exercise FUMEX to 

which the Halden Project provided the data. He was also a member of the 

FRAPCON peer review team. He is a member of the German nuclear society.

H-10
APPENDIX I

TUTORIAL PRESENTATIONS TO THE PIRT PANEL 

This appendix contains information presented to the PIRT panel with the objective of 

assisting the panel members to develop a common understanding of LOCA phenomena 

and fuel behavior in PWRs and BWRs. 

I-1 PWR Large-Break LOCA: Impact of High Burnup Fuel 

Panel members Lawrence E. Hochreiter, Mitchell E. Nissley, and Bert Dunn prepared 

this review for the PIRT panel. 

1-1.1. Introduction 

The Large-Break Loss of Coolant Accident (LBLOCA) will be examined for two 

different types of Pressurized Water Reactors (PWRs). First, the typical response of a 

Westinghouse plant, using the approved Appendix K Evaluation Model (EM) will be 

examined, then the plant response using the approved Westinghouse best estimate 

calculation will be examined. The response of a Babcock and Wilcox lowered loop PWR 

will also be examined. In addition to the system response for these plant designs, the 

fuel temperature and cladding behavior will be examined. 

A LOCA is defined as a failure of the main reactor coolant piping which is sufficiently 

large enough to cause a safety injection signal in which the Emergency Core Cooling 

system (ECCS) is initiated. This is usually a break which is larger than 3/8 of an inch. 

Breaks smaller than this size are classified as leaks since the normal plant charging 

systems have sufficient make-up flow to prevent draining of the reactor coolant system. 

There are two classifications of breaks; Small Breaks are breaks that are less than one 

square foot in flow area, and large breaks are breaks that are larger than one square foot 

in area up to and including a full size double-ended guillotine rupture of the main 

reactor piping. A Double-Ended cold Leg Guillotine break (DECLG) can have a flow 

area of approximately nine square feet. 

The LBLOCA has been used as the design basis accident for the light water reactors, 

and it's really a "battleship in the desert" type of approach in which the vendor is asked 

to analyze this accident without regard to the probability of the accident occurring. The 

real purpose of analyzing the accident is to verify that the design and function of the 

emergency core cooling systems, which are provided for the safety of the plant, which 

recover the plant and limit the fuel damage to acceptable levels. The LBLOCA is used as 

a calculation exercise, basically to verify the performance of the emergency core cooling 

system. 

In 1965 and earlier, the LOCA was defined as the failure of the largest connecting pipe 

to the reactor coolant main piping, which is the pressurizer surge line. This is a hot leg

I-1
break. This break is a "good break", because all the flow has go to go through the core 

to go out through the break, so very good core cooling is obtained. About 1966, the 

break was moved to the failure of the main reactor coolant piping, which was a much 

more severe break because the break was larger and not all the flow would go through 

the core. This change resulted in very high-calculated fuel rod cladding temperatures. 

As a result, additional safety equipment was added to plants. On the Westinghouse 

plants, cold leg accumulators were added and the low-head pumped systems were up 

graded and so forth. 

There were a whole series of safety concerns that were occurring in the late 1960s and 

early 1970s, when a large number of these plants were being ordered, and the power for 

the plants was increasing very rapidly. There were a series of hearings that were held in 

the early 1970s called the Core Cooling Hearings, and these hearings provided the basis 

for the regulations that came out in 10 CFR 50.46 and Appendix K of the General Design 

Criteria. The regulations that were a result of those hearings were acceptance criteria on 

the performance of the safety systems and the resulting temperatures and oxidation 

levels and cool ability of the core that were allowable for light water reactors. The 

requirements included: the peak allowable fuel rod cladding temperature cannot exceed 

2200 degrees Fahrenheit; the local cladding oxidation cannot exceed 17 percent; the 

core-wide oxidation cannot exceed one percent; the core must maintain a coolable core 

geometry; (This was never really well defined, in a sense, because fuel rod ballooning 

and bursting is allowed to occur, but the intent here is that a rod-like geometry exists in 

the core, because the heat transfer database is based on a rod-like geometry to generate 

the heat transfer models and correlations that are used to calculate the performance of 

the fuel); and ensure that the safety systems provide long-term cooling of the core to 

remove the decay energy. These are the criteria that have been used to evaluate the 

performance of the plant safety systems, by analysis, using a calculation model of the 

plant. In addition to the criteria, the results of the Core Cooling Hearings were to 

develop a detailed list of modeling requirements, which were specified in Appendix K 

of the General Design Criteria, which reflected the safety analysis knowledge at that 

time. This froze safety analysis technology in a 1972 timeframe and it was very difficult 

to improve the models at that time. One had to license new models that still conformed 

to the Appendix K requirements, which were very prescriptive. The results of applying 

the Appendix K Evaluation Models for the plant safety analysis resulted in lower linear 

heat rates for the plants. The LBLOCA became the limiting transient for Westinghouse 

plants and the LBLOCA is usually the most limiting transient for the other vendor's 

designs as well. 

In 1988, after a substantial NRC confirmatory research program that lasted almost 20 

years, in cooperation also with EPRI and with industry, there was a significant amount 

of effort spent to try to determine the conservatisms in the requirements in Appendix K, 

and to provide a database which could be used then to perform more realistic 

calculations of the LBLOCA accident. One approach was to realistically model the plant 

response to the accident. That is, given the accident, how to more realistically calculate

1-2
the performance of the safety systems and the performance of the reactor system itself? 

This was actually something that the NRC was encouraged to do immediately after the 

Appendix K rule came out, by the American Physical Society. It was felt that the NRC 

would be better served if it had best-estimate analysis capability, because the NRC 

could then make better judgments regarding how the reactor system responded to the 

accident. 

In 1988, after the confirmatory research program was complete and the margins 

associated with the Appendix K evaluation models could be quantified, the rule was 

changed to allow the use of best-estimate methodology. The modified rule permitted 

the additional experimental information that had been generated to improve the 

analysis capability. The ECC performance requirements remained the same, however, 

the main difference now was that a best-estimate approach could be used which 

resulted in a more accurate model which could better predict the performance of reactor 

safety systems such that a higher linear heat generation rate could be calculated for the 

plant. So in a sense, the best-estimate calculations have regained margin in terms of 

kilowatts per foot for the plants using improved safety analysis methods and the results 

of the confirmatory research program. 

1-1.2. The Large-Break LOCA Transient Behavior 

Figure 1 shows a schematic of a Westinghouse four-loop plant. In Westinghouse 

designs, the power per loop is roughly the same, approximately 300 megawatts each for 

each loop. In this fashion, the pumps, steam generators and piping sizes are the same 

for a 2,3, and 4-loop plant design. As seen in the figure, there are a steam generator, hot 

leg, loop seal, reactor coolant pump, and cold leg for each loop. When doing a LOCA 

analysis, the worst break location must be determined. Typically, for a large break 

LOCA, the worst break location is in the cold leg. The cold leg, at the pump discharge, 

has been found to be the most limiting-break location for all the designs, not only 

Westinghouse but other designs as well. 

For a large break in the main reactor cooling piping system, the break flow area is very 

large (typically 9 square feet) so that the transient is very rapid. There are three periods 

for the large-break LOCA: blowdown, refill, and reflood as shown in Fig. 2. The 

blowdown period which is approximately 30 seconds long is when the initial inventory 

that is in the reactor system, is lost out the break. The passive cold leg accumulators will 

start to inject as the plant depressurizes and there can be some high pressure pumped 

safety injection also occurring as the plant is depressurizing. The timing of these 

systems is design specific and depends on the set point pressure of the accumulators 

and the start-up time of the emergency diesels for the plant.

1-3
REACTOR COOLANT PUMP

Westinghouse NUCLEAR STEAM SUPPLY SYSTEM 

MB 3595

1-4
2500

-'2000 

SLOWDOWN REFILL

cc 1600 

C 1000 C. 

0 LPSI 

6 00 ACCUMULATOR INJE CTION 

0 INJECTION |N EC I_ 1 I " I ,!, 

0 5 10 is 20 25 30 35 40 

TIME (SEC) 

The blowdown is a very rapid transient and there will be flow reversals that occur in 

the core. The core will quickly go through critical heat flux, and the reactor will shut 

down due to voiding in the core. No credit is taken for control rods to shutdown the 

fission reaction in the core, at least for PWRs. The core flow reversals that occur depend 

on the pump and break interaction, and depend upon how many loops there are in the 

plant, which means how many pumps there are in the plant. One can calculate more 

upflow or more downflow depending on the number of pumps (loops) and the break 

size. This will result in different degrees of cooling depending on whether the flow is up 

or down in he core and the amount of liquid carried in the flow. The high-pressure 

safety injection will automatically initiate for most plants during blowdown, but this is 

a very low flow, high head system. Most of the water that is injected from this system 

basically goes straight out the break. 

Cold leg accumulators will initiate in the Westinghouse plants and in the Babcock & 

Wilcox plants at a set point pressure of approximately 600 psi. These are passive 

injection systems. The accumulators are large tanks full of cold borated water, which are 

isolated from the primary system by check valves, and when the system pressure drops 

below the set pressure in the check valves these tanks start to inject. In the Combustion 

Engineering plants most of the cold leg accumulators are set at a set point pressure, 200 

psi, so there is a longer delay before those accumulators start to inject. 

ECCS bypass occurs during blowdown as shown in Fig. 3. ECCS bypass means that the 

accumulator flow that is injected into an intact cold leg can be entrained by the steam 

flow coming up the downcomer and swept around the downcomer annulus and out the 

break. If there is a high upflow in the downcomer annulus, the flow that is injected can 

go around the annulus and out the break instead of penetrating down the annulus to fill 

the vessel as seen in Fig. 3. In the original Appendix K rules, there were stipulations on 

the amount of credit that can be taken for the injection flow from the accumulators. It

1-5
was determined that the accumulator flow would be bypassed until it could be shown 

that the steam flow up the vessel was low enough that the liquid flow from the 

accumulators could flow down the downcomer into the lower plenum.

S- ACCUMULATOR 

REACTOR INJECTION 

COOLANT 

PUMP

Cold Leg Break Steam Flow Path 

Since the 1972 core cooling hearings, there have been full-scale tests run on ECC bypass 

in Germany as part of the NRC program of confirmatory research. What was observed 

in those experiments was that for situations in which there was accumulator injection in 

the loops on the opposite side of the break, the flow from these accumulators would 

penetrate into the downcomer and begin to refuel the lower plenum, so ECC bypass 

earlier would actually end earlier. The experiments also showed that the accumulator 

in the broken loop was completely lost and the accumulator in the cold leg next to the 

broken cold leg would also bypass a large fraction of its flow during blowdown. Best 

estimate computer codes have been used to predict the accumulator flows and bypass 

flow rates in a more accurate manner. However, the code predictions tend to 

underestimate the amounts of water, which penetrate the downcomer and hence, over

predict the amount of accumulator water, which is bypassed. This is a conservative bias 

in the best-estimate computer codes.

1-6
Toward the end of blowdown, the minimum vessel inventory is reached because most 

of the initial reactor coolant system inventory is blown out the break. 

Now as indicated earlier, there are three phases for the large break LOCA. The 

blowdown lasts approximately 30 seconds. This is when the primary inventory is 

vented out the break; the core goes through DNB, voids, and shuts down. The cladding 

heats up and this is primarily due to the stored energy temperatures of 1600 degrees 

Fahrenheit can be exceeded. The next period is refill, which is approximately ten 

seconds in duration, during which refill of the reactor vessel begins, primarily filling the 

lower plenum to the bottom of the heated length of the fuel rods, and to beginning to 

reflood the core. 

During the "refill period," the lower plenum is being refilled. ECC bypass is ended, and 

the core is in a stagnant steam environment and is heating up almost adiabatically. 

There can be some radiation heat transfer to the guide tube thimbles and rod-to-rod 

radiation heat transfer. 

The low-pressure safety injection system starts to activate since the reactor system has 

depressurized and the reflood process has begun. Figure 4 shows the reflood process as 

the core begins to reflood. Fuel rod burst can occur in this time period because the 

cladding temperatures are very high. They are generally in excess of 1600 degrees 

Fahrenheit, and now the system pressure is very low, so the differential pressure across 

the cladding from the inside of the cladding to the system itself is very large, and clad 

ballooning and failure can occur in this timeframe. 

Again the safety systems that the plant has to cope with the LBLOCA transient are 

shown in Fig. 5. There are the cold leg accumulators where there is typically one per 

loop, one per cold leg. There is a high pressure (high head) safety injection system with 

injection into each loop. Some plants also have intermediate pressure injection systems. 

All plants have low pressure safety injection systems, which inject into the cold legs. 

The different safety injection systems will initiate at different pressure levels depending 

on the plant design. The systems are loaded on emergency diesels, so in the analysis the 

start time of the diesel and the delay times of the signals to start the systems must be 

considered. These pump systems pull suction from the refueling water storage tank, 

which is outside the containment. The refueling water storage tank typically contains 

about 350,000 gallons of borated water. The pump systems will inject water from the 

tank up to 20 minutes and then will automatically switch over and pull suction from the 

reactor sump. It is the pump systems that provide the long-term decay heat removal, 

post accident, long term cooling for the core. The highest flows into the reactor vessel 

come from the accumulator and then the next highest flow comes from the low pressure 

injection systems.

1-7
Cold Leg Break Steam Flow Path Schematic Drawing 

1-8
I, . 

L.n......... 

••1olCore Cooling System

Once the lower plenum is refilled, the core then begins to reflood as seen in Fig. 4. 

Reflood typically starts at about 40 seconds after the accident and can last until up to 

600 seconds before the core is completely quenched and the fuel rods are approximately 

at the saturation temperature in the entire core. The accumulators are designed and 

sized such that there is sufficient water in the accumulators to fill the downcomer with 

water. This provides a driving head for reflood. The reflooding situation is a passive 

gravity-driven situation. Basically, the accumulators fill the downcomer, which 

provides 16 feet of driving head to forcd the flow into the core. The flow goes into a hot 

core, generates a two-phase mixture, provides cooling and exists the core, goes through 

the hot leg, into the hot steam generators through the cross over leg, through the reactor 

coolant pump (which acts as a large resistance), to the cold leg, into the downcomer 

annulus, around the annulus and out the break. The reflood flow path is shown in Fig. 4 

and the details of the reflood process in the core are shown in Fig. 6. 

Since the system is depressurized so that the primary pressure now is well below the 

secondary side pressure, the steam generators secondary side is much hotter as 

compared to the primary side saturation temperature. The stagnant water on the 

secondary side (note, the secondary side is isolated for the transient) is typically 500 to 

525 degrees Fahrenheit, so as a two phase mixture is swept into the steam generator 

tubes on the primary side, it will all evaporate or a large fraction of it will evaporate.

1-9
This will result in an increased steam flow rate of superheated steam leaving the steam 

generator, which will result in a large pressure drop across the reactor coolant pump since the pump is a large hydraulic resistance. Since, there is only 16 feet of water in the 

downcomer to force flow to the core and to overcome the pressure drops in the steam 

generator and the reactor coolant pump. Therefore, the flooding rate into the core is 

very low; typically it is one-inch per second or less for the majority of the reflood 

transient. This is called the "steam binding effects" for PWR reflooding.

0-O.47 if ,

DISPERSED FLOW AND HEAT TRANSFER

I TRANSITION. 'IN FILM FLOW 

OUENCH FRONT 

LEATE EUBBLY. 

ING SLUG FLOW

LIOUID 

FLOW

Hydrodynamic and Heat Transfer Regions Nea Quench Front

• - Typica Condlt;on in Rod Bundle During Reflood

1-10
If the plant design has a very low containment back pressure, like a Westinghouse ice 

condenser plant, the steam specific volume is very large at low pressure, such that for 

the same mass that gets evaporated there is a much larger volume of steam that has to 

vented resulting in larger pressure drops in the steam generator and pump such that 

the flooding rate for these designs are even lower. 

The important thing to note is that reflood is a gravity driven process. Pumps are 

injecting flow into the cold legs. The best they can do is keep this downcomer filled. 

Once the downcomer is filled, all subsequent water that is injected spills out of the 

break. After this point, no amount of pumping capacity contributes to putting cooling 

water in the core. 

The peak cladding temperature, while it's different for different designs, for the 

Westinghouse plants, it will generally occur at reflood, either at the beginning or later 

during the reflood transient. For the B&W plant that is shown, the peak cladding 

temperature is calculated to occur at the very beginning of reflood. 

Reflood is most critical portion of the transient. The heat transfer is generally the lowest 

here. The fuel rods will continue to heat up as the quench front advances and the peak 

cladding temperature is being pushed up the fuel as the quench front moves up. 

The fuel rods are being cooled by dispersed flow boiling, heat transfer regime as seen in 

Fig. 6. If the rods have not burst during refill, they generally will burst during reflood, 

which results in flow blockage within the core. The flow blockage results in flow 

diversion and an additional peak cladding temperature penalty. As indicated earlier, 

the core flooding rates are low. In a Westinghouse plant they quickly drop below one 

inch a second flooding rate, so it take a long time to fill the reactor vessel and to provide 

the cooling. The dispersed, two-phase nonequilibrium mixture of superheated steam 

and drops cools the fuel rods. The calculated peak cladding temperatures exceed 1800 OF 

such that zirconium/water reaction can occur at these high cladding temperatures. 

The PCT generally occurs during reflood. The fuel rods will eventually cool down and 

quench. It takes approximately 5 minutes to quench the core. 

1-1.3. Westinghouse Appendix K Evaluation Model LBLOCA Calculations for a Four 

Loop PWR 

Calculations have been performed for a four-loop Westinghouse PWR using an 

approved Appendix K type of an evaluation model. One of the requirements of 

Appendix K is that different break sizes be simulated by changing the discharge 

coefficient on the break. The event timing for different break sizes are given in Fig. 7 

and the calculated results for the different break sizes are given in Fig. 8. As Fig. 8 

indicates, the 0.6 Double-Ended Cold Leg Guillotine (DECLG) is the most limiting break 

for this particular plant. The most limiting break size is plant size and type dependent 

and must be determined by performing a series of calculations for different break sizes 

or discharge coefficients to obtain a break spectrum. Figure 7 gives the timing for the

1-11
different DECLG transients for this plant. As the figure indicates, for the 0.6 DECLG 

the accumulator injection occurs at approximately 15 seconds, pump injection occurs at 

about 30 seconds, and blowdown ends at approximately 30 seconds. The bottom core 

recovery, which means the end of refill, occurs 13 seconds later (approximately 43 

seconds). With blowdown ending at 30 seconds, the accumulators are empty at 53 

seconds. This is the time period to fill the lower plenum and to fill the downcomer with 

the remaining flow in the accumulator, so a large fraction of the accumulator flow has 

been lost in this Appendix K analysis. A large fraction of the accumulator flow is lost 

because they have started injecting at 15 seconds but credit is taken only for the flow, 

filling the vessel toward the end of blowdown at 30 seconds. The hot rod has burst 

during the refill period in this case and the peak cladding temperature is occurring 

approximately three minutes into the transient as indicated in Figs. 7 and 8. Figure 8 

gives the calculated value and location of the peak cladding temperatures for this range 

of breaks. For the most limiting break size, and the peak cladding temperature was 

calculated to be 2000 'F and it occurred at the eight foot elevation. . The calculated PCT 

is below the Appendix K limits and the local zirconium/water reaction is only 8 %, 

which is also under the limit of 17 % such that this represents an acceptable licensing 

calculation for this plant. The fuel rod burst occurred at six feet, the peak power location, but peak temperature occurred at eight feet as indicated in Fig. 8. What is 

usually found is that the peak, even with a symmetrical power shape, is usually above 

the peak power location. Therefore, power shapes, which are more limiting, are those 

shapes in which the power is skewed to the top of the core. The peak location is 

displaced toward the top of the core because reflooding is from the bottom of the core, 

and this tends to push the peak temperature location up the rod. 

CD = 0. 4 CD = 0.6 = 0.8 

DECLG DECLG DECLG Time (Sec) (Sec) (Sec) 

Break 0 0 0 

Reactor trip signal 0.777 0.751 0.735 

SI-signal 1.65 1.33 1.15 

Intact loop accumulator 20.00 15.10 12.50 

injection 

Pump injection 30.65 30.33 30.15 

End of bypass 38.029 29.348 25.334 

End of blowdown 38.035 29.348 25.334 

BOC time 53.637 43.601 39.175 

Intact loop accumulators 60.247 53.441 50.225 

empty 

Hot rod burst time 57.08 40.00 44.26 

Peak clad temperature (PCT) 67.44 165.32 89.92 

time

1-12
LARGE BREAK LOCA RESULTS FUEL CLADDING DATA

MINIMUM SAFEGUARDS 

CD =0.4 CD = 0.6 cD = 0.8 

DECLG DECLO DECLG 

RESULTS 

Peak clad temperature, F 1804.1 2014.0 1776.0 

Peak clad temperature 7.00 8.00 8.O0 

location, ft 

Local Zr/HzO reaction, 2.64 5.39 2.27 

maximum % 

Local Zr/H2 O reaction, 6.25 8.00 6.25 

location for maximum 

reaction, ft 

Total Zr/H1 0 reaction, % <0.30 <0.30 <0.30 

Hot rod burst location, ft 6.25 6.00 6.25 

Figure 9 shows the calculated core flows during the blowdown period. There is 

stagnation initially in the core and the flow exits out of the top and the bottom of the 

core. As blowdown continues, there is a positive core flow period and then core flow 

period during blowdown. This is caused by the interaction of the main reactor coolant 

pumps, the break, and the flow resistances between the core and the break. The pump 

flow in this case is for a four-loop plant, which is greater than the break flow early in 

time before the pumps cavitate. The difference between the break flow and the pump 

flow results in positive flow through the core, so the core heat transfer improves. But at 

some time, the pumps cavitate and loose their pumping capability. The break is still 

demanding the same flow, so the flow then reverses through the core and there is a 

negative flow through the core, which also provides some core cooling. For a 

Westinghouse three-loop plant, this zero flow point on Fig. 9 would shift upward 

because now instead of having three pumps to provide the flow to the break, there are 

only two, remembering that the cold leg pipe size (break size) is the same for all 

Westinghouse plants. If a two-loop plant is examined, all the flow is negative, or down 

through the core, during blowdown because the break demands more flow than the one 

pump can provide.

1-13
CORE EPOTTOti

j I I

i I 

-OP

U * .. 1 . - S

I.' I C 

CJ 

--3 

LI.

I I I I

17.5 20. 22.S 2=-. 27-. 50.

The pressure transient is shown in Fig. 10 and indicates that the containment pressure is 

reached in approximately 30 seconds. The accumulators begin injecting when the 

pressure typically reaches 600 psi. Again the accumulators are a passive injection 

system which provide high flow but for a relatively short period of time. The 

accumulator injection flow is shown in Fig. 11. In a four loop plant there are three 

accumulators that are injecting into the intact cold legs. I J il il I I I I 

-2 MW. . COPE BOTTOM I-) TOP * (. ] 

SO D il I 1 I i I ' I 

______ i i

12.5 H. c?.5 17. S 20. 22.5 2i. 30. 

7'Atz ,SEC)

1-14

E''..

I I 

4iI0O. I I

-2000.1 U . I I 

SI I I I I I I I 1 1

I .0o. I 

2.

I i I

2.S E. 7.E IV. 23. 5 5. 

II.C tSsEC

11 .406Aý

I 1 1 I .,v • m

I 

I I I i I

i

I I I I I

I

I i i i
I I I I I

So 5DOO. 

...Q 

&LI 

AVODD. U 

M,1 

3000. 

2ODD. 

"- ooD

2.S S. 7.2- 0. 2.- IS. 17.S 

TIM1E (SEC)

20. 22.S 25. 27.S SD.

In the reflood period, the downcomer level increases and provides the driving head to 

force the water into the core. To overcome the pressure drop in the loops. a liquid level 

in the core is built up as seen in Fig. 12. The downcomer level is the effective driving 

head to provide cooling into the core. The fuel rod quench front moves up the fuel rods 

as the core level increases resulting in cooling of the upper elevations of the core such 

that the temperature turn over and the rods will eventually quench. 

22.-[ 

DOTWNCOMER 

LL. 1 . 7 

I I.

1-15

7000.

!i I I I I , I 

i I"I 

iil j I ' IY

0.

I i I ! i i i !! !
The calculated hot rod temperatures at the burst location and the peak location are 

shown in Fig. 13. In this calculation, the hot rod peak cladding temperature is located 

approximately two-feet above the burst location. When ballooning and burst occurs, 

the gap conductance in the fuel rod at the location of burst is significantly reduced 

because the cladding has moved away from the pellet. Therefore, the pellet is thermally 

isolated from the cladding such that the cladding acts like a large fin. Since there's less 

stored energy, in the cladding, it is much easier to cool, so the cladding easily cool at the 

burst location. The fuel rod burst location is usually not the limiting peak cladding 

temperature location. However, there are cases when it can be. This becomes very plant 

specific.

* • . - -, 

.25E.+D4 

.17SE404 

* 75E0., 

.GE:4. 

-7GO4

La 

W tA 

M 

L9 

0 

0 

0 

43: LU

* j=......~ . I I I ii [ t 

- , 1j !.• '"- ... 1 

_ _I I; 

I I 'I 

_ BURST e--- PEAK I ,j 

KB III• 9

LJ.

T I ME

I SO. 

SEC.

20D. 250.

The peak cladding temperature location, in this calculation, is at the eight-foot elevation 

on the rod, whereas this burst location is at the six-foot elevation on the rod. Double

sided zirconium/water reaction is considered at the burst location. However, as the 

temperature plot indicates, the cladding temperature decreases upon burst such that the 

effects of double-sided reaction are not significant. In the Appendix K calculations, 

there is no direct fuel-coolant heat transfer. All heat transfer assumes the presence of 

the cladding and the thermal resistance of the fuel pellet-clad gap. The fuel is assumed 

to stay within the rod and fuel is assumed to stay in a rod like geometry, even in the 

burst zone.

1-16

9J3.

0 0 Z
1-1.4. Babcock and Wilcox Lowered Loop Appendix K Evaluation Model LBLOCA 

Calculations 

Next, a Babcock and Wilcox (B&W) lower loop Appendix K Evaluation Model 

calculation will be examined. This plant design has lowered primary loops and uses 

once through steam generators. The break of interest is an 8.5 square foot DECLG 

break on the pump discharge. The rated plant power is 2772 MWt. The calculation uses 

a bottom peaked power shape, which is limiting for this plant design, since the PCT 

occurs at the very beginning of reflood which will be explained shortly. 

Figure 14 shows the B&W plant. There are two unique features. One is the once

through steam generator, so this is like a Three Mile Island type unit, and there are 

barrel check valves in the upper downcomer. The purpose of these check valves is to 

provide a vent path for flow, reflood flow that would be coming out of the core. 

Instead of having to go up the long and high hot leg configuration (the candy-cane) and 

down through the hot steam generator, the flow can vent out through the check valves 

in the upper plenum to the downcomer, around the downcomer annulus and can go 

out through the break. The vent valves provide a short circuit flow path to the break. 

There are approximately 10 large vent valves around the vessel. These are basically 

flapper valves and are shown in Fig. 15. During normal operation they are held closed 

because of the positive pressure difference in the downcomer, relative to the upper 

plenum, because of the pump flow entering the downcomer. During reflood there is a 

higher pressure in the core and upper plenum so that the flapper valves can open and 

vent out the two-phase mixture from the core directly to the downcomer annulus, 

which then flows to the break. This process short-circuits the reflooding portion of the 

transient because there is no need to force the flow through the huge, hot, once-through 

steam generators because the reflood flow can vent out through the vent valves. 

Therefore, reflood is not the limiting portions of the transient for this type of plant 

design. The time period that is more limiting in this design is the end of refill just before 

the beginning of reflood. 

The pressure transient is shown in Fig. 16 for the B&W plant during the blowdown 

period. Since this for a double-ended break, the blowdown proceeds very quickly and 

the vessel is basically depressurized in 30 seconds. For this design there will again be a 

short positive core flow but a very, very strong negative flow down to the core, as seen 

in Fig. 17. At the end of blowdown, the reverse flow through the core diminishes 

because by the end of blowdown, there is little water remaining in the reactor system. 

Because a vent path through the barrel check valves to the downcomer and break exists, 

the loop resistance is negligible and the calculated flooding rate into the core is much 

higher than a typical Westinghouse plant as shown in Fig. 18. The flooding rate into the 

core for a Westinghouse plant would be typically be around one inch a second. The 

flooding rate for the B&W plant is typically two inches a second.

1-17
,PRIMARY

SECONCARY

1-18
Figure 1. Vent Valve Arrangement (B&W 136Z14E) 

LACKSItEE 

Coll -,S. uppol 

LCu

SECTION Z-Z

-I

w 

C) 

0 

0 x

Zdo
REACTOR VESSEL UPPER PLENUM PRESSURE.

28

a. 

U? 

5 

0 

0

HOT CHANNEL MASS FLOW RATES.

TIMI. SECONDS

1-20

TIME. SECONDS

so 

40 

0 0 

15 

-0 

1⁄4-.
CORE FLOODING RATE..

TIME._SECONDS

The peak cladding temperatures that are calculated for this design typically occur at the 

end of the refill period or beginning of the reflood time period. Given the good 

reflooding heat transfer behavior, with the higher reflood rates due to the vent valves, 

the peak cladding temperatures quickly turn over and the rods and quench. 

As a result, B&W plants tend to have higher allowable linear heat rates because the vent 

valve plants are not reflood limited. The calculated peak clad temperatures are 

occurring during the end of blowdown and the refill period, where the PCT then is 

more controlled by the stored energy in the fuel and not by the duration of the reflood 

transient, as it was in the Westinghouse plant. 

The limiting axial power profile is located 2.5 feet from the core inlet at the bottom of 

the core. The reason for this location is because of the downflow period at the end of 

blowdown and the beginning of refill in which the flow is coming down through the 

vessel. For this flow situation, it is more conservative to have power peaking at the 

bottom of the core further from the source of cooling, which in the B&W plant case is 

the upper plenum during blowdown. 

The calculated LBLOCA transients for the B&W plant are also shorter as compared to 

the Westinghouse transients, since reflood is not limiting. Examining Figs. 19 and 20 for 

these calculations, the PCTs occur earlier as indicated in the figures and the whole 

transient is much shorter compared to the Westinghouse plant, so there is a real benefit 

in having barrel check valves. Figures 19 and 20 show both the cladding and fuel

1-21

5,

0 

0 rU

14

12 

'10 

B8. 

0 

'4 U

0
temperatures at the burst location as well as at the peak non-burst location. At the burst 

location, the cladding has deformed and has moved away from the fuel pellet so it takes 

longer for the fuel temperature to decrease (Fig. 19) as compared to the fuel 

temperature at the non-burst location in which the fuel/cladding gap heat transfer 

coefficient is larger (Fig. 20). 

HC FUEL & CLAD TEMPERATURES AT RUPTURED LOCATION. 

2400 . LEGEND 

• . 1500 -= •Fuel, Seg6 

2000:• . C0.1ad, Seg 6 

1250 

=.1600, ,, 

1000 

1 750 

4 0 0 . . .o0 S o.. 

00 

0 40 80 120 160 200- 240 

"TIME. SECONDS 

HC FUEL-& CLAD TEMPERATURES AT PEAK UNRUPTURED LOCATION. 

2400 LEGEND 

1500 

Fuel, Seg 7 

2000 • Clad, Seg. 7 

• . .. 1250' i Soo. ./ 10 

ioo 

~1200 

too 750'

.. 120 

TIME, SECONDS

1-22
1-1.5. Westinghouse Best-Estimate Model LBLOCA Calculations for a Four-Loop 

Plant 

The previous calculations, which were shown, were Appendix K type calculations using 

the types of models that were approved in the 1972 timeframe. There have been 

improvements and modifications made to these models, but generally these models 

used the prescriptive approach given by Appendix K. Next, a sample best estimate 

calculations, again for a Westinghouse plant, will be examined. Now these calculations 

use a best estimate code, in this case WCOBRA/TRAC. The analysis conforms to the 

rule change in 1988 that allows using a best estimate computer program. 

The timeframes for the best-estimate transient are pretty much the same as the 

Appendix K calculation for the Westinghouse plant as seen in Fig. 21. There is still have 

roughly a 30 second blowdown period followed by a 10 second refill period, and then a 

reflood period that can last up to 5 minutes until the core is completely quenched. The 

change that is different in these calculations is that they are performed using a best 

estimate method, which gives a more accurate representation of the plant response to 

the imposed LBLOCA transient. The calculated allowable peak linear heat rates are 

higher for the best-estimate calculations, as compared to the Appendix K calculation. 

The Appendix K models and requirements were more limiting as compared to the best

estimate models. As a result, the allowable peak linear heat rates for the best-estimate 

calculations are higher, so the peak kilowatts per foot in the analysis that is being 

analyzed is higher. 

The benefit of the best-estimate calculations, relative to Appendix K calculations, is that 

best-estimate LBLOCA calculations would allow the plant to operate at higher kilowatts 

per foot since the safety analysis of the plant would be performed at higher kilowatts 

per foot limit. 

The temperatures that are calculated using the best-estimate analysis are shown in Fig. 

22 for this particular plant. In a best estimate methodology, the nominal value PCT is 

calculated and then the uncertainties are convoluted in a statistical fashion to calculate a 

95th percentile PCT value, replaces the Appendix K value as the licensing basis for the 

plant. The PCT values shown in Fig. 22 are not that dissimilar from what was shown 

earlier for older Appendix K calculations. The difference is that the best-estimate 

calculations are done at higher kilowatts per foot, so that there are temperatures that 

are typically 16-17-1800 0 F at blowdown and approximately 2000 0F during reflood. 

The best-estimate calculations still must conform to the same original 10CFR50.46 

requirements that the Appendix K calculations had to achieve. The difference is that 

the same criteria can be achieved with best-estimate calculations, but at a higher linear 

heat rate.

1-23
B 

L 

0 

0 

w 

N 

R 

E 

F 

I 

L 

L

0 sec. Break occurs

Reactor trip (pressurizer pressure or high containment pressure) 

Pumped SI signal (pressurizer pressure or high containment pressure) 

Accumulator injection begins 

Pumped ECCS injection begins (offsite power available) 

Containment heat removal system starts (offsite power available)

20-25 sec. I End of bvvass

I.-

End of blowdown

Pumped ECCS injection begins (loss of offsite power) 

Containment heat removal system starts (loss of offsite power) 

Bottom of core recovery 35-40 sec.

R 

E 

F Accumulators empty 

L 

0 

0 

D 5 nin. Core quenched 

L 

0 

N 

G Switch to cold leg recirculation on RWST low level alarm 

T 

E 

R 

M Switch to hot leg/cold leg recirculation 

C 

0 

0 

L BEST ESTIMATE LARGE BREAK LOCA TIME SEQUENCE OF EVENTS 

I 

N 

B 

BEST ESTIMATE LARGE BREAK LOCA ANALYSIS RESULTS

Component 

PC~rv~mze

Blowdown Peak First Reflood Peak Second Reflood Peak 

<15081F <1681°F <13841F

<1760"F

Maximum Oxidation

Total Oxidation

<1976 0F 

<11% 

<0.89%

<1964°F

1-24

I
Having higher linear heat rates allows the utility to operate the plant in a more cost 

effective manner. The best-estimate analysis methods identify LOCA margin that is 

kilowatts per foot margin that can be used to more efficiently operate the plant. LOCA 

margin can be used to make changes in the design, usually in the fuel, which requires 

the use of higher kilowatts per foot. For example, removing the fixed poison rods in the 

guide tube thimbles in a PWR fuel assembly, which requires higher allowable peak 

linear heat rates. Another example is using lo-lo leakage loading patterns so the fuel on 

the edge of the core is almost dead, such that the power is pushed into the center of the 

core, resulting in higher kilowatts per foot. As fast as LOCA margin is developed, it is 

used by the utility to improve the cost effectiveness of the plant. Therefore, similar 

peak cladding temperatures are calculated using the best-estimate methods, however, 

the plant now operates more efficiently or is using more efficient fuel designs such that 

the fuel cycle costs are lower. 

Figure 23 shows the locus of peak cladding temperatures anywhere in the core. The 

peak is occurring at the value is that is nearly 2000 0F, and it is occurring at about 9.7 

feet. The PCT is occurring at roughly the same time frame as in the previous Appendix 

K analysis. Again, while the PCT is about the same, kilowatts per foot, and the 

integrated power in a hot assembly in the best-estimate analysis is higher than in the 

Appendix K analysis.

2200 

2000 

1800 

1600 

. 1400 0 

E t200 

I1000 

800 

600

12

11 

10 -' 

9I 

10 

7

1-25

17 SM ,- _r4-5~ •, 7-
Figure 24 shows the flow rate at the top of the core. In this figure and in the following figures, there is an initial steady state period before the transient, which is modeled in 

WCOBRA/TRAC as indicated on the figures. At 20 seconds, the break opens and the 

transient begins. Figure 24 shows the flow reversals during blowdown as well as the 

flow out of the core during the reflood period. 

Figure 25 shows the calculated pressurizer pressure and indicates that the system 

pressure is quickly down to containment pressure in 30 seconds after the break is 

initiated. 

Z *o 

W2 E 

*00 

(s/wq1) 91D8 MO1J SSDIl 

E 

o I 0 I , I .0

1-26
Figure 26 shows the liquid inventory in the lower plenum, and indicates that most of 

the water is initially blown out of the lower plenum, even in a best-estimate calculation. 

But the cold leg accumulators begin to inject and the lower plenum is quickly refilled 

and stays water filled for the remainder of the transient. The time period during which 

the water has been blown out of the lower plenum is the minimum mass inventory in 

the vessel. The real point of interest is how quickly the emergency core cooling systems 

refill the vessel. 

Figure 27 is a plot of the reactor vessel water mass and shows how quickly the ECC 

refills the vessel. Most of the water inventory is due to the accumulator injection, which 

is then followed by the pumped injection, which maintains the mass inventory in the 

system.

1-27
Figure 28 shows the accumulator flow from one cold leg accumulator, which injects 

very rapidly the vessel. The accumulator injection recovers the mass in the vessel and 

the pump system continues to inject and maintains flow into the vessel for an extended 

period of time as seen in Fig. 29.

-

( S

0 1 D~

C 

I I I

A 0 _ S S Dr

C 

_____ _____ _____ _____I_____ _____ B 02

t + 4 L 0 

E 0

0 

C

(S/wqj) 9OID8 MIJj SSDVJ

1-28

C 

Sn

CO
The accumulator injection also fills the downcomer such that the liquid level increases 

as seen in Fig. 30. It is the level in the downcomer that provides the driving head for 

flow to enter the core and to begin reflood as seen in Fig. 31. In this calculation, level in 

the downcomer is recovered with the accumulator injection, and then it is maintained 

due to the pump injection. At the same time, the flow from the downcomer increases 

the liquid level in the core, since the head in the downcomer is forcing water into the 

core. This is what provides the reflood cooling which results in turning the fuel rod 

temperature around (PCTs) in the reflood portions of the transient. As Fig. 31 indicates, the core level is initially recovered early in reflood and then remains approximately as 

constant. 

35 1- 

=20 -. 

10 

o REFERENCE S N I LOOP I DOWNCOMERT

I COLLAPSED UQUID LEVEL

1-29
Effects of High Burnup on the LBLOCA Calculations

In this part of the presentation, the effects of high burnup on the LBLOCA were examined and some conclusions on burnup effects were developed from the LBLOCA analysis. Figure 32 shows calculated fuel rod internal pressures for two different fuel rod codes. This solid line is the FUELROD code, which is a PC code that is used at Penn State in a Reactor Engineering class. The dashed line in Fig. 32 is FRAPCON code calculation. These are calculations have been done for the same fuel rod design, plenum design, and power history. The result is for rod internal pressures, which reflect the differences between the codes, primarily the fission gas release model in each of the codes. Comparison calculations have been made between FRAPCON and FUELROD at Penn State and everything else agrees pretty well except for the amount of fission gas that is released in the calculations. So there can be a range of calculated internal pressures for the rod design. The purpose of showing these two calculations is to indicate that there can be a range of internal pressures. For high burn-up fuel, the fuel rod internal pressure can exceed the system pressure. The FUELROD code calculated an internal pressure of about 2700 psi inside the rod during operation at approximately 80,000-MWd/metric ton burn-up, where as FRAPCON calculated a little over 2000 psi. 

Looking at a loading pattern for core, most of the designs today are going to lower 

power on the outside of the core to basically provide some degree of shielding to the reactor vessel. Utilities are using these low leakage loading patterns so the powers of the outer row of assemblies are very low relative to a normalized power of unity. An example of such a loading pattern is shown in Fig. 33. The fuel assemblies on the outside edge of the core represent once and sometimes twice burned fuel assemblies, so burned fuel assemblies are moved out to the outside edge, and the fresh assemblies are loaded into the center regions of the core. The once and twice burned assemblies are the assemblies which are going to have the higher initial internal pressure in the fuel rods because they have been in the core for a longer time, but at the same time, they are the low powered assemblies. Most of the burned assemblies are loaded on the outside edge of the core. However, there are loading patterns that load a twice-burned assembly in the very center of the core to act as a power suppresser or poison. This twice-burned assembly can be driven to higher powers, then the edge assemblies, by the high power fresh assemblies, which surround it. 

Considering the core nodding scheme used in the WCOBRA/TRAC Model that was used for the best-estimate calculations as seen in Fig. 34, WCOBRA/TRAC models this 

type of a loading pattern with different types of assemblies and different fuel rods. There is a unique representation for the low power fuel assemblies. All the low powered edge assemblies are lumped together as a single channel in the calculation. The interior of the core region, in the center, is broken up into fuel assemblies, which 

are under open holes, support columns or freestanding mixers, or fuel assemblies, which are under guide tubes. There is a separate channel, which represents the hot 

assembly.

1-30

1-1.6.
Figure 3-2 Inner Pressure Versus Time 

3000 ---------- r ---- --- r -----r ---------- r ------ r- -------- --- --------------- •' 

2500 --------------- ---- ---- -------- ---------- I ---------

Cyl I 

~l0----------- -------- ------- --------------- 1----#-----I---------------------- -- -

IIIIII I I I I 

1 L - I I -I I 

i a I - fuelrod I 

I I I I I . . .I 

2500i IFrapcon3 

I II I a IIIIII I I I 

0 r -- - - - - r a 

I I ,I I , I a 

0 200 400 600 800 1000 1200 1400 1600 1800 2000 

Time (Days)
WESTTNGEOUSE PROP3ET•ARY CLASS 2C

211.80 MMD/b= 

10 PPM 

1.383 CORE F-DELTA-H 1 1.629 CORE FQ

lOB 12 104 f 21 11 11 12 94 .884 1.333 -908 1.028 1.000 1.070 1.186 .403 .911 1.374 .951 1.062 1.023 1.137 1.337 .606 

1-048 1.625 1.094 1.211 1.165 1.318 IL1.590 .696 62410 26256 '" 52628] 48482 48393 147051 22856 46228 

1.333 1.079 1.328 1.295 1.061 1.309 1.105 .344 

1.374 

'12 if 

1.117 1.383 

212 221 2 12 10B 1.371 1.160 1.381 1.306 .622 1.625 1.285 1.627 1.609 1.339 1.629 1.546 .719 

2625 49664 28247 28172 45940 27171 22638 46783 10k 1 11-r i- r I; -. ___ _ "V .-

.908 

.951 

1.094 

52628

1.328 

1.383 

1.626 

28233

1.050 

1.104 

1.266

113 12 - i 1. 22 ' 12 10,k. 1.028 1.294 1.053 1.069 1.299 1.1•9 .381 

1.062 1.371 1.109 1.174 1.373 1.311 .673 1.211 1.608 11 1.261 1.355 1.611 1.542 .773 48482 28156 45866 46414 28363 23311 1 43262

1.053 

1.109 

1.260

1.044 

1.096 

1.231

1.254 

1.357 

1.596

.702 

.969 

1.115

1.000 

1.023 

1.165 

48393

1.070 

1.137 

1.318 

47051

1.186 

1.337 

1.590 

22856

9k 

.403 

.606 

.696 

46228 

1

1.061 1.160 

1.339 

45960

12 

1.309 

1.381 

1.628 

27871

1.105 

1.306 

1.547 

22646

ion 

.344 

.622 

.718

46790

1.044 

1.097 

1.251

49090 128363 .LA 

1.254 

1.357 

1.596 

26816

1.299 

1.373 

1.611

1.109 

1.311 

1.542

23313 386

1.164 

1.343 

1.580

11 

.587 

-920 

1.049

11 

.103 

. 969 

1.115 

40631

10k .382 

.674 

.773 

43257

I 43257

11 

.587 

.920 

1.050

lIAmarx 3cw

NUMBER OF POWER BURNUPS 

ASSEMBLIES SHARING TOTAL CYCLE

4 

9 

12 

64 

68

.403 

.404 

.557 

.949 

1.240

46228 6908 

48523 7241 

46382 11047 

45585 20622 

26163 26163

CYCLE 10 ASSEMBLYWISE POWER AND BURNUP AT 

21180 MWD/MTU, HFP, ARO, EQUILIBRIUM XENON

1-32

REGION 

IDENT. 

9A 

10B 

10k 

11 

12

I
Note: 

In Figure 

(1) ROD 5 is shown as LP, 

(2) ROD 4 as OH/SC, 

(3) ROD 3 as GT, 

(4) ROD 2 as HA, and 

(5) ROD 1 is HOT rod. 

In the LOCA calculation that is done, there are typically have five fuel rods which 

model the different assemblies in the core. There is a fuel rod representing the low 

power region; a fuel rod representing those fuel assemblies under guide tubes and fuel 

rod representing fuel assemblies under open hole support columns and free-standing 

mixers, (these are generally at the same power). There are also two fuel rods, which are 

modeled, in the hot assembly. There is the hot assembly average rod, which determines 

the hot assembly fluid conditions, and the hot rod, which is the highest powered rod in 

the core. The hot rod has included all the engineering uncertainties which makes it the 

hottest rod. Therefore, there are a total five rods, which are represented in the COBRA 

TRAC core model.

1-33
The calculated temperature response for these rods is shown in Fig. 35. In this figure, Rod 1 is the hot rod, and has a relative power of 1.73 and a delta H value of 1.73; Rod 2 

is the hot assembly average rod, and it has a relative power of 1.66. These are the rods, which represent the inner region of the core, Rods 3 and 4; and the Rod 5 represents the fuel assemblies on the outside edge of the core. So Rod 5 would represent the majority of the rods, which have been once, or usually twice, burned. This figure shows the 

temperature response of the different fuel rods to the LOCA transient as calculated and indicates that the hot assembly, of course, has the highest fuel rod temperatures. Rods 3 and 4 which represent the majority of the fuel rods in the core itself, have lower 

temperatures, approximately 1400 0F, and then Rod 5 which represents the rods on the outside edge of the core have even much lower temperatures like 400 OF. 

Figure 36 shows the corresponding fuel rod internal pressures, in psi, for each of the rods. In this particular calculation the hot rod burst was locked out, that is, the calculation prevented the hot rod from bursting even though it would have been calculated to burst. This was done to obtain a more conservative estimate of the hot rod PCT. The time at which the hot rod would have burst (Rod 1 in Fig. 35), can be obtained form examination of the hot assembly average rod, Rod 2 in Fig. 35. Rod 2, which is the hot assembly average rod is calculated to burst. Rod 1 would have been calculated to burst almost at the same time, perhaps a little earlier than Rod 2. Both 

Rods 1 and 2 would have burst in this calculation. 

Rods 3 and 4, which represent the majority of the rods in the core are calculated to not 

burst. Rods 3 and 4 are mostly once-burned rods. Some of them could be a little more than once-burned rods. Rod 5, which represents the low power assemblies on the edge of the core, which have the highest internal gas pressure are also not calculated to burst. 

Now, compare the calculated temperatures and internal pressures from Figs. 35 and 36 

to the burst criteria in NUREG-0630 shown in Fig. 37. Rods 3 and 4 lie outside the 

amount of burst data that is given so these rods would not be calculated to burst. Rod 5 is at such a low temperature that even though it has a high internal pressure it is not 

calculated to burst. Rods 1 and 2 would be calculated to burst, not because they have a 

high pressure but because they are at such a high temperature. 

To make Rods 3 and 4 burst, they would have to have the internal pressure increase 

from 600 psi to approximately 1000 psi, such that their points would lie on the NUREG

0630 burst curve. So those rods would have to have much higher pressure to be calculated to burst. However, these rods are not at such a high pressure and hence they are not calculated to burst in the calculation. In addition to the higher internal pressure, the fuel rod must also achieve high temperatures to burst. For a very high burnup rod, calculated temperatures in excess of 1200 degrees Fahrenheit are required for burst. Both the internal pressure and cladding temperature must be met for the fuel rod to be calculated to burst. As a rule of thumb, a high burnup rod, which would tend to be at relatively low power, would have to heat up to 1300 F (700 C) in order to burst.

1-34
Rod 1 - Hot rod, relative power - 1.73 F1 

Rod 2- Hot assembly rod, relative power - 1.66 

Rod 3 - Average rod under guide tubes, relative power = 1.2 Rod 4 - Average rod under other structures, relative power -1.2 

Rod 5 - Peripheral assembly rod, relative power = 0.3

2000

1600

j'1200

CL 

Seo E8 - -- - - - - - - N .... .

50 100 150 200 250 300

Time (seconds)

Figure 3; Best Estimate Analysis of the LBLOCA " 

(Clad Temperature Comparison at Rod_1 PCT Location)

-------------- --- --- r-------------- ----- 

- - - - - - - - -I - - - - - - - - -

400-

0

0

( ,I CA

j
.r - -- -- -- --- - -- ---r - -- --- -

IRod-1 ig0•'Rol

Rod 1 - Hot rod, relative power - 1.73 

Rod 2 - Hot assembly rod, relative power - 1.66 Rod 3 - Average rod under guide tubes, relative power = 1.2 Rod 4 - Average rod under other structures, relative power - 1.2 Rod :5 -Peripheral assembly rod, relative power-- 0.3

------ . . . . . --_ •,?-X.L -.. -• -_• ' • - - --r - -- - - -'-- -- - - - -- - -- - --- --- --- __ _- -

SI 

* I 

I * I II,, 

* I 

* I 

- - - - - - - - - - - - - - - - - - - - - ---- -- I 

-- - - - - - --- 

* I 

n I" To

I 

'r' 

-~ -------------- - - - - -

50 100 150

Time (seconds) 

Figure Best Estimate Analysis of the LBLOCA. 

(Rod Internal Pressure Comparison)

1000 

800 A

CL 

0. 

EL )

600

400 2

200

0O

U 250 300

'ý I a6le

- - - - - - - - - - - - - - - -- - - - - - - - - - - - - - - - -... . . - - - - - - - - - - - - - - ---

200
n

o 

0 o . . 

0

"3 

C)rf,

C:_ A 

m' CC)j 

0 5 10 15 20 25 

ENGINEERING HOOP STRESS-(KPSI) 

ORNL correlation of rupture temperature, as a f~nction of enqineerinci hoop stress and 

*temperature-ramp rate with data from internally heated Zircaloy claddinci in aqueous Satmospheres. 

In a LOCA the fuel rod internal pressure decreases because the energy source has been removed as the reactor shuts down. Also, all system pressures come down and the 

fission gas plenum is at a much lower temperature so the internal pressure of the rod 

will decrease. This effect is shown in Fig. 38 for different fuel rods at the beginning of 

life (BOL), middle of life (MOL), and the end of life (EOL). This internal pressure 

calculation assumes perfect axial pressure transport such that there is no internal 

resistance to the gas pressure within the fuel rod. If the temperature for Rod 5 is 

examined, which is like about 400 degrees, this point is not even on this burst curve.

1-37
Rod Internal 

PGAS 

- - - - PGAS 

-- - - PGAS

Pressure: 

2 0 a 

2 0 0 

2 0 0

BOL . ROD INT 

ROD INT 

ROD I NT

MOL. EOL 

GAS PRESSURE 

GAS PRESSURE 

GAS PRESSURE

2500 

2000

150

TIME (seconds)

1-38

I 

I' 

'I 

' A 

\ 

I \ 

-- - I -H

0 

C,, 1500

a__ 

h. 

0 0 

L.5 

500 

0

U 3U 1 U 200 250 300
This analysis indicates that at end of life, for the most burned fuel assemblies, which 

have the highest internal pressure, the rods are not likely to burst because their 

calculated temperatures are too low. If Rods 3 and 4, which constitute the majority of 

the rods in the core, don't burst, the majority of these fuel assemblies, which are going 

to have powers around 1.2 or lower, will not burst. Therefore it is only going to be the 

very high power fuel assemblies that burst, and these fuel assemblies are going to be the 

fuel assemblies with the least burnup. 

WCOBRA TRAC was used for the transient calculations. The variations of the thermal 

conductivity of the fuel with burnup has not been considered since the majority of the 

fuel rods would seem to be at relatively modest bumup prior to where thermal 

conductivity degradation with burn up would become important. 

1-1.7. Conclusions 

The bottom line is that for the highest bumup fuel assemblies, particularly on the edge 

of the reactor core, the assembly power and the corresponding calculated temperatures, 

for a LOCA, are sufficiently low that fuel rod burst is not calculate even if these 

assemblies have the highest internal gas pressure. If the loading pattern places a twice

burned assembly at the center of the core, surrounded by high power assemblies; the 

effect of the higher power assemblies is to drive the burned assembly to higher powers 

than would otherwise be the case. In this situation, with the higher internal gas 

pressure within the fuel rods, the high burnup fuel assembly could experience burst. 

However, most of the burned assemblies, in which the relative power was 1.2 or less, 

were not calculated to fail. Therefore, in a realistic LOCA calculation, the majority of 

the higher burnup fuel assemblies are not calculated to fail. This result is dependent on 

the loading pattern, which is chosen. : For the purpose of satisfying the 10CFR50.46 

criteria, the WCOBRA/TRAC plant model only analyzes the five fuel rods as discussed 

above. There is no fuel rod census throughout the core to see how many rods burst. 

That is not part of the acceptance criteria, although, perhaps, it should be.

1-39
1-2. PWR Small-Break LOCA: Impact of High Burnup Fuel

This review was prepared for the PIRT panel by A. F. Gagnon of Westinghouse Electric 

Company and panel member Lawrence E. Hochreiter of The Pennsylvania State 

University. 

As part of the High Burnup Phenomena Identification and Ranking Table (PIRT) effort 

for the Loss of Coolant Accident, it was requested to put together a PWR small break 

LOCA tutorial for the PIRT panel. A loss of coolant accident is a failure of the reactor 

pressure boundary, either a crack, split, or a break in the main reactor piping or a 

failure of a connecting line, such that the break is large enough that the reactor system depressurizes. When the primary system depressurizes, safety injection and reactor 

scram signals will be activated. In Westinghouse plants, breaks larger than 

three-eighths of an inch, result in break flows which is lager than the normal make-up system capacity at full pressure, such that the reactor system depressurizes. Westinghouse plants have instrumentation lines that come out of the bottom of the 

reactor vessel and so one designs for failure of those instrumentation tubes, such that it 

becomes a leak, not a LOCA. The high head charging system is sized to have enough makeup flow capacity, that if one of the instrumentation tubes failed, which has 

happened, the safety injection signal does not occur. 

Therefore, small-break LOCAs are breaks, which are typically larger than three-eighths of an inch, since smaller leaks don't depressurize the reactor coolant system, and are 

called leaks. Typically SBLOCA have break areas that are larger than three-eighths of 

an inch and less than one square foot. The limiting small break is determined by the 

interplay between the power level, and the axial power shape that is used, the capacity of the high head safety injection system, and the setpoint of the cold leg accumulators. 

The analysis and results that will be shown are typically what is calculated using a 

conservative Appendix K safety analysis model. 

Typically, what's used in the Appendix K analysis is an artificial power shape, which is 

bounding for all expected operational power shapes. The bounding shape has the peak 

power location (peak kilowatts per foot) at the top of the core, because as the reactor 

vessel drains the core is uncovered from the top. As the two-phase mixture level drops below the peak power location, the fuel rod transitions from being cooled by nucleate 

boiling to being cooled by very low steam flow. As a result, the fuel rod temperature increases. The specific axial power shape is also chosen such that the integral of power is low at the bottom portion of the core. In this manner, only a small amount of steam is 

generated below the peak location, which then provides the steam flow for the cooling of the uncovered top portion of the core. With a high power at the top of the core and 

relatively low power at the bottom of the core, poor heat transfer and higher calculated 

peak cladding temperatures (PCTs) are calculated at the peak power location.

1-40
The break size is also important, because this is the leakage path from the primary 

system. The high-head safety injection performance is also very important, since this 

represents the make-up flow to the system. The accumulator set-point pressure is 

important since this determines if and when in the transient, the accumulators inject 

and refill the vessel. 

Parameters such as the high head safety injection are plant dependent, and they can be 

different for different plants, and different vendors. The accumulator set-point pressure 

is also different for different vendor's plants. Therefore, the limiting break is that break 

size that is large enough such that the safety injection system can't make up the flow, 

such that the core uncovers. However, the limiting break must also be small enough 

such that the plant does not depressurize rapidly to the accumulator set -point, since at 

the set point the accumulators will inject and refill the reactor vessel. Therefore, the 

limiting breaks are those in the range where the break flow rate is significantly larger 

than the high head safety injection system; but, at the same time, not so large that the 

system is immediately depressurized to the point that the accumulators inject. 

Figure 1 shows the types of safety systems for a typical PWR. The cold leg accumulators 

are passive tanks of cold borated water that are isolated from the primary system by 

check valves and when the primary system depressurizes below their set point 

pressure, water is injected into the reactor downcomer at a high flow rate. Figure 1 also 

shows the high pressure, intermediate pressure and low pressures safety injection 

pumped systems. Most plants have both high-pressure and low-pressure safety 

pumped systems. These systems can come on at a varying range of pressures. For 

some plants, the high head injection pressure can exceed the safety valve set-point, 

particularly for older plants. This is not necessarily a good idea, because if there is a 

spurious safety injection signal, the charging flow from the safety injection pump could 

result in a liquid solid system, which will result in opening the pressurizer safety 

valves. Most of the high-pressure safety injection systems inject starting at 1,800 PSI. 

The low-pressure safety injection systems start to inject at approximately 150 PSI. 

Reactor scram and safety injection typically occurs when the reactor coolant system 

depressurizes below 1,900 PSI. 

Analysis of small break LOCAs is really verifying the performance or capacity of the 

high-pressure safety injection system to mitigate the small break transient. Single failure 

is assumed such that one of the high-pressure safety injection pumps is lost. 

For most Westinghouse plants, the limiting break size is typically between two and four 

inch sizes. The smaller breaks are terminated by the high-pressure safety injection 

system, and the larger, small-breaks, are terminated due to the accumulator injection. 

So, it is this smaller range between two to four inches that's more limiting. For both 

Westinghouse and B&W plants have the accumulator set-point pressure at about 600 

PSI. This pressure limits the maximum size of the small break, because the larger the 

break, the more quickly reactor system depressurizes and then the accumulator comes 

on and refills the vessel to terminate the accident.

1-41
INSlmt CONIAIN~rN?

I

ACCUMNATOR 

"""- -- " -

1 CO11 LE1O U OT LE ii I I 

PUMP 

I

•OU 1. $.t 

CONIAIN14NT 

"lNJECTt STOlON ON TANK

Intermediate ý Pressure Pump

Figure 1. Emergency Core Cooling System (Schematic)
Combustion Engineering plants have lower accumulator set point pressures, typical 200 

PSI. It is my understanding is that they dropped the pressure down to 200 PSI basically 

to improve the plant large break LOCA performance since less accumulator water 

would be bypassed. By decreasing the accumulator setpoint pressure, there is a larger 

range of small breaks, which can be more limiting. For the Combustion Engineering 

plants, the small break peak cladding temperature (PCTs) are typically higher than for 

Westinghouse and Babcock and Wilcox plants and are closer to the large-break LOCA 

PCTs. 

A Westinghouse reactor system is shown in Fig. 2. For small break LOCA behavior, a 

reactor trip on low pressurizer pressure will occur at typically 1,900 PSI. Full credit is 

taken for control rod insertion, such that the high pressure safety injection system only 

has to remove the core decay heat and the stored energy that is in the primary system, 

mostly in the metal structures.

REACTOR COOLANT PUMP

NUCLEAR REACTOR VESSEL

1-43
High-pressure safety injection system will occur about 1,800 PSI. The steam generators are isolated, and the emergency steam generator feed flow starts. Energy is removed from the primary system by both the break and the energy removal from the steam generators. The time period in which energy is removed by the steam generators depends on the break size, but usually for the limiting breaks it is for several hundred seconds. The primary system is in single and two-phase natural circulation with the 

energy being removed from the generators, in addition to energy going out the break. The plant decay heat is typically two or three percent and the generators are sized for 100 percent so it is easily remove the core decay energy from the generators, which will 

help cool the system down. 

Single phase and then two-phase natural circulation will continue until the reactor system has drained sufficiently that there will be a break in natural circulation in the generators because the mixture level is low enough that liquid can no longer carry over the U-tubes in the generators. At this point the plant will transition into a reflux condensation mode in the generators. Vapor with, some liquid droplets, will be carried into the up-hill side of the generator where the steam will be condensed and the condensate will run back down into the upper plenum. Some of the steam will also be carried over the U-tube to the downhill side of the generator where it is condensed and the condensate flows into the cold leg. During all phases of natural circulation, the high pressure safety injection system continues to inject subcooled borated water into either 

the cold legs or into the downcomer. 

As the transient progresses, the primary system slowly looses mass and the mixture level in the vessel is decreasing until there is a level in the hot legs and in the cold legs, natural circulation has ended and basically, the reactor vessel is acting like a boiling pot in reflux condensation. The core is still covered, and is in nucleate boiling. There is really no temperature excursion at this point, and the stored energy in the fuel has been removed. The energy that is being removed is the core decay heat any stored energy that is in the structures. 

As the mixture drains down further to the cold-leg elevation, the generators are empty, there can be a liquid level in the reactor cold-leg loop seals as seen in Fig. 2. The loop seals can be full of water and there will be a mixture level that may be in the upper plenum above the top of the core. For a cold leg break the loop seals will have to be vented such that core generated steam can vent out the break. To vent the loop seals, the upper plenum pressure will increase which will depress the mixture level in the vessel. Water will actually be forced out of the core, up into the down 

comer, to counterbalance the elevation head of the loop seals. This can lead to a core 

uncovery and a short temperature excursion. Once the loop seal vents, steam will then blow through the loop seal and out the break, and the mixture level will return to its original position in the vessel. Once steam vents out the break, the system depressurizes faster. Core uncovery can occur as the vessel mixture level is depressed to vent out the loop seal, however, it is only momentarily. It doesn't normally take very

1-44
long to vent the loop seal and then recover the mixture level, after which the system 

slowly continues to lose inventory. 

The portion of the core that is covered with a two-phase mixture remains in nucleate 

boiling. When the mixture level drops down into the core, there is a very sharp 

interface that exists between the covered region in the core and that portion of the core, 

which is uncovered. The covered region is in nucleate boiling where the cladding 

temperature is approximately five degrees above the saturation temperature. The 

uncovered region of the core is in steam cooling, and is the region in which the peak 

cladding temperature is calculated to occur. The covered and uncovered regions of the 

core are shown in Fig. 3. 

Depending upon how deep the core uncovers, the axial power shape that is assumed, 

and the core power level, the steam above the mixture level will quickly superheat. As 

a result there is very poor cooling, and the fuel rods heat up very quickly to high 

temperatures due to the decay power at the hot spot. From a fuel rod point of view, is 

the difference between the small and the large break, at this point is that in the small 

break, the stored heat in the fuel rod is not a factor and does not contribute to the fuel 

rod heat-up. In the large break, the fuel rod stored energy is a factor and does 

contribute to the fuel rod PCT. 

The resulting small-break PCT does dependent on how long the heat-up occurs in the 

transient and the location of the mixture level in the core, or the location of where the 

front is between nucleate boiling and steam cooling. 

The Westinghouse Appendix K small-break analysis method uses the NOTRUMP 

reactor systems code. It is a one-dimensional, network flow path, control volume type 

code that uses a five-equation drift flux formulation. NOTRUMP calculates the 

draining of the reactor system, counterflow and countercurrent flooding, and mixture 

levels in the core and different volumes in the reactor system. It treats thermodynamic 

non-equilibrium, since the steam can be superheated at temperatures much higher than 

the saturation temperature, in the presence of liquid droplets. 

NOTRUMP has flow regime dependent drift flux models to predict the mixture levels, 

two-phase pressure drop and the void fraction distribution in the mixture within the 

reactor vessel. The code can also predict the two-phase stratification in the hot legs and 

the cold legs as the vessel drains. NOTRUMP models the core and the structural energy 

release into the fluid and the steam generators as well as the break from the primary 

system. It uses point kinetics for the neutronics modeling but this is a small effect 

because the reactor shuts down as the control rod scram and it is really the decay power 

that is being removed. NOTRUMP also models the safety injection system including 

the high-pressure safety injection as well as the accumulators.

1-45
Figure 3. Cold Leg Break Steam Flow Path Schematic Drawing
Again, NOTRUMP is the hydraulic calculation, as seen in Fig. 4, and it passes the hydraulic 

information, such as the mixture level, the core flow, the pressure, and so forth, to a hot rod heat 

up calculation, which is the LOCTA code which is also seen in Fig. 4. LOCTA models the core 

average rod as well as the hot rod in the core. It has the thermal model for the rod as well as the 

mechanical model. The mechanical models is not the detailed mechanical design code model, but 

it's a suitable mechanical model that will predict the mechanical cladding behavior such as 

swelling, burst and rupture on the cladding as well as the peak cladding temperature. LOCTA 

also performs its own calculation of the local vapor temperature, which is the heat sink for the 

fuel rod model. The hot rod model in the LOCTA code has all the uncertainties added on such 

that the calculation represents the peak FQ, F delta H, and so forth. For the mechanical 

calculation, it includes the NUREG 0630, burst criteria and burst strain curve and includes the 

Appendix K required metal-water reaction and performs a double-sided reaction calculation if 

rod burst occurs. 

DECREASE IN REACTOR 

COOLANT INVENTORY 

N CORE PRESSURE, CORE 

0 FLOW, MIXTURE LEVEL, L 

T AND FUEL ROD POWER 0 

R HISTORY C 

T 

UA 

M 0<TIME<CORE COVERED A 

P 

CODE INTERFACE DESCRIPTION 

FOR- SMALL BREAK MODEL

1-47
Two sets of Appendix K small-break calculations, one calculation for fresh fuel which has higher peaking factors and fuel which has higher burnup at 30 GWd/t and 54 Gwd/t will now be examined. These calculations are for typical Westinghouse three-loop plant. Different plants will respond differently to the small-break LOCA, because of the sizing of the high-head safety injection system. Typically, Westinghouse four-loop plants have higher high head safety injection capacity so the small break peak temperatures that are calculated for these plants are much lower than the temperatures calculated for three-loop plants. 

The first set of calculations reflect typical FSAR small break LOCA calculations for fresh fuel. The initial conditions for the calculations are given in Table 1. Fresh fuel maximizes the peak linear heat rate in the calculation. However, at the same time, fresh fuel has very low burnup and the calculation tends not to predict fuel rod swelling, burst and ballooning, and failure of the cladding. The second set of calculations 

consider the effects of burnup on the fuel rod behavior and response. In these cases, the fuel rods do balloon and burst. The information in Table 1 is for a three-loop plant at the beginning of a cycle. The peak linear heat rate includes the effects of the calorimetric 

error or uncertainty. The plant power is 2775 megawatts. It has a rather high total peaking factor (FQ). This plant has fuel assemblies with a 17 by 17-rod array. The 

reactor system conditions are given in the table along with the plant initial conditions 

before the calculated transient. The accumulator set point pressure is 615 PSI, and it is 

assumed that the plant has1O percent steam generator tube plugging. 

Table 1. 

INPUT PARAMETERS USED IN THE ECCS SMALL BREAK ANALYSIS 

Core Power 2775 MWt 

Peak Linear Power (included 102% factor) 13.663 KWt/ft 

Total Peaking Factor, FO 2.46 

Axial Peaking Factor, FNz 1.4302 

Power Shape See Figure 15.6.5.2-3 

Fuel Assembly Array 17x17 

Accumulator Water Volume (nominal) 1,000 ft3 

Accumulator Tank Volume (nominal) 1,450 ft3 

Accumulator Gas Pressure 615 psia Safety Injection Pumped flow See Figures 15.6.5.2-1 and 15.6.5.2-2 

Initial Loop Flow 9,612 lb/sec 

Vessel Inlet Temperature 554.8 OF 

Vessel Outlet Temperature 622.2 OF 

Reactor Coolant Pressure 2300 psia Steam Pressure 925.5 psia 

Steam Generator Tube Plugging Level 10%

1-48
Table 2 gives the time sequence for a range of small breaks. In the analysis, a range of 

breaks need to be considered to examine the trade-off between performance of the high 

head safety injection system; the break size, the core power, and the accumulator set 

point. The interactions of all these parameters determine the most limiting break. 

Typically four breaks are examined; two, three, four, and six-inch breaks as shown in 

the table. The reactor trips on low pressurizer pressure. The S, or safety injection signal, 

quickly follows. Safety injection starts being delivered. In some plants, there is a delay 

time for safety injection. The time at which core uncover occurs is also given in the 

table. For the smallest break, it takes almost 1,500 seconds before core uncovery occurs. 

The larger the break the shorter the time scale for the accident transient. Also, the 

larger the break the quicker the reactor system will depressurize and the accumulator 

will terminate the transient because it will start to inject sooner. 

The time when the peak cladding temperature occurs is also given in Table 2. For the 

two-inch break, there is no accumulator injection. In other words, as the decay heat 

falls off, and the high head safety injection is maintained, it will eventually overcome 

the loss of inventory in the system and recover the vessel. 

The three-inch break has the highest PCT. This break requires the accumulator injection 

to terminate the transient and as the break size gets larger, the accumulator injection 

occurs earlier in time and the core is recovered earlier in time such that the resulting PCT is lower. The same criteria are used to evaluate the small-break LOCA as are used 

to evaluate the large-break LOCA. Usually the most limiting criteria is the calculated 

peak clad temperature which is limited to 2200 0 F. The small-break LOCA transients 

are very long and for the smaller breaks it take several thousand seconds before the core 

is completely covered by a two-phase mixture and the fuel rods return to nucleate 

boiling. 

While the cladding temperature is elevated, as long as it stays below the metal-water 

reaction temperature, the amount of oxidation of the cladding is small and the 17% 

criteria of Appendix K is not violated. If the cladding temperature does remain above 

1800 0 F for extended periods of time then the analysis would be limited by the cladding 

oxidation. The peak cladding temperature is still usually the most limiting criteria for 

the transient. 

The results of the different break size calculations are given in Table 3 and indicate that 

the three-inch break is the most limiting and its peak clad temperature is slightly over 

1800 0F. Again, this analysis has been performed for fresh fuel, so the fuel rods do not 

burst. The PCT elevation is at 11.75 feet which is due to the assumptions on the axial 

power distribution.. The maximum calculated oxidation is about 12 percent for the 

three-inch break. Therefore, the calculation has met the Appendix K criteria of a PCT 

is less than 2,200 degrees, oxidation is less than 17 percent, and less than one percent of 

core wide hydrogen generation. Since the peak cladding temperature and oxidation 

limits are within the Appendix K requirements the core remains in a coolable geometry. 

The plant has also been designed for long-term cooling.

1-49
Table 2.

NOTRUMP TRANSIENT RESULTS

Event Time (sec) 2 Inch 3 Inch 4 Inch 6 Inch 

Break Initiation 0 0 0 0 

Reactor Trip Signal 42.5 16.2 9.41 6.07 

" S-Signal 56.8 28.0 20.3 13.9 

SI Delivered 83.8 55.0 47.3 40.6 

Loop Seal Clearing* 960 416 210 42.5 

Core Uncovery 1468 700 436 18-1.5 

Accumulator Injection N/A 1246 609 259.9 

RWST Empty Time 3910 3878 N/A N/A 

PCT Time 2822.5 1482 764.2 317.7 

Core Recovery** > 6000 > 5000 2828 382.2 

* Loop seal clearing is defined as break vapor flow > 1 lb/s 

** For the 2 and 3 inch cases, where core recovery is > TMAX, basis for transient termination can 

be concluded based on the following arguments: (1) The RCS system pressure is decreasing which will increase SI flow, (2) Total RCS system mass is increasing, (3) Core mixture level 

has begun to increase and is expected to continue for the remainder of the accident 

Table 3. 

BEGINNING OF LIFE (BOL) ROD HEATUP RESULTS 

2 Inch 3 inch 4 Inch 6 Inch 

PCT (OF) 1276 1829 1531 1418 

PCT Time (s) 2822.5 1482.0 764.2 317.7 

PCT Elevation Uft) 11.5 11.75 11.5 10.75 

Burst Time (s) N/A N/A N/A N/A 

Burst Elevation (ft) N/A N/A N/A N/A 

Max. Local ZrO7 (%) 0.23 3.98 0.49 0.16 

Max. Local ZrOz Elev (M) 11.5 11.75 11.5 11.0 

Core-Wide Avg. ZrO2 (%) . 0.03 0.53 0.08 0.03

1-50
Figure 5 shows the high head safety injection flow. The lower curve is the flow that's in 

the broken loop, and the higher curve is the flow that is in the two intact loops. As the 

figure indicates, as the system pressure drops, the flow rate increases as expected. In 

NOTRUMP, the intact loops are lumped, this is a three-loop plant, so the intact loop 

flow actually represents the flow that is in two loops. One of the things that is done for 

the ECCS system design relates to the orifices in the injection lines, which are designed 

it to minimize the amount of flow that is loss due to the break. Therefore, each line has 

roughly the same flow per loop, whether it discharges into a broken cold leg or into an 

intact loop. The flows are actually tested in the plant. 

Figure 6 shows the axial power distribution that is used in the small break calculations. 

The peak power occurs near the very top of the core resulting in a very conservative 

shape. The peak power location is occurring roughly at the 10-foot elevation, and 

most of the power is in the upper portion of the core with lower power in the bottom 

part of the core. Since the core uncovers from the top, the two-phase mixture level is 

moving down and uncovering the top portion of the core. If there is lower power 

generation at the bottom of the core, the amount of steam that is generated is smaller, so 

that the steam flow flowing upward to cool the top portion of the core is smaller. 

Therefore a power shape of this type is specifically chosen to penalize the calculation. 

Figure 7 shows the reactor system pressure. The pressure quickly drops from 2250 psia 

down to approximately the safety valve setpoint on the secondary side of the steam 

generators (approximately 1250 psia). The main reactor coolant pumps are tripped very 

early in the transient. Energy is then removed from the primary system in this time 

period, by natural circulation, using the steam generators. As the primary system 

loosing more and more inventory and the pressures starts to come down. It is the break 

that's doing the depressurization. At approximately 1,100 seconds or so, the 

accumulator setpoint pressure is reached and the accumulators can inject which is a 

large source of water to refill the vessel. 

Figure 8 shows the mixture level in the vessel. The top of the core is at 22 feet in the 

figure. The mixture level decreases and then it starts to move into the core at 

approximately about 700 seconds or so, and reaches a level of approximately six feet in 

the core and then it slowly recovers. The accumulators start to come on at this point 

and then the core mixture level is recovered as shown in the figure. In this calculation, it 

does take a long time to completely refill the core.

1-51
-I______________

MNPS 1&2 FSAR

DECREASE IN REACTOR 

COOLANT INVENTORY

0 5 10 15 2D 25 30 35 40' 45 50 

SAFEIYINPtIQN FLO"S OQRISEUNVS

Figure 5. 

SMALL BREAK SAFETY INJECION FLOW RATES 

BREAKS LESS THAN 6 INCHES

1-52

Io 

z
MNPS 1&2 FSAR

DECREASE IN REACTOR 

COOLANT INVENTORY

2 4 6 8 10 

CORE LOCATION (FO) 

Figure 6. 

SMALL BREAK HOT ROD POWER SHAPE

1-53

12 

10

12
DECREASE IN REACTOR 

COOLANT INVENTORY

Figure 7. 

3-INCH PRESSURIZER PRESSURE

1-54

2500

2000 

a- 15 00 

=3 L. 

C" 1000 

c5 

500 

0-

0
"- 32

30 

-- 28 

26 

S24

22J 

S20 

18 

16

0 1000 2000 3000 

Time (s) 

Figure 8. 

3-INCH CORE MIXTURE LEVEL

1-55
Figure 9 shows the calculated peak cladding temperature, which indicates when the top of the core becomes uncovered. Once the core uncovers, the temperature immediately increases above the saturation temperature since the fuel is now being cooled by weak 

steam cooling. The calculated peak temperature reaches approximately 1,800 degrees before it turns-over and comes back down. In this calculation, the peak clad 

temperature is turning over as the vessel refills. As the core begins to refill two effects 

occur which help cool the peak temperature. Refilling the core reduces the vapor temperature, which is the heat sink for the heat transfer from the fuel rod. The second 

effect is that more of the core becomes covered with a two-phase mixture such that a 

higher steaming rate will occur which improves the single-phase steam convection heat 

transfer coefficient. Also, the figures indicate that the peak temperature is occurring 

over several hundred seconds during which the core is slowly refilling. During this time the decay heat is also decreasing slightly as time progresses. All of these small 

changes result in improved cooling at the peak location and the termination of the clad 

temperature rise. 

Figure 10 shows the calculated heat transfer at the PCT location. The high heat transfer 

is when the core is covered with a two-phase mixture while the low heat transfer 

coefficient is when the PCT location is in steam cooling. There are two orders of 

magnitude difference. 

Figure 11 shows the calculated vapor temperature at the PCT location. The steam does 

get quite superheated and it is not significantly different than the cladding temperature. The Steam temperature then decreases as the core refills, which decreases the cladding temperature. The heat flux at the hot spot sees the benefit of the increased steam flow, which results in a higher heat transfer coefficient, and the benefit of a larger delta T, both which increases the heat flux from the rod which turns the temperature over. 

Westinghouse also ran some additional 3-inch small-break LOCA calculations to 

examine the effects of high burn up situations. The hydraulic transient for the high burn up calculations and the beginning of cycle calculations shown previously, is nearly identical as seen in Table 4. The higher burn up rods have reduced power to reflect the 

fuel burn down that has occurred. These rods also have higher internal pressures since 

there are additional fission gases within the rods due to the burn up effects. 

Figure 12 shows the pressure behavior in the vessel, which is basically, the same, since 

the pressure is responding to the break and the safety injection systems. The 

accumulator set point is 615 PSI. The vessel mixture level shown in Fig. 13 is also the 

same. Twenty-two feet is the top of the core. As the figure indicates, the core becomes 

uncovered and then it slowly recovers due to the accumulator injection. Figure 14 

shows the core average vapor temperature, which reaches 1,300 degrees Fahrenheit. 

The vapor temperature increases once the core location is uncovered. The same axial 

power shape is used for these calculations as was shown previously.

1-56
DECREASE IN REACTOR 

COOLANT INVENTORY

Figure 9. 

3-INCH PEAK CLAD TEMPERATURE

1-57

2000 

1800 

1600 

1400 

2 1200 

o. 1000 

E 

800 

600 

400 -
DECREASE IN REACTOR 

COOLANT INVENTORY

DECREASE IN REACTOR 

COOLANT INVENTORY

0 1000 2000 3000 4nnn Ann

T ime ( s )

Figure 11. 

3-INCH HOT SPOT FLUID TEMPERATURE

Figure 10. 

3-INCH HOT SPOT HEAT TRANSFER COEFFICIENT

1800

1600 

1400

00

S1200 

S1000 

E 

I-- 800

5 

10 

C14 

1 0 

0 

2 

10 

S10 

I10 

0D 

10w.

600 

400

•v•u I'

Pl
3 - Inch Cold Leg Break

- RCP Trip 

- S - Signal 

- SI - Delivery

- Loop seal clearing 

- Core Uncovery

16.8 sec 

30 sec 

64.5 sec 

349.5 sec 

470.6 sec

- Accumulator Injection 1036.5 sec

- Core Recovery 3333.3 sec

Figure 12. 

2500 

2000 . ...... 

1.) -2 1500 ...

a) 

U) 

U) 

a) 

C1000

500 

0

Transient 

9

Run 

0

3 inch Break w/High Tavg 

0 PRESSURIZER

Time (s) 

1-59
Figure 13. Transient Run 3 inch Break w/High Tavg

I U 0 C0RE

Time (s)

Figure 14. 

1400 A

1200-

1000-4

800 

600 -

Transient Run 3 inch Break w/High Tavg

6 0 0 TOP CORE VAPOR

2000 

Time (s)

1-60

35

S30 

S25 

=> 20 

15

3--

400 I 1000

I I I I It I I I I

b 3000 4000 0
Figure 15 shows the rod temperatures and the local hot channel vapor temperatures. 

This particular rod has 30 thousand-megawatt days per tonne burn up. The hot channel 

vapor temperature is higher than the core average vapor temperature shown in Fig. 14 

as expected since the hot channel enthalpy rise is larger than the core average enthalpy 

rise. Figure 16 shows the fuel rod internal pressure and the system pressure. There is an 

internal pressure increase, because the temperature of the rod is increasing, and then 

the pressure comes down, due to clad swelling and then burst occurs at approximately 

1100 seconds and the rod internal pressure decreases to the system pressure. The 

differential pressure at burst or the delta P at burst is shown on the figure. Now, this 

point can be plotted on the NUREG-0630 burst curve, given in Fig. 17, which is what's 

used in calculating the burst. 

The calculated oxidation level can be calculated for the 30 thousand-megawatt day per 

tonne burn up fuel rod in the calculation. The oxidation level at the PCT location, 

without burst, in this calculation is around 4.1 percent as seen in Fig. 18. These were 

initially about 3.95% of initial oxidation before the transient so there was an increase in 

the oxidation that occurred due to the transient. However, since the calculated PCT is 

only around 1,600 degrees or a little bit less, so there is not a significant amount of 

oxidation generated during the transient. At the burst location, one has to calculate 

double sided oxidation as per the Appendix K requirements. As seen in Fig. 19, the fuel 

rod had initially roughly 3.9 percent oxidation, but total oxidation at the end of the 

transient is only about 5.3 percent. So because of the double-sided oxidation at the 

burst location, additional oxidation is calculated for the 30 thousand-megawatt days per 

tonne bum up rod. The total amount of oxidation is higher, but still within 17 percent 

requirement. 

The results for a 54,000-megawatt day per tonne rod, for a 3-inch break are very 

similar. The small-break hydraulics are the same so only the hot rod behavior will be 

examined. Figure 20 shows the cladding temperature, which is increasing from the 

point of core uncovery and then turns-over as the core, is recovered. The hot channel 

vapor temperature is also plotted on Fig. 20. The core average vapor temperature peak 

was about 1,300 degrees Fahrenheit. The fuel rod temperature is lower than in the two 

previous cases since the rod power is lower due to the burn up. In this case the peak 

temperature was only 1500 'F so not much oxidation will occur. Since the burn up is 

now 54,000 megawatt days per tonne, the rod pressure is higher as seen in Fig. 21. The 

rod pressure increases as the rod temperature increases and there is some swelling 

which occurs, which brings it down, before it bursts. The rod then bursts with the 

differential pressure as seen on the figure. As Dr. Meyer indicated, the PIRT panel 

should think about the high burn up mechanical properties that could have an effect on 

this process, because the ballooning strains might be much smaller. The pressure would 

not drop as seen in the figure and burst could occur earlier. It might occur at a slightly 

different set of pressure and temperature conditions. This could have some effects in 

this area, since the NUREG-0630 curves that are used are all for fresh fuel.

1-61
HIGH BURNUP 

sblocta

STUDY 15x15 Fuel 

20.0 FP 

5:43:59.40 1858872525

Non-IFBA 30K Burnup 

C1999/01/25 X2000/07/20 

lioness

Hot Rod PCT 

Channel Fluid Temperature

1800 

1600

-i, 1400

Time (s)

Figure 15

ON

Q) 

ED 

qP 

c:)_ 

t---

1200 

1000 

800

600 

400
HIGH 

sblocta 

2000 

1800 

1600 

.o 1400 cn 

1200 

= 1000 

C!) 

co 

® 800 

600 -• 

400 -. 

200

0

BURNUP STUDY 15x15 Fuel 

20.0 FP 

15:43:59.40 1858872525 

System Pressure 

Rod Internal Pressure (Hot

2000 

Time (s) 

Figure 16.

Non-IFBA 30K Burnup C1999/01/25 X2000/07/20 

lioness 

Rod)

4000
0 

0

3 I " i .20 c/s. 

O_ 13 a O /.• 14 C,/S" 

0~ 

C', 

0 iO 20 25 

ENGINEERING HOOP STRESS .(KPSI) 

ORNL correlation of rupture temperature. as a function of enqineerinq hoop stress and Figure 17. temperature-ramp rate with data from internally heated Zircaloy claddina in aqueous 

atmospheres.
HIGH BURNUP STUDY 15x15 Fuel Non-IFBA 30K Burnup 

sblocta 20.0 FP C1999/01/25 X2000/07/20 

15:43:59.40 1858872525 lioness

__0. Oxide

4.15 

4.1

4.05

4

Thickness

3.95

At PCT Elevat ion

2000 Time (s)

Figure 18

U) Ct) 

C-

-

0
- HIGH BURNUP STUDY 15x15 

blocto 20.0 

15:43:59.40 1858872525 

Maximum Oxide Thicknes 

5.4 

5.2--__ y 

n 4.8

.) 4.6

) 4.4

S4.2

4-

3.8

Fuel Non-IFBA 30K Burnup 

FP C1999/01/25 X2000/07/20 

lioness

Time (s)

Figure 19.

1-66

t-" 

C.-

0
HIGH BURNUP STUDY 15x15 Fuel 

sblocta 20.0 FP 

10:30:19.30 1625140151

Hot Ro 

Channe

d 

I

PCT 

Fluid T emp e r a t u r e

Time (s) 

Figure 20.

Non-IFBA 54K Burnup C1999/o1/25 X2000/07/20 

calico

CL 

0-_) 

E 

(1) 

I---

1600 

1400 

1200 

1000 

800 

600 

400
HIGH BURNUP STUDY 15x15 Fuel Non-IFBA 54K Burnup 

La 20.0 FP C1999/01/25 X2000/07/20

10:30:19.30 1625140151

System Pressure 

Rod Internal Pressure

2500 

2000

Time (s)

Figure 21.

sbloc

calico

(Hot Rod)

C/) 

0

00 1500 

1000 Cl) 

U) 

ci) 

C500 

0

4000
The calculated oxidation for the 54,000 megawatt days burn up case at the PCT location 

is shown in Fig. 22. The initial oxidation on the rod at the beginning of the calculation, 

due to the higher burn up is roughly 13 percent. The increase in the amount of 

oxidation at the PCT point is not very much due to the lower temperatures during the 

transient. The oxidation at the Burst node is shown in Fig. 23 and again there is a very small, gain in the total oxidation even though the double-sided oxidation was 

considered. In these cases, the accident does not really add any significant oxidation, as 

compared to the initial oxidation level on the fuel rod at the beginning of the accident 

due to the increased burn up history. Therefore, as the fuel rod bum up increases, the 

initial oxidation value at the beginning of the transient increases, however, with the 

higher burn up, the rod power are sufficiently low that there is no significant heat up of 

the highly burn up rods and the additional oxidation due to the transient is very small. 

Therefore, the Westinghouse small-break Appendix K calculations show that the 

calculated PCTs are well below the 2,200 degrees Fahrenheit limit for all burn ups considered. Fresh fuel, gives the highest PCTs since the linear heat rate is the highest 

but the hot rod is not calculated to burst in these calculations. As the burn up increases, 

the calculated PCT will decrease since the linear heat rate is lower but the fuel rod will 

burst due to the increased fission gas pressure within the rod. The fuel rod burn up does not have to be very large before a burst is calculated. Burst usually will be 

calculated at burn ups greater than about 10,000-megawatt days per ton. 

If best-estimate small break calculations were performed, much lower PCTs would be 

calculated, primarily because of the lower decay heat. Lower decay heat helps in many different aspects. It will make the worst small break size larger, and since the 

accumulator setpoint is fixed, the window for which the core could uncover would 

become smaller. The transients are all going to be shorter, so the time at elevated 

temperatures will be less so the heat ups, for thousands of seconds, which was 

calculated in the Appendix K calculations, would not be calculated in the best-estimate 

calculations. 

Also, best-estimate calculations would use more realistic power shapes, instead of 

bounding power shape and the peak power and the PCT will not be at 11-1/2 feet. In 

some of these Appendix K calculations, bursting is predicted to occur at 12 feet which is 

unrealistic because there is no fuel at 12 feet or perhaps just one pellet, and so the burst 

occurs where the fission gas plenum is, which is sort of crazy. So, a realistic power shape is going to reduce the power at the top of the core. Since the power will be 

reduced at the top of the core, there will be more power at the bottom of the core, which 

will increase the steam flow at the PCT elevation and improve heat transfer there. 

Appendix K calculations use a conservative break flow model which maximizes the 

flow lost from the reactor system. If a best-estimate break flow model is used, the flow 

out of the system would be less, which will make the high head safety injection system 

perform more effective.

1-69
HIGH BURNUP STUDY 15x15 Fuel Non-IFBA 54K Burnup 

sblocta 20.0 FP C1999/01/25 X2000/07/20 

10:30.19.30 1625140151 calico 

Oxide Thickness At PCT Elevat ion 

13.234 

13.232 

• 13.23 . - . . . . . .. . . . . . . . . . 

co 13.228 . . . . . . . . . . . . . . . . . . . . . . . . . o .. . . . . 

.2 2 6 . . . . . . . . . . . . . . ... . . . . . . . . . . . . . . . . . . . . . . 

13 2 2 . . . . . . . . . . . . . 

13-2 . .- . . .. 

13.224 

o 13.222-' 

13.22

13.218 - ' I I i J . t i i I

Time (s)

Figure 22.
. - HIGH BURNUP STUDY 15x15 Fuel Non-IFBA sblocto 20.0 FP C1999/01 

10:30:19.30 1625140151 calico 

Maximum Oxide Thickness

54K Burnup /25 X2000/07/20

0 500 1000 1500 

Tim

20 

e

00 2500 3000 

( S)

Figure 23. 

1-71

14

13.9 

13.8

13.7 

13.6

U) 

Ur) 

Q) 

C-) 

C13.5 

13.4 

13.3 

13.2 41! I I - I I I I I I I I

3500 4000

.,I

tf

I ! I 1 I i I I I
There are also three-dimensional effects, which are not accounted for in the Appendix K 

calculations, which would result in a thermal siphon effect where the higher power regions will suck in steam from the lower power regions and a higher steam flow for 

the higher power fuel assemblies will be calculated. Also, for best-estimate calculations 

more realistic rod powers and fuel rod internal pressures would be used such that burst 

would probably not be calculated. So, if a true best estimate calculation were 

performed the expected PCTs would be nowhere near the licensing limit. In fact, it is 

likely that core would not uncover for some of the breaks that are now calculated to 

uncover with Appendix K calculations. The resulting PCTs would be much, much lower 

and none of the rods are likely to burst.

1-72
1-3 BWR LOCA

Panel member Jens G. M. Andersen prepared this review for the PIRT panel. 

The following is a presentation of the BWR LOCA. The presentation is based on the 

GE BWRs. Slide 2 is a brief outline of the presentation. The presentation will be based 

on the design basis accident, which is the double-sided break in the suction side of the 

recirculation line. It will cover the phenomena that occur during the LOCA. 

BWR LOCA %.= 

*Design Basis Accident 

- Double Sided Break in Recirculation Line - Suction Line 

SPhenomena 

- Blowdown 

- Refill Reflood 

* BWRP6 BWRl4 BWRI2 

*Typical LOCA Transient 

*Small Break LOCA 

*Fuel Exposure 

3GMt' MA1y .(W 2 

The discussion of the LOCA phenomena will be divided into two phases, the blowdown 

phase and the refill/reflood phase. The presentation will furthermore cover the 

differences between the different plant types, the BWR/6, which is the latest design in the 

U S., and the BWR/4 which is the design for which there are the most plants. Finally the 

BWR/2, which is the early BWR designed with the external recirculation loops without jet 

pumps, will be covered. There are two BWR/2 in the United States. In addition to the 

typical design basis LOCA transient, the presentation will address issues that exist for 

small break LOCAs, and it will discuss the effect of fuel exposure on the fuel performance 

during a LOCA transient.

I- 73
The BWR/6 design basis accident (See Slide 3) will be covered first. During normal 

operation, the feedwater comes in at the top of the downcomer just below the dryer skirt. 

The feedwater will flow down the downcomer into the bottom of the downcomer where 

the suction for the recirculation line is located. Water is taken out through the 

recirculation line, enters the recirculation pump and is then reintroduced back into the jet 

pump in the downcomer through the drive line for the jet pump. There the drive flow 

mixes with the suction flow that comes directly from the downcomer. For simplicity, the 

driveline is shown at the drawing coming in at the top of the jet pump. In reality, the 

drive line enters the downcomer at the same elevation as the suction line and then 

continues vertically to the top of the jet pump inside the downcomer. There is a 180 

degree bend at the top before the drive flow enters the jet pump. Both of the recirculation 

line connections are the bottom of the core, to avoid having any pipe connections on the 

vessel in the high flux region outside the core. 

BWR/6 LOCA GNF 

 jMM Design Basis Accdent 

j_ -dWW CDouble Ended Break In Reclrculalon 

Line - Suction Line 

LLPCI 

JIWOA May 2W 3 

The flow then exits the bottom of the jet pump and continues up to the core. 

The rectangular boxes in the core region on Slide 3 represent the fuel channels. The region 

in between the fuel channels is the bypass region. Typically, about 10-11% of the active 

recirculation flow flows in the bypass region. The dominant flow path for the bypass flow 

is up through the inlet orifice in the fuel channels, and then through the leakage paths 

which are small holes in the side of the nose pieces to the fuel channels.

1-74
The fluid flows from the inside of the channels through the leakage holes and couple of 

additional leakage paths into the bottom of the bypass region. The flow through the 

leakage holes is the dominant flow path into the bypass region. It is important, because it 

has a very significant impact on the performance during the refill/reflood phase of the 

LOCA. 

Boiling and vapor generation will take place in the channels, and the two-phase mixture 

flows from the core region through the upper plenum and into the steam separators. The 

steam is separated from the liquid in the steam separators and flows through the steam 

dryers into the steam line. The main flow path for the liquid is from the separators into 

the mixing region and back into the downcomer where it mixes with the feedwater flow. 

The inlet subcooling at the bottom of the core is approximately 20 F. 

The worst break, the design basis accident for the BWR/6, is a break in the suction side of 

the recirculation line. This is the lowest pipe connection for any large pipe relative to the 

position of the core. 

The height of the jet pumps is two-thirds of the height of the core. With a break in the 

recirculation line, it is possible to reflood the core with a collapsed water level up to the 

height of two-thirds the height of the core before the water will spill out to the top of the 

jet pump and flow out to the break location. 

There are other possible break locations - in the LPCI line or the spray lines, but those 

would be break locations above the top of the active fuel in the core, and the break sizes 

are smaller. There are also possibilities for break in the feedwater line and the steam lines. 

These other breaks are usually much less severe than the break in the recirculation line. 

The design basis accident is a double-ended break in the suction line (See Slide 4). There 

are two suction lines going into two recirculation pumps in the BWR. 

After the recirculation pumps, the flow continues to a header that feeds the drive lines for 

10-12 jet pumps per recirculation loop. The suction pipe is a large pipe. Following a 

break, the reactor scrams, typically on high drywell pressure. However, because of the 

depressurization, and the associated fluid loss there is a very fast increase in the core 

average void fraction. The negative reactivity caused by the voiding shuts down the core 

immediately. 

Immediately after the break, fluid will start flowing from the downcomer out through the 

suction line to the break. There will be flow reversal in the drive line and fluid will flow 

to the break from the other side. As a result of the flow reversal in the discharge side of 

the recirculation line, all the drive flow to half the jet pumps is lost.

1-75
e Break in Recirculation Line

e Scram 

IIIIIII steam Line ' Loss of Power 

9 Flow Reversal in Broken Loop 

Feed Water Jet Pump 

LPCS/HPCS * Coast Down of Intact Pump 

e Large Reduction in Core Flow 

and Early BT= 1 sec. 

. PCT Dominated by Stored 

Energy (kW/ft, hgap) 

D Isolation 

* Depressurization and Loss of 

Liquid Inventory 

JGMA May 2000 

The flow reversal in the drive lines and the break flow out the suction line causes the exit 

flow from half of the jet pumps to reverse. It is conservatively assumed that all power is 

lost, which means that there is no power to the recirculation pumps and the flow in the 

intact recirculation loop will start to coast down. The coast-down time of the recirculation 

pumps is on the order of 10-15 seconds, so it's a relatively slow coast-down. 

Consequently, there is a very large, almost instantaneous reduction in the core flow. This 

leads to an early boiling transition in the core, typically within one second after the break.

-- 76
Following the early boiling transition, there is very fast increase in the cladding 

temperature. During steady state, there is a parabolic temperature profile in the fuel. 

Following the boiling transition, the film boiling heat transfer is much less than the 

nucleate boiling heat transfer in normal operation, and therefore the temperature 

distribution in the fuel will become equalized. Therefore the early peak cladding 

temperature (PCT) is dominated by how much stored energy there initially is in the fuel, 

and that is given primarily by the linear heat generation rates and the fuel heat transfer 

parameters. These are the two dominant parameters that control the amount of stored 

energy and the early PCT. 

Subsequent to the break there is an isolation of the primary system. The closure of the 

isolation valves typically occurs four seconds into the transient. 

As a result of the break, the system starts depressurizing and there is a loss of liquid 

inventory. The loss of liquid inventory leads to a drop in the water level in the 

downcomer, and when water level has reached the top of the jet pump, a flow path is 

established where steam can flow directly out to the break (See Slide 5). This leads to a 

significant increase in volume flow through the break and in the rate of depressurization. 

BWR/6 LOCA 

* Increased Depressurizatlon 

Following Jet Pump Uncovery 

* Flashing when Tsat(P) < T, 

* Increased Core Flow from Lower Plenum Flashing 

I i I II -. Quenches Fuel

PCT .tGbM M.y 'MO -T

Core Flow
The initial subcooling in the downcomer and lower plenum is approximately 20 F, 

therefore when the pressure drops below the saturation pressure corresponding to the 

liquid temperature, flashing of the liquid will take place. The combination of the large 

increase in the depressurization following the jet pump uncovery and the flashing of 

super-heated liquid, causes a surge in core flow. The surge in core flow from the lower 

plenum quenches the fuel and the cladding temperature is returned to the saturation 

temperature. Lower plenum flashing typically occurs 10 seconds into the transient. 

In summary there is a rapid reduction in the core due to the flow reversal in half the jet 

pumps and a subsequent early boiling transition. Once the pressure drops to the point of 

lower plenum flashing there is a surge in core flow, which quenches the core and brings 

the temperatures back to the saturation. This is referred to as the early boiling transition 

and the first peak in the cladding temperature for the BWR. 

As the LOCA progresses, the liquid inventory and pressure continue to drop, and 

eventually the two-phase level drops inside the core region. A two-phase level will also 

form in the lower plenum region (See Slide 6). 

The reason that levels may form in both in the lower plenum and the core region is 

counter current flow limitation (CCFL) at the inlet to the fuel bundles. This core uncovery 

leads to a second boiling transition, which typically happens 20 seconds into the transient. 

BWR/6 LOCA 

* Continued Depressurlzation 

e Loss of Liquid Inventory 

S*Core Uncovery Leads to 

I• Second BT 

Fft Wow~. a HPCS comnes an (DIG Startup) 

* LPCS and LPCI Comes on 

when P < Shutoff Head 

*Worst Single Failure 

Falure: LPCt DIG 

- AvOMMabl HPCSLPCS&PCUAOS 

JG4A .4ay 2I00 6

T- 78
The high pressure core spray into the upper plenum will come on early in the transient. 

It is controlled by the time it takes to start up the diesel generator that powers the high 

pressure core spray pump. This pump typically comes on 35-40 seconds into the 

transient. 

Diesel generators also power the low-pressure core spray into the upper plenum and the 

low-pressure coolant injection into the top of the bypass region. These systems, however, 

will not come on until the system pressure has dropped below the shutoff head for the 

pumps, which is typically on the order of about 200 psi. 

A conservative requirement for the LOCA analysis is to make the assumption of the worst 

single failure. The BWR/6 has one high pressure core spray system, one low pressure core 

spray system, and it has three LPCI systems injecting into the bypass region. In addition 

to this, there is the automatic depressurization system ADS). The ADS is of importance 

primarily for small breaks and it's purpose is to lower the pressure and allow the low 

pressure systems to come on. The worst single failure for the BWR/6 is to assume a 

failure of one of the diesel generators powering two of the LPCI systems. 

Several different combinations of failures are analyzed, but the diesel generator failure is 

the worst and leaves only the high pressure core spray, the low pressure core spray and 

one LPCI system that injects into the bypass region. 

When the pressures is low enough that it has dropped below the shutoff head for the 

LPCI pump, liquid is injected into the bypass region and quickly fills up the bypass region 

(See Slide 7). The only significant way this liquid can drain out of the bypass is to flow 

backwards through the leakage hole into the fuel channels and then drain through the 

side entry orifice into the lower plenum. However, because of the depressurization and 

release of stored energy from the vessel wall and guide tubes in the lower plenum, the 

liquid in the lower plenum will be flashing. The flashing leads to a significant flow of 

steam up into the core region and at the same time there is liquid trying to drain into the 

lower plenum. For the BWRs, in order to control core flow distribution and stability, there 

are relatively tight inlet orifices at the bottom of the channels. The flow area at the inlet 

orifice is typically a third or less of the active flow area in the channel. Counter current 

flow limitation will occur at this point and liquid will be held up in the core. Two-phase 

levels are therefore formed both in the core region and in the lower plenum. 

In the upper plenum, there is injection of the high pressure and the low pressure core 

sprays. There is a flow restriction at the top of the fuel channel. In older fuel types, the 

most limiting flow area used to be the upper tie plate in the channel. For modern fuel 

designs, the flow area in the upper tie plate has been increase, and the limiting flow area 

is typically the top spacer in the fully rodded region of the bundle.

.- 79
BWR/6 LOCA 

9 LPCI Fills Bypass Region - Liquid Enters Core through Channel 

Leakage Path Steam Line - Steam Holds from Lower Plenum Flashing up Liquid in Core due to CCFL 

Feed Water in Inlet Orifice 

9 HPCS/LPCS Enters Upper 

LPCS/HPCS Plenum 

LPCI- Liquid Level Forms in Upper Plenum. 

Holdup in Upper Plenum due to 

Steam from Lower Plenum (flashing) 

and Core (heat transfer). 

CCFL at Upper Tie Plate 

- Down Flow through Peripheral Low -I +. . Power Bundles Helps Refill and 

Reflood Core 

JGMA MaY 2000 7 

Steam originating from the flashing of fluid in the lower plenum and core region plus steam generated in the core due to the heat transfer from the hot fuel will flow up through the top of the bundles. Counter current flow limitation can therefore occur at the spacers or the upper tie plate and will then hold up liquid in the upper plenum. CCFL typically exist for the average and high power bundles, while for the peripheral channels, which 

are at much lower power than the central channels there is typically not enough steam 

flow to cause CCFL and down-flow is likely. 

The phenomena in the upper plenum and core regions are of particular interest for refill/reflood part of the transient. In the upper plenum, there is interaction between the 

core spray system and the water level, and in the core region there can be three different flow patterns (See Slide 8). There are two different scenarios dependent on whether the 

water level in the upper plenum is below or above the core spray spargers. 

The cold ECC water is injected through the spargers, which are pipes that runs along the periphery of the upper plenum and contains spray nozzles along the length of the pipe.

S-- 80
BWRIB LOCA GNF 

"* Upper Plenum Phenomena 

" Level Below CS Sparger 

co- Gondens"on on ECG - Spyolstilbutln 

- CCFLalUpperTW PIa Holds up Liquid Jm Upper hMemn 

- Upper Plenum L&val Raises 

Os/H SO Level Above CS Sparger - Not Enough Stearn for Condensaticm 

LPCI a E - Sucoln buhe Petphra 

Bundles Caushng Subcooled CCFL 

Break Down 

- Upper Plan= Level Drops 

a Parallel Channel Effects in Core 

- Co:Cutm Up Flow 

- Cortier Currt Flow wft Level 

- Co Current Down Fow 

* Refloocing of Core Prior to 

Refilling of Lower Plenum 

JIGMA Wy 2000 8 

When the water level is below the core spray sparger, the nozzles spray water as droplets 

into the vapor space m upper plenum. The condensation heat transfer is very large on 

these droplets and data have shown that there usually is enough condensation to remove 

all the subcooling once the droplets have traveled a distance on the order of 4 to 5 nozzle 

diameters into the upper plenum. There is complete condensation on the ECC water, 

which means that the water is saturated once it reaches the top of the two-phase level. 

Therefore the water below the two-phase level in the upper plenum will be saturated. 

CCFL will occur at the top of the bundles due to the up-flow of steam, and the water can 

not drain into the core as fast as it is supplied from the spray sparger. Consequently the 

water level in the upper plenum will start rising. 

This quickly leads into the other situation where the two-phase level in the upper plenum 

is above the core spray sparger. In this scenario, the water essentially shields the cold 

water from the steam. Enough steam can not get to the cold ECC water to provide

J- 81
sufficient condensation to heat the water up to saturation. Therefore, the liquid below the 

two-phase level becomes subcooled, and cold water starts penetrating to the top of the 

core and in particular along the shortest path down into the peripheral bundles. 

These bundles are also the lowest power bundles, and subcooled CCFL breakdown will 

occur. Subcooled CCFL breakdown is a phenomenon that occurs when the amount of 

condensation required to remove the subcooling of the water flowing down exceeds the 

upward steam flow. When this condition occurs, the cold water penetrating into the 

channel condenses all the available steam flow and the flow pattern changes to co-current 

down-flow of cold water in these peripheral channels. 

As a result of the subcooled CCFL breakdown in the peripheral channels, the upper 

plenum starts to drain and the two-phase level drops until the core spray is uncovered 

again. The upper plenum enters a mode where the water level slowly oscillates up and 

down around the location of the ECC sparger, the oscillation occurs with the time period 

of maybe 10 to 20 seconds. Consequently during the refill-reflood period a water level is 

maintained and oscillating around the core spray spargers. 

In the channels three different flow patterns can occur. There can be down-flow in the 

peripheral channels because of the CCFL breakdown that occurs when the cold water gets down to the top of the channels. There can be a pattern of co-current up-flow where steam 

from the lower plenum and core is vented up through some of the channels at a rate close 

to or exceeding the shut-off point on the CCFL curve, i.e., the point where liquid down 

flow is not possible. This typically occurs in the highest power channels. For the 

remaining part of the channels, there will be a counter-current flow pattern. There is 

counter-current flow at the top of the channel where the upward flowing steam limits to 

the amount of liquid flowing down into the channels, and as a result a two-phase level is 

formed inside the channels. This flow pattern occurs for the majority of the channels. 

The number of channels in co-current up-flow, co-current down-flow, counter-current 

flow is controlled by several phenomena. One of them is the requirement to have the 

same pressure drop between the inlet and the exit of the channels, which is a pressure 

drop between the lower plenum and the upper plenum. The number of co-current 

up-flow channels is given by the number of channels it takes the vent flow that is 

generated from lower plenum flashing plus the vapor generation in the channels. The 

number of co-current down-flow channel is given by the number of channels it takes to 

drain the core spray flow that is injected into the upper plenum. The remaining channels 

will be in the counter-current flow mode. The two-phase levels in these channels are such 

that the pressure drop between the upper and lower plena is the same for all channels. 

This phenomenon was not seen in some of the earlier experiments in the seventies such as 

the Blow Down Heat Transfer (BDHT) loop, and then later on the Two Loop Test 

Apparatus (TLTA) and in the eighties the Full Integral Simulation Test (FIST) facilities 

under the Refill/Reflood and FIST programs. In those facilities, the simulation of the

.J7- 82
boiling water reactor was scaled to one single fuel bundle. The three simultaneous flow 

patterns were not possible as only one channel existed, and the major flow pattern seen in 

the experiment was the counter-current flow pattern. As a result the liquid in the test 

facility with a single bundle would initially drain slower into the lower plenum and the 

two-phase level in the lower plenum would drop down to the bottom of the jet pumps. 

When the level drops below the bottom of the jet pumps, steam that is generated from the 

lower plenum flashing will vent up to the jet pumps. This reduced the amount of steam 

flowing up through the core and as a result of that counter-current flow limitation would 

diminish and allow more liquid to drain down from the upper plenum and core regions 

into the lower plenum. During the transient while the level existed at the bottom of the jet 

pump, liquid would be entrained out through the top of the jet pumps and flow out 

through the break. 

During the refill-reflood program, however, tests were also conducted in the Steam Sector 

Test Facility (SSTF). In this test facility a full height 30 degree pie-shaped sector of the 

boiling water reactor was simulated. 58 parallel fuel channels were simulated in the SSTF 

and all three flow patterns were observed in this these tests. In these tests the level in the 

lower plenum did not drop down to the bottom of the jet pumps. Because of the parallel channel effects allowing all three flow patterns at the same time, there would be enough 

channels in down-flow to drain the flow out of the upper plenum while maintaining the 

level at the bottom of the core spray spargers. 

Furthermore, as a result of holding up water due to CCFL at the side entry orifice at the 

bottom of the fuel channels, the core in the BWR typically refloods before the lower 

plenum is completely filled. Consequently, in the PIRT tables, we do not distinguish 

between the refill and the reflood phase, but we consider the two together, since they 

occur simultaneously. Eventually the refilling and reflooding will restore the liquid level 

in the core and it will completely quench the core. Inside the jet pumps the water level 

rises to the top of the jet pumps, which is two-thirds height of the core. With the vapor 

generation in the core region and the corresponding high void fraction, the two-phase 

level inside the core shroud needed to balance the static head is significantly above the 

top of the active fuel, and it completely quenches the core (See Slide 9). 

To summarize the LOCA transient for the BWR/6, there is a second boiling transition that 

occurs, typically 20-35 second into the transient. A minor core heat up is experienced, and 

the core is reflooded and quenched 100-150 second into the transient. The PCT in a 

nominal best estimate calculation is below 1000 F. The upper bound PCT at the 95% 

probability / 95% confidence level is a couple hundred degrees higher and in the 1200

1300 F range. This is substantially lower than the 2200 F licensing limit set forth in 

1OCFR50 Appendix K. 

For the BWR/4, the ECC configuration is slightly different (See Slide 10). There is a 

low-pressure core spray system that injects liquid into the upper plenum similar to the 

BWR/6. There is no high-pressure core spray system, instead, there is a high-pressure 

_f- 83
coolant injection system that injects coolant into the downcomer region. The LPCI, 

instead of injecting into the bypass, connects to the drive line of the recirculation system. 

There are one low pressure core spray system, one high pressure coolant injection system, 

and four low pressure coolant injection systems, two into each recirculation system. 

BWR/6 LOCA GNF 

and Reflooding Restores 

llMlUnl Liquid Inventory and Quenches 

F Rod VVS-- D•nor IL"V• at Top of Jet.4 Pump - Two.P2-u Level Abomve ONO 

LPOS 

PCT 

BWR/6 PCT < 1000 °F 

JuOu MayL ý 9 

BWRI4 LOCA GN]Fr 

9 Early Part of Transient Similar 

to DWRM - Eady Boiling Transition 

Steam L~a -. ewet due to Lower Plenmm 

ll l - - *LCI - PII Injection Into jet Pump 

LPC * Corm Uncovery Leads to 

LM, LWSecond BT 

=DU]flB 7* -- HPCI Comes on (DiG Startup) 

SLPCS and LPCl Comes on 

when P < Shutoff Head 

* Worst Single Failure 

-AFaihm•. Sotmy --AwvtabW- I LPCS, I LPCl in each• 

100p, ADS 

"JYA Me " ;42

_Z- 84
The early part of the LOCA transient is very similar to the BWR/6 LOCA because it is 

dominated by the phenomena prior to the onset of the ECC systems (See Slide 11). Similar 

to the BWR/6, there is an early boiling transition, the fuel heats up and subsequently 

rewets due to lower plenum flashing, when the downcomer level has dropped below the 

top of the jet pump and the pressure has dropped below the saturation pressure 

corresponding to the liquid temperature in the lower plenum. The HPCI comes on based 

on the startup time for the diesel generators. 

BWR/4 LOCA 

* LPCS Injection 

Steam Line - Phenomena Similar to BWR/6 

- Level Forms in Upper Plenum 

Feed Water * LPCI Injection 

- Condensation on LPCI 

HPCI - Lower Plenum Refills 

-LPCS - Core Refloods 

_.LPC1 LPC Less Liquid Hold Up in Core 

iDuring Refill Phase for BWR/4 

Leads to Higher PCT for Second 

u L~iu -JPeak 

JGMA May 2000 I1 

The LPCI and the low pressure core spraying will come on when the pressure has 

dropped below the shutoff pressure for the pumps. Because of the time delay for the 

onset of these systems, the water level drops into the core region and there will be a 

second boiling transition as part of the core is uncovered. 

Evaluating the various failure combinations and applying the single failure criterion, the 

worst failure for the BWR/4 is the battery failure. Battery failure will prevent the startup 

of one diesel generator, which will powers two of the LPCI pumps, and it will also 

prevent the opening of the injection valve for the high pressure coolant injection into the 

downcomer. This leaves one low pressure core spray system and two LPCI systems, one 

T.-85
for each recirculation loop. The LPCI is also affected by isolation valves in the discharge side of the recirculation loop preventing the LPCI from being lost out of the break 

location. The BWR/4 also have an ADS system, and similar to the BWR/6 the ADS 

system is not significant for the large break design basis accident. 

For the BWR/4, when the two-phase level drops down into the lower part of the core, levels form in both the core and lower plenum (See Slide 11). This is very similar to the 

BWR/6. The phenomena associated with the low pressure core spray injection into the 

upper plenum are also very similar to the BWR/6. 

For the low pressure coolant injection, cold water is injected into the jet pump drive line 

and enters the jet pump through the nozzles. There is a very large condensation rate on 

the LPCI flow right inside the mixing region of the jet pump and steam is drawn into the 

jet pump from the downcomer region to sustain this condensation. This was tested in the 

SSTF as part of the Refill-Reflood test program. These tests showed that, as long the water 

level in the jet pumps does not cover the injection flow there is close to complete condensation inside the jet pumps on the LPCI flow coming through the nozzles and 

therefore saturated liquid will penetrate into the lower plenum. 

For the BWR/6, the LPCI liquid was injected into the bypass and it would flow through the leakage holes into the bottom of the channel and be held up inside the channel by the 

CCFL. This does not happen for the BWR/4. The liquid is injected into the jet pump and 

from there into the lower plenum. Therefore less liquid is held up inside the core region and the level drop in the core is larger compared to the equivalent level drop for the 

BWR/6 before the core starts to reflood, and it takes longer time to quench the core. 

Consequently, the second peak in the cladding temperature is higher for the BWR/4 than 

for the BWR/6. As the lower plenum fills and a corresponding filling of the jet pumps 

occur, the condensation on the LPCI in the jet pumps will be reduced and cold water will 

penetrate into the lower plenum. The subcooled water stratifies at the bottom of the lower 

plenum and has relatively little impact on the two-phase level and the refilling of the 

lower plenum. As a major part of the liquid in the lower plenum remain saturated, particularly at the two-phase level, the flashing from the lower plenum continues and 

holds up liquid in the core region providing some cooling of the fuel. The liquid from the 

LPCI injection quickly fills up the lower plenum and when the two-phase level gets up to 

the bottom of the core, the core starts reflooding and subsequently quenches to terminate 

the transient. 

In summary the temperature transients for the BWR/4 is quite similar to the BWR/6 

(See Slide 12). There is level inside the jet pump at the top of the jet pump, and a two

phase level inside the core shroud that is above the top of the active fuel.
BWR/4 LOCA

CW*

* Refilling and Reflooding 

Restores Liquid Inventory and 

Quenches Core 

- Downcomer Level at Top of Jet Pump 

- Two-Phase Level Above Core

PCT 

BWR/4 PCT = 1000 TF

JGMA May 2000 12

00 

4

TNWM
Because it takes a little longer to reflood the core for the BWR/4 the core heat-up is larger before it is quenched. Typically for BWR/4 the nominal best estimate PCT is in the order 

of 1000 F. The upper bound PCT at the 95/95 level is again a couple hundred degrees 

higher, and could be as high as 1400-1500 F. 

Slide 13 summarizes the major phenomena for a LOCA in a the jet pump BWR. Critical 

flow in the break location affects the liquid loss from the primary system and the 

depressurization of the system. Interfacial shear, controls the void fractions in the system 

and the two-phase levels in the various regions. Counter-current flow limitation is 

extremely important for the BWR. It occurs at the side entry orifice and at the top restrictions in the fuel channel, which is either the upper tie plate or it is the top spacer in 

the fully rodded section of the fuel bundles (the last spacer before the end of the part 

length rods). Interfacial heat transfer together with the depressurization determines the 

flashing in the lower plenum and in the core region as well as other regions of the primary 

system. Interfacial heat transfer is also the controlling phenomenon for condensation on 

the ECC as it is injected into the system either through the core spray nozzles in the upper 

plenum, as LPCI into the bypass for the BWR/6 or into the jet pump drive lines for the 

BWR/4. Condensation is the controlling phenomenon for subcooled CCFL breakdown, 

which is a major controlling phenomenon for the parallel channel effects that exists in the 

core of the BWR during the refill/reflood phase. Wall friction is 

important, because, as discussed earlier in the presentation, there can be three parallel 

flow patterns in the core during the refill-reflood part, co-current up-flow, counter-current 

flow with the two-phase level and co-current down-flow. The pressure drop has to be 

matched between lower plenum and the upper plenum. For co-current up-flow the 

dominant part of the pressure drop is wall friction. For the down-flow it is friction and 

static head, and for the counter-current flow it is predominantly static head associated 

with the two-phase level in the channels. Finally heat transfer is important. Fuel heat 

transfer (pellet conductivity and gap conductance) controls the initial stored energy in the 

fuel. The early boiling transition in the core happens primarily because of the flow 

reduction while there is still plenty of liquid in the core region, and the stored energy has 

a major impact on the early PCT. Once boiling transition occurs, film boiling becomes the 

dominant mode of heat transfer and a major controlling phenomenon for the PCT. 

Early in the transient, following the first boiling transition, the film boiling heat transfer is 

for the inverted annular type flow regime, and later on, for the co-current up-flow and 

counter-current flow regimes in the core region, prior to quenching, the film boiling heat 

transfer is for dispersed annular flow. Both modes of film boiling exist in the BWR.

"T-,88
Major Phenomena 

*Critical Flow 

*Interfacial Shear 

- Void Fraction 

- Two-Phase Levels 

lrnterfacial Heat Transfer 

- Flashing 

- CondensatIon 

- Subcooled CCR. Break Down 

eWall Friction 

*Heat Transfer 

- Boiling Transition 

- Film Boiling 

- Rewetting 

- Fuel Gap Conductance 

JGtA ,Var2 13 

Finally rewetting is important, it is controlled by two phenomena: the minimum film 

boiling temperature and when sufficient liquid is present to maintain a wetted condition. 

For the BWR/4 and /6, because the PCT is in the order 1000 degrees, metal-water 

reaction is not significant. There is no significant metal-water reactions at these low 

temperatures. Fuel rod failure mechanisms are also not significant because there is not 

enough stress on the cladding to cause failure at these low temperatures.

1•-- 89
For the BWR/2, the situation is different. The BWR/2 is the older generation of the BWRs 

without jet pumps (See Slide 14). In the BWR/2, the recirculation flow is taken out at the 

bottom of the downcomer and recirculation pumps inject it back into the lower plenum. For the BWR/2, the design basis accident is the double-sided break in the discharge side 

of the loop. This produces a large break connected directly to the lower plenum, and due 

to the large size of the recirculation pipe, it becomes impossible to reflood the core for the 

BWR/2. The only means to maintain cooling is the core spray system There are two 

independent core spray systems that inject water into the upper plenum once a LOCA has 

occurred for the BWR/2, and because the design basis accident is a break connected to the 

lower plenum, all core flow is lost almost instantly and there is an early boiling transition. 

There is no early quenching from lower plenum flashing as most of the liquid from the 

lower plenum is lost out through the break. 

BWR/2 LOCA 

* Early BT Like Jet Pump BWRs 

* Core Can Not Be Reflooded 

Steamm Linne Main Heat Transfer Phenomena: 

Core Spray Heat Transfer 

Feed Water - Convective Heat transfer to 

Superheated Steam 

CS - Radiation Heat Transfer to Channel 

Box and to Two-Phase Mixture - Conduction Controlled Rewetting 

- Metal-Water Reaction 

e PCT Controlled by Balance 

Between Decay Heat and CS 

Heat Transfer 

JGMA May 2000 14 

Quenching of the fuel for the BWR/2 does not occur until on the order of half an hour into 

the LOCA transient. During the heatup, the only cooling of the fuel is provided by the 

core spray system injecting water into the upper plenum and from there into the fuel 

channels. A two-phase level never forms in the upper plenum, because the up-flow of 

steam is not sufficient to hold up water in the upper plenum. The up-flow of steam into 

the upper plenum is low, because part of the steam generated trough vaporization in the 

core will leave the system through the bottom of the core and out through the break.

T--90
The spray distribution in the upper plenum and CCFL for the high power channels 

controls the drainage of liquid into the channels. 

The core spray heat transfer is the dominant mode and the major controlling phenomenon 

for the PCT. Core spray heat transfer consists of convection heat transfer to superheated 

steam with entrained droplets and radiation heat transfer. 

Film flow will occur on the channel box and the fuel rods after quenching. Thermal 

radiation heat transfer is a very dominant mode for the BWR/2. There is radiation heat 

transfer to the two-phase mixtures to the channel box. The radiation heat transfer to the 

channel box is the dominant mode as the channel box will quench relatively early in the 

transient because of the core spray water entering the channels and the bypass region. 

The quenching of the channel box is dominated by conduction controlled rewetting, and 

the quench front moves down from the top to the bottom of the channel box. 

The PCT becomes very high for the BWR/2 because of the relatively low value of the core 

spray heat transfer coefficient and because thermal radiation is only significant at high 

temperatures. Metal-water reaction will occur at high temperature, and at high 

temperatures close to the 2200F limit, it can become comparable in magnitude to the 

decay heat. The LOCA transient is significantly longer for the BWR/2 compared to the jet 

pump BWRs. The PCT typically occurred within 100-150 seconds for the jet pump BWRs. 

For the BWR/2, after the initial blowdown, the transient is an almost quasi-steady state 

transient, where the temperature of the fuel is determined by a balance between the 

energy that is being generated, which is the decay heat, plus some contribution from the 

metal-water reaction, and the heat transfer, which is the core spray heat transfer. Typically 

the PCT occurs 600-800 seconds into the transient. Quenching of the fuel rods could be 

very slow. It can be half an hour into the transient because the core never refloods and 

quenching of the fuel is mainly due to conduction controlled quenching from the top. For 

the fuel, the quench front propagates at a velocity that is in the order of millimeters per 

second. 

In summary, for the BWR/2 there is an early boiling transition, and the system, once the 

blowdown is over, enters a quasi-steady state mode controlled by a balance between the 

decay heat and core spray heat transfer. The cladding temperature is controlled by the 

magnitude of the core spray heat transfer and the gradual reduction of the decay heat, 

and the PCT occurs at the point where the energy generation becomes less than the core 

spray heat transfer. In the order of half an hour into the transient, the quench front has 

moved low enough on the fuel rods to quench the temperatures.

r7--91
Nominal best estimate PCT for BWR/2 is in the order of 1700 F (See Slide 15). Upper bound PCT at the 95/95 level is on the order of 2000 F. Licensing calculations for BWR/2 

with the SAFER/CORCL codes can give PCTs close to the 2200 F limit.

BWR/2 LOCA

* Core Spray Heat Transfer 

Controls PCT

PCT

BWRP2 PCT = 1700 OF 

IGA May WOa I.

_ 1-92

RLPý
Major phenomena for the BWR/2 are critical flow and interfacial shear (See Slide 16). 

Critical flow controls the blow down, while interfacial shear controls phenomena such as 

void fractions, two-phase levels, CCFL and the spray distribution in the upper plenum. 

The spray distribution in the upper plenum is very important for the BWR/2 as there will 

be no level in the upper plenum. It is not significant for the jet pump BWRs, because a 

level is maintained in the upper plenum during the refill and reflood phase. Similar as for 

the jet pump plants, interfacial heat transfer is significant for the BWR/2 as it controls 

flashing, condensation and subcooled CCFL breakdown.

Major Phenomena GE;

oCritlcal Flow 

.Jnterfacial Shear 

- Void Fraction 

- Two-Phase Levels 

- CCFL 

- Spray Distribution 

* Interfacial Heat Transfer 

-Fahing 

- Condensation 

- Subcooled CCFL Break Down 

eWall Friction

"* Heat Transfer 

- Boiling Transition 

- Film BoAng 

- Rewetting 

- Radiation Heat Transfer 

- Conduction Controlled 

Rowemtng 

"* Metal-Water Reaction 

"* Fuel Rod Failure

JGMA Hay 2y" 16

For heat transfer, thermal radiation heat transfer is a very dominant mode of heat transfer. 

For a fuel rod in a BWR/2, typically half or 50% of the heat transfer is thermal radiation. 

The dominant part of the radiation heat transfer is to the channel box once it is quenched 

by the core spray water flowing down on the inside and on the outside of the channel box. 

Quenching time for the channel is typically on the order of half a minute into the 

transient. Rewetting is a very important phenomenon for the BWR/2 because the 

quenching of the channel box and the associated change in emissivity has such a strong 

impact on the heat transfer. Finally because of the high temperatures obtained in a 

BWR/2 LOCA, metal-water reaction and fuel rod failure mechanisms are important 

phenomena. 

T93
The previous was a summary of the design basis LOCA for the various BWR types and 

the major controlling phenomena for these events. The following will contain a short 

discussion of the small break LOCA (SBLOCA). 

When LOCAs are analyzed, the entire break spectrum is evaluated to demonstrate that 

the design basis accident is indeed the limiting LOCA. For a SBLOCA (See Slide 17), the 

system may not depressurize, because only small amount of liquid leaves the system 

through the break while the vapor generation can make up for the volume loss. In this 

case the system is not depressurizing, however, the level will drop slowly because of the 

loss of liquid inventory. This is where the ADS system becomes important for the BWR. 

The Automatic Depressurization System (ADS) is needed to lower the pressure, because 

as long as the pressure stays high liquid from any of the low pressure systems can not be 

injected into the BWR, and the single failure criterion can eliminate the high pressure 

system. Most of the ECC systems are low pressure systems. Once a low level in the 

downcomer occurs, the ADS is tripped. The ADS valves will open and depressurize the 

system, allowing the low pressure systems to come on. In a way, the ADS converts the 

small break LOCA to an event that resembles a large steam line break LOCA. 

Small Break LOCA GNF 

*No Depressurization 

*Level Drop Due to Loss of Inventory 

*ADS Trip on Low Level 

e Low Pressure Activates LPCI and LPCS 

eADS Turns Small Break Into Large Break 

JGMA Miky 21XA 17

-""- 94
For the small break LOCA, there is no immediate reversal of the flow in half the jet 

pumps, therefore there is no drastic reduction in the core flow, but rater a slow reduction 

given by the coast down of the recirculation pumps. The early boiling transition is 

typically avoided for the small break LOCA and only the second boiling transition will 

occur. 

The last slide (Slide 18) shows the impact of fuel exposure. Typically, the thermal 

mechanical design limit determines the maximum linear heat generation rate; in this case 

here, in this case, the maximum average planar linear heat generation rate (MAPLHGR) is 

shown. The thermal mechanical design limit is an envelope of power versus exposure and 

as long as the fuel operate within that envelope, all of the classic design and licensing 

limits are met. The top curve on Slide 18 shows the thermal mechanical design limit for 

the jet pump BWRs. It is the design and licensing limit for the fuel, and the PCT for the jet 

pump plants, as shown earlier, are calculated with the fuel operating at this limit. This is 

the reason for the conclusion that the jet pump BWRs are not limited by the LOCA events. 

The BWR/2 is LOCA limited. Early in the transient, where the linear heat generation rate 

is high, the linear heat generation is limited by the peak cladding temperature. For the 

licensing calculation, the PCT gets close to the licensing limit of 2200F. The 2200F limit 

determines what the maximum linear heat generation can be in the fuel bundle. At some 

point, with increasing the exposure, as fission gas pressure or the pressure in the gap 

increases, the fuel rods may fail. When that happens, metal-water reaction is calculated 

on both sides of the cladding, both on the outside and on the inside. 

Under these conditions the limit of 17% maximum local oxidation can be reached before 

the PCT reach the 2200F limit. Typically for intermediate exposures as shown in Slide 18, 

the linear heat generation rate is limited by oxidation limit. In the calculation of the LOCA 

transient, it is assumed that there is no initial oxidation of the cladding, which maximizes 

the energy generations from the calculated metal-water reaction. The assumption for the 

initial oxidation on the fuel is currently subject to some discussion in the technical 

community. In conclusion, while the jet pump BWRs are not limited by the LOCA, the 

BWR/2 is limited by the LOCA.

_ -- 95
Fuel Exposure - MAPLHGR

*MAPLHGR = Maximum Average Planar Linear Heat 

Generation Rate

_ --" 96

WGN;ý

IMWA MJA 2Wt)O i8
THE HISTORY OF LOCA EMBRITTLEMENT CRITERIA

G. Hache 

Institut de Protection et de Suret6 Nucldaire 

Cadarache, France 

H. M. Chung 

Argonne National Laboratory 

Argonne, Illinois, USA 

Abstract 

Performance of high-bumup fuel and fuel cladding fabricated from new 

types of alloys (such as Zirlo, M5, MDA, and duplex alloys) under loss-of

coolant-accident (LOCA) situations is not well understood at this time. To 

correctly interpret the results of investigations on the performance of the 

old and new types of fuel cladding, especially at high bumup, it is 

necessary to accurately understand the history and relevant databases of 

current LOCA embrittlement criteria. In this paper, documented records 

of the 1973 Emergency Core Cooling System (ECCS) Rule-Making Hearing 

were carefully examined to clarify the rationale and data bases used to 

establish the 1204°C peak cladding temperature and 17% maximum 

oxidation limits. A large amount of data, obtained for zero- or low

bumup Zircaloy cladding and reported in literature only after the 1973 

Rule-Making Hearing, were also evaluated and compared with the current 

criteria to better quantify the margin of safety under LOCA conditions. 

1. Introduction 

Because of major advantages in fuel-cycle costs, reactor operation, and 

waste management, the current trend in the nuclear industry is to increase fuel 

discharge bumup. At high bumup, fuel rods fabricated from conventional 

Zircaloys often exhibit significant degradation in microstructure. This is 

especially pronounced in pressurized-water reactor (PWR) rods fabricated from 

standard Zircaloy-4 in which significant oxidation, hydriding, and oxide 

spallation can occur. Thus, many fuel vendors have developed and proposed 

the use of new cladding alloys, such as low-tin Zircaloy-4, Zirlo, M5, MDA, 

duplex cladding, and Zr-lined Zircaloy-2. Performance of these alloys under 

loss-of-coolant-accident (LOCA) situations, especially at high burnup, is not 

well understood at this time. Therefore, it is important to verify the safety 

margins for high-burnup fuel and fuels clad with new alloys. In recognition of 

this, LOCA-related behavior of various types of high-burnup fuel cladding is 

being actively investigated in several countries [1-6]. However, to correctly 

interpret the results of such investigations, and if necessary, to establish new 

embrittlement thresholds that maintain an adequate safety margin for high

burnup operation, it appears necessary to accurately understand the rationale,

-1- 97
history, and data bases used to establish the current LOCA criteria, i.e., 

maximum cladding temperature limit of 1204'C (2200'F) and maximum 

oxidation limit of 17%. For this purpose, documented records of the 1973 

Atomic Energy Commission (AEC) Emergency Core Cooling System (ECCS) 

Rule-Making Hearing were carefully examined and the relevant databases were 

reevaluated in this paper. Since the establishment of the current criteria, large 

amounts of data were obtained in many countries for zero- or low-burnup fuel 

cladding. The results of these investigations were also critically evaluated to 

determine the validity of the current criteria and safety margins for a wider 

range of conditions. 

2. Primary Objectives of Current Criteria 

In 1967, an Advisory Task Force on Power Reactor Emergency Cooling [7], 

appointed to provide "additional assurance that substantial meltdown is 

prevented" by core cooling systems, concluded that: 

"The analysis of (a LOCA) requires that the core be maintained in place 

and essentially intact to preserve the heat-transfer area and coolant

flow geometry. Without preservation of heat-transfer area and 

coolant-flow geometry, fuel-element melting and core disassembly 

would be expected... Continuity of emergency core cooling must be 

maintained after termination of the temperature transient for an 

indefinite period until the heat generation decays to an 

insignificant level, or until disposition of the core is made." 

This rationale makes it plainly clear that it is most important to preserve 

the heat transfer area and the coolant flow geometry not only during the short

term portion of the core temperature transient but also for long term. 

Consistent with the conclusions of the Ergen Task Force, the U.S. Atomic 

Energy Commission (AEC) promulgated Criterion 35 of the General Design 

Criteria [8] which states that: "... fuel and clad damage that could interfere with 

continued effective core cooling is prevented." It also promulgated Criterion 

3 of the Interim Acceptance Criteria for ECCS for LWR [91 which states that: 

"The clad temperature transient is terminated at a time when the core geometry 

is still amenable to cooling, and before the cladding is so embrittled as to fail 

during or after quenching." 

These criteria were subjected to a Rule-Making Hearing in 1973, which was 

extensively documented in the Journal of Nuclear Safety in 1974 [10,11]. 

During the hearing process, the last part of the Criterion 3 was replaced by the 

modified Criterion 1 and the new Criterion 2 of the Code of Federal Regulations, 

Title 10, Part 50.46, Article (b), commonly referred to as 10 CFR 50.46 [121. 

Thus, the AEC Commissioners wrote:

J- 98
"In view of the fundamental and historical importance of maintaining 

core coolability, we retain this criterion as a basic objective, in a more 

general form than it appeared in the Interim Acceptance Criteria. It is 

not controversial as a criterion... Although most of the attention of the 

ECCS hearings has been focused on the events of the first few minutes 

after a postulated major cooling line break, up to the time that the 

cladding would be cooled to a temperature of 300'F or less, the long

term maintenance of cooling would be equally important [13]." 

There are two key factors to consider to evaluate the change in coolable 

geometry of core, a brittle mode and a ductile mode of deformation in fuel 

cladding. The ductile mode is related to cladding ballooning, burst, and coolant 

channel blockage. This mode will not be treated in this paper. Our focus in 

this paper is on the change in coolable geometry due to cladding embrittlement 

and failure. 

3. Metallurgy of Cladding Embrittlement 

In 1960s, Wilson and Barnes performed laboratory tests simulating steam 

reactions with Zircaloy-clad fuel rods at high temperatures. They observed 

embrittlement of oxidized cladding well below the melting temperature of 

Zircaloy, either during the test itself or during removal of the specimen from the 

oxidizing furnace. The results were reported in Argonne National Laboratory 

(ANL) progress reports and synthesized later in Ref. 14. At the same period, 

investigators in Oak Ridge National Laboratory (ORNL) conducted TREAT Test 

No. 6 with Zircaloy cladding in steam and observed that the specimen was 

severely embrittled by oxidation [15]. Also at about the same period, many 

tests were conducted that simulated reactivity-initiated accident (RIA) in 

SPERT-CDC and TREAT reactors. Results of metallurgical examination in these 

tests showed that embrittlement was caused by severe microstructural 

modification of the cladding. Brittle cladding cross sections exhibited oxide 

layer, oxygen-stabilized alpha-phase layer and a region of acicular prior beta

phase. The results were later reported by Fujishiro et al. [16]. 

As a result of these observations, the scientific community was alerted to 

the fact that oxidation of Zircaloys above the alpha-to-beta transformation 

temperature results in the formation of inherently brittle phases, i.e., Zr oxide, 

oxygen stabilized alpha-Zr (fcc structure), and diffusion of oxygen into the 

underlying beta phase (bcc structure). This is shown schematically in Fig. 1. 

Ductility of cladding could be severely degraded if the degree of oxidation is 

high. It was also realized that, if the embrittled cladding fragments into small 

pieces, the coolability of the core could be seriously impaired.

'-99
ALPHA PRIOR BETA

Original OD Surface 

" Before Oxidation

.4 0.625 mm

oxide 

oxide 

alpha 

alpha Load-Bearing 

. Metal 

prior beta

0.12

Distance from OD Oxide Surface (mm) 

Fig. 1. Schematic illustration of microstructure (top) and 

oxygen distribution (bottom) in oxide, stabilized alpha, 

and prior-beta (transformed-beta) layers in Zircaloy 

cladding after oxidation near 1200'C.

._-"- 100

28.0 ID

6.7 1 .2 0 

2 

0 

C3) 

X 0 0.7

0 0.25 0.60

OXIDE
Significantly embrittled cladding can fragment during the quenching phase of a LOCA. The action of rewetting by ECCS water involves the collapse of the 

vapor film that covers the cladding outer-diameter (OD) surface prior to 

subsequent transition to nucleate boiling. This event takes place at a more or 

less constant temperature, i.e., the Leidenfrost temperature. For oxidized 

Zircaloy-4 cladding rewetted by bottom-flooding water, ANL investigators reported that rewetting occurs in the range of 475-600'C [171. The abrupt change in the heat transfer conditions induces large thermal-shock stress, which can fracture the cladding, if it is sufficiently embrittled by oxidation. 

Below the Leidenfrost temperature, there is continued risk of 

fragmentation after quenching. In accordance with the opinions of the Ergen Task Force and the AEC staff and commissioners mentioned earlier, other 

experts also wrote a similar opinion for OECD Committee on Safety of Nuclear 

Installations (CSNI) [18]: 'The ability of the cladding to withstand the thermal

shock stresses of quenching during rewetting or post-LOCA forces is related to 

the extent and detailed nature of oxidation during the transient. The post LOCA forces, which need to be taken into account, are the hydraulic, seismic, handling, and transport forces." 

There are two primary factors that exacerbate the susceptibility of oxidized 

cladding to post-quench embrittlement in comparison with susceptibility to 

fragmentation during quenching: i.e., (1) more pronounced effect of oxygen dissolved in beta phase at lower temperature of loading (i.e., more pronounced after quench than during quench) and (2) more pronounced effect of hydrogen uptake which may occur during irradiation (e.g., in high-burnup Zircaloy-4) or 

during transient oxidation in steam (e.g., from cladding inner surface in contact 

with stagnant steam near a ballooned and burst region). For cooling rates 

typical of bottom flooding of core (i.e., 1-5°C/s), most hydrogen atoms remain in 

solution in the beta phase at Leidenfrost temperature, and in such state, hydrogen has little effect on the fracture resistance of an oxidized Zircaloy. However, when load is imposed at temperatures below the Leidenfrost 

temperature, precipitated hydrides strongly influence the fracture resistance of 

cladding. Eutectoid decomposition of hydrogen-stabilized beta phase at 

temperatures below =55 0 °C [19] is the major factor that causes this deleterious 

effect (see Fig. 2). 

4. Opinion of Regulatory Staff and Commissioners during 

1973 Rule-Making Hearing 

4.1 Reluctance to Neglect Effects of Mechanical Constraints 

Some factors during a LOCA, such as ballooning of the rod near the spacer grid, rod-grid spring chemical interaction, and the friction between the fuel rod 

and spacer grids, can restrict the axial movement of the cladding. Also, guide tubes in a PWR fuel assembly are mechanically fixed to the spacer grids.

_T-1oi
Because of these factors, fuel rods during reflooding will be subject to tensile 

load that is produced due to the differential axial shrinkage between a cladding and the guide tube. Rods may interact each other due to ballooning or bowing. For high-burnup fuels in which tight pellet-cladding bonding is common, axial 

shrinkage can be restricted if the tight bonding remains unchanged after 

ballooning and burst. These constraints will remain after quench, when 

deleterious effects of oxygen and hydrogen are far more pronounced.

Temperature (*C) 

0 10 20

Concentration (at.%) 

30 40 -50

Zr

-863C 

beta

0.07 

Z~ral pha

E. Zuzek et al. 

Bulletin of Alloy Phase Diagrar 

Vol. 11, No.4 (1990) 385-395

/i+l: ii~

++++++ idelta + ++ I 9.+ # + lA , :+++] [] ++: ] + S + + + . : + elta d+f +++ion 

.+ 5+++++:++ ++:~ ++ S+ + + + + + T+ ; ++ + + ++: + + + + ' S+ + +++++++++ ++++ +? ++++ ++ ++ ++++ * U+++ !'I 

S. .... .. .. . . . . .. ! S U

0 0.4 0.6 0.8 1.0 1.2 1.4 

Hydrogen Concentration (wt.%)

In recognition of this, the AEC Staff wrote during the 1973 Rule-Making 

Hearing that "the loads due to assembly restraint and rod-to-rod interaction 

may not be small compared to the thermal shock load and cannot be 

neglected [20]." Subsequently, it was concluded that: 'The staff believes that 

quench loads are likely the major loads, but the staff does not believe that 

the evidence is as yet conclusive enough to ignore all other loads [21]." 

Then, the Commissioners added: "There is some lack of certainty as to 

just what nature of stresses would be encountered during the LOCA.... (We want) to draw attention to the fact that it may not be possible to anticipate 

and calculate all of the stresses to which fuel rods would be subjected in a 

LOCA. Although we believe the calculations of thermal shock stresses are 

worthwhile and informative, we agree with the regulatory staff that they are not 

sufficiently well defined to depend on for regulatory purposes [13]." 

Before 1973, no thermal-shock quench test was performed on 

mechanically constrained cladding specimens. Then in early 1980s, Uetsuka et

"7-C 102

&D

900 

8W0 

700 

6W~ 

500. 

400

300

200

100-

U .- ' i _ . . . . l .. . . . . . .. . .. . . .

Figure 2. 

Zr-H binary 

phase diagram 

(from E. Zuzek et 

aL, Bulletin of 

Alloy Phase 

Diagrams, Vol. 

11, No. 4, 1990, 

pp. 385-395).

Lb~ 1. 2-0 I I I .. .. | 7 • : • •.

I:
al. performed quenching tests on cladding sections under severely constrained 

condition [22]. In their experiment, cladding tube was fixed at the bottom but 

was allowed to freely elongate in axial direction during oxidation at high 

temperature. As a result, cladding length increased freely because of thermal 

expansion and oxide-induced creep. At the end of the isothermal oxidation, the 

specimen top was fixed to the crosshead of an Instron tensile facility. Then, the 

load-time curve was continuously monitored during quenching. Thus, at 

Leidenfrost temperature, the cladding tube was subjected to combined axial

tensile and thermal-shock stresses. The results of the tests are summarized in 

the Fig. 3. Similar tests were also performed on unconstrained tubes (Fig. 4).

Oxida1ion 

TMeg 1300

10000 

8000 

6000 

2000 

1000 

800 

600 

400 

200

Oxidation Temperature (OC) 1200k 1100 1000 900

I I I I ITotal Oxidation Calculated 

with Baker-Just Eq. 

H. Uetsuka et al. 40% 200/15% 

J. Nucl. Sci. Tech. 

Z 20 (1983) 941-950 

a ..41 z 

0 0 

* 0 

.* / 

3 

00 

-0 0 ^ 

- 0/ 0 

/ / * Failed on Quench~ing 0

o Survivred 

I 1 nnrl I'V 6.0

I

7.0 8.0 

1/T x 10000 (inverse K)

I

I

Figure 3. 

Failure-nonfoailure 

boundary for fully 

constrained 

Zircaloy-4 after 

oxidation in steam 

Sand quenching as 

function of 

oxidation time and 

30 temperature; total 

oxidation 

calculated with 

Baker-Just 

10 equation is also 

indicated (from 

5 Uetsuska et aL, J. 

Nucl ScL Tech. 

20, 1983, pp. 941

950).

Oxidation 

Time (min.)

A comparison of the results from the two contrasting- types of test shows a 

large effect of the mechanical constraint. However, it is difficult to conclude 

whether the degree of constraint in the experiments of Uetsuka et al. is 

prototypic of a LOCA or unrealistically too severe. The 17% oxidation limit, 

calculated with Baker-Just correlation, appears to be adequate for protection of 

constrained rods against thermal-shock failure (Fig. 3), whereas a large margin 

is evident for unconstrained rods (Fig. 4).

J-103

,I,•• I •! I•,i i o: i

I'
Unlike other bundle tests such as NRU, REBEKA, JAERI and ORNL 

multirod tests that were entirely devoted to the study of ballooning, burst, and 

flow-channel blockage, some of the tests in Phebus LOCA program was devoted 

to the study of embrittlement [231. The fragmented Rod 18 of the Test 219, 

exposed to =1330°C, is especially interesting (see Fig. 5). For this oxidation 

temperature, results of calculation with PRECIP-II Code [24] indicates that the 

0 content in the beta phase was higher than 0.9 wt.%, a threshold 0 

concentration found to be associated with thermal-shock failure or survival 

[17]. Rod 18 fragmented despite it was oxidized to an equivalent-cladding 

reacted (ECR) value of only -16%. This observation indicates a deleterious 

bundle effect, i.e., an additional mechanical constraint. 

As a conclusion, results of the JAERI constraint quench test and the 

PHEBUS-LOCA Test appear to justify the reluctance of the AEC staff and 

commissioners to neglect the effect of mechanical constraints on the 

susceptibility to thermal-shock failure.

Oxidation Temperature (°C) 

1400 1300 1200 1100 1000

I I I I 

H. Uetsuka et al. 

J. Nucl. Sci. Tech. 

20 (1983) 941-950

I I 

Total Oxidation Calculated 

with Baker-Just Eq. 

40% * 20% 

/7 / /10%

-I

-1

-i

* Failed on Quench 

O Survived Quench

Figure 4. 

Failure-nonfailure 

boundary for 

unconstrained 

100 Zircaloy-4 after 

oxidation in steam 

60 and quenching as 

function of 

oxidation time and 

30 temperature; total 

20 oxidation 

calculated with 

Baker-Just 

10 equation is also 

shown (from 

5 Uetsuska et aL, J. 

Nucl. ScL Tech. 

20, 1983, pp. 941

950).

Time 

(min.)

___Tj- 104

Oxidation 

Time (min.)

Time 

(s) 

10000 

5000 

2000 

1000 

500 

200 

100

6.0 7.0 8.0 

1/T x 10000 (inverse K)

7

I I
313 4 

T, 28 31 FIS 3 

F23 i~o !I fj~a, 3, 6!; 

"'_2 

75 as I 

LOA2 94 Tes Bud Leve 7578

Figure 5. 

Cross-section of 

portion of fuel 

assembly from 

PHEBUS LOCA

219 Test (Ref. 

23), showing 

fragmentation of 

cladding.

4.2 Preservation of Ductility and Consideration of Results 

from Unconstrained Quench Test

At the end of the 1973 Hearing, the AEC Commissioners wrote: 

"...Nevertheless we find the quench results encouraging in that they provide 

assurance that the 2200'F limit is conservative. Our selection of the 2200'F 

limit results primarily from our belief that retention of ductility in the 

Zircaloy is the best guarantee of its remaining intact during the 

hypothetical LOCA... The thermal shock tests are reassuring, but their use 

for licensing purposes would involve an assumption of knowledge of the 

detailed process taking place in the core during a LOCA that we do not 

believe is justified [ 13]." 

Without much ambiguity, this conclusion clearly expressed the belief that 

retention of ductility was considered the best guarantee against potential 

fragmentation under various types of loading (thermal-shock, bundle 

constraints, hydraulic, handling, and seismic forces). During the 1973 Hearing, 

results from unconstrained quench tests (simple thermal-shock test) were 

considered only corroborative and reassuring. However, their use for regulatory 

purposes was not accepted. 

Results of later investigations on unconstrained or partially constrained 

cladding [17,18] showed a large margin of survival under thermal shock relative 

to 17%-ECR and 2200'F (1204'C) peak temperature limits. Such results are 

summarized in Fig. 6. No fragmentation occurred for ECR < 17% for all 

oxidation temperatures, whereas significant margin of survival was observed for

,_1 -105
oxidation temperatures <1204'C. The results in Fig. 6 were limited for thermal

shock tests in which cladding tube or ring was directly quenched from the 

maximum oxidation temperature without slow cooling through the range of 

beta-to-alpha-prime transformation. For slow-cooling conditions, more 

pronounced margin of survival was observed [17]. 

100 I I I 

Nonirradiated Zircaloys Figure 6. 

Direct Quenching after FAIL INTACT Failure boundary 

80 Oxidation at High Temperature Chung & Kassner *0 

H + * of partially 

"a 70 constrained 

*'o 1204*C 1204C • limit limt Grandjean & Lebuffe V Z-caloy c cladding 

a 60 tubes or 

5 unconstrained 

M= rings after 

S4o 6 oxidation at high 

temperature and 

0 °direct quenching S20 _ ORO from peak 

Aoo AA 8 17% limit oxidation 

10 990 '000A 

A oa O oA* temperature (from ,, aAq- Refs. 17 and 29). 1100 1300 1500 1700 1900 

Oxidation Temperature (°C) 

5. 17% -Oxidation Criterion 

5.1 Establishment of 17% Criterion During 1973 Rule-Making Hearing 

The rationale for establishment of the two criteria in 10 CFR 50.46(b) is 

described in this section. As indicated in a few reports [17,18] that reviewed 

the results of the LOCA-related tests performed before and after the 1973 

Hearing, the 17%-ECR and 1204'C criteria were primarily based on the results 

of post-quench ductility tests conducted by Hobson [25,261. 

Figure 7 summarizes the results of Hobson's ring compression tests 

performed at 23-150'C. Zircaloy-4 cladding tubes were oxidized in steam on 

two sides, followed by direct quenching into water. Then, short ring specimens 

cut from the oxidized tube were either compressed slowly to a total deflection of 

3.8 mm or squashed by impact loading. After the test, the broken pieces of the 

ring was assembled back to determine the degree of brittleness. Zero ductility 

was defined on the basis of the macroscopic geometry of the broken pieces and 

the morphology of the fracture surface on microscopic scale. Each data point in 

Fig. 7 indicates failure type, test identification number, oxidation time in min., 

oxidation temperature in OF, and first maximum load in pound.

-. 106
ZIRCALOY SLOW0-COMPRESSION TESTS 

O.~t/i.CROSSHEAD SPEED "0,2• -. n- LONG RINGS 

I I LOAD tFIRST MAXIMUM) it)) "0 TOTAL DUCTILITY Deformation& ONE TO THREE FRACTURES 

Temperature FOUR FRACTURES -- r : -.. • .. lt ZERO OUCTJLý•TY . 

(OF) 4W0 FTEST TEMPERATURE VARIAT ION

Co 

o :.0 

0:l(i

0 

M 

S

2 '2 

0 7

0 

0

NOTE: THE CURVES OASEO ON THE SLOW COMPRESSION 

TESTS EXCLUDE THE SPECIMENS EXPOSED AT 

2400'F.

1~0 

• o-, 4,. 

10 . " 

I

0 

0

Hobson 1973 

Ref. 25

2W0*

100 

0

-NI

0l

oa 

"ZERO OUCTILITY kt 

SLOW COMPRESSION 1.

0 

0I

0g 

0 

NI

0 

o4 '2

I

o 00 

M 0 

"*• fT T 

I4., a

0 010 0". 0.000o 

• I*. \

0 0 0,0 

0"0 

.40

0 

T 

0 

0 

0

"TOTAL DUCTILITY" 

ORNL- 4758

I gocO 0

0 

0

0.5 0.6 0.7 0.8 

Beta-Layer Thickness Divided by As-Oxidized Wall

0.9

Figure 7. 

Ductility of two-side-oxidized Zircaloy rings as function of slow- or 

fast-compression temperature and fraction of transformed-beta-layer 

(from Hobson, Ref. 25 and 26). 

The dashed line on the left side of Fig. 7 denotes the.zero ductility domain 

for slow-compression rate. This domain is valid only for oxidation temperatures 

of <2200'F or <1204*C. During the 1973 Hearing, ORNL investigators 

suggested to consider a zero-ductility temperature (ZDT) no higher than the 

saturation temperature during reflood, i.e., =135°C. Zero-ductility threshold at 

this temperature is equivalent to a beta-layer fraction of =0.58, or a fraction of 

combined oxide layer plus alpha layer thickness (defined as XT) of =0.42 (based

_r- 107

300.

I

"I I I

I
on as-oxidized cladding wall). The latter fraction corresponds to =0.44 if it is 

calculated based on fresh nonoxidized cladding wall (defined as Wo). 

The threshold fractional thickness of the combined oxide and alpha layer (XT/Wo, defined as Xoa in Fig. 8) of 0.44, which corresponds to zero ductility 

threshold for slow compression at 135°C, was the key number in the 

establishment of 17% oxidation criterion in the 1973 Hearing. During the 

hearing, the AEC Regulatory Staff wrote: 

"Giving due credit to the numerous quench experiments and the ORNL 

zero ductility experimental data points for both impact and slow 

compression, the staff suggests that an embrittlement criterion be 

based on a calculated XT/WO that shall not exceed 0.44. This is 

equivalent to a zero ductility temperature of about ... 275°F based 

on the slow compression tests [201." 

Then, it was concluded: 

'To preclude clad fragmentation and to account for effects noted in the 

tests described above, a limit of XT/Wo < 0.44 was earlier suggested 

by the Regulatory staff as an embrittlement criterion (Exhibit 1113, 

page 18-18). This limit was inferred from quench tests and 

mechanical tests. Criterion (b)(2) is now proposed as a better method 

of specifying a similar limit on the extent of cladding oxidation. The 

bases for proposing this method are described below: (The) use (of the 

17 percent reaction limit) with the Baker-Just equation is 

conservative when compared to the previously suggested limits of 

XT/WO < 0.44. This is shown in Figure 8 (of this paper) for isothermal 

conditions. Four lines of constant calculated XT/Wo (two for 0.44 

and two for 0.35) are constructed on the plot of percent reaction 

versus a parameter proportional to the square root of exposure time. 

The solid XT/Wo lines are based on Pawel's equation (Exhibit 1133) 

(Ref. 27 of this paper), and the dashed lines are based on Exhibit 09, 

page 9, Figure 5 (Ref. 25 of this paper). As can be seen, the 

XT/Wo = 0.44 lines are both above the 17 percent reaction line..." 

Results of a total of five key tests and calculations are summarized in Fig. 

8, a complex but the most important step used to reach the 17% oxidation 

limit. They are: (1) equivalent cladding reacted (ECR) calculated as function of 

oxidation temperature and square root of time based on Baker-Just correlation, 

(2) two broken curves which define the time and temperature to reach the 

threshold fractional thickness of the combined oxide and alpha layer (denoted 

as Xoa) of 0.44 and 0.35, as determined based on the data given in Ref. 25, 

Page 9, Fig.5, (3) two solid curves that define the time and temperature to reach 

the threshold fractional thickness of the combined oxide and alpha layer of 0.44

108
and 0.35, as determined based on the method of Ref. 27, (4) six ECR-(time)0 .5 

curves from the thermal-shock tests of Hesson et al., Ref. 14, and (5) results 

from Combustion Engineering (CE) ring compression tests after one-sided 

oxidation. 

Equivalent Cladding Converted to Oxide - (% Calculated with Baker-Just Equation)/ Oxidation 

34 2700 2600 2500 2400 2300 2200 Temperature 2100 

32 

in 

Thq al-Shock Failures 

30- L-7609, Hesson 19 

28-20 rmHobson & Ritenhos 

24-/ 26 

18 oa. 0. 0.4 

20- 1800 

18- Xa 0.4 ýý17% O~xidation Line 

16 . . .. Xoa = 0.35 

12

10- 'Xoa = combined thickness 

of oxide and alpha layers 

8- /divided by fresh cladding S~wall thickness 

6

Per Pawel, J. Nucl. Mater. 

50 (1973) 247-258 N = 1 for one-side oxidation 

2 N = 2 for two-side oxidation 

(0.343NICladding Thickness in mm) x Square Root of Time in s 

2 4 6 8 10 12 14 16 18 20 22 24 26 28 30 32 34 36 38 40 42 44 46 48 

a 0.250 ( 11 '5~ 31 4 66 7' A 9' 10 12 14 116 118 2'042 A Ab 2 :4 36 38 

Oxidation lime (min.) for 0.686-mm-thick Cladding 

Figure 8. 

Summary of multistep procedure used to establish 17% oxidation 

criterion during 1973 Rule-Making Hearing (from Docket RM-50-1, April 

16, 1973). Note equivalent cladding oxidized was calculated per Baker

Just correlation. For comparison, time to reach threshold fraction of 

combined oxide and alpha layers of 0.44 is shown as determined per 

Hobson and Rittenhouse (ORNL-4758, January 1972) and Pawel (J. 

NucL Mater. 50, 1973, pp. 247-258). 

By definition, ECR parameter varies depending on cladding wall 

thickness, either due to differences in fuel design or due to ballooning and

j..7-109
burst during the heatup phase in a LOCA. Figure 8 shows how to take account 

of the effects of variations in wall thickness and one- vs. two-sided oxidation. 

Two of Hesson's thermal-shock experiments resulted in cladding 

fragmentation at calculated ECR values of =19 and =30%, as indicated in the 

figure. The other four did not fail at ECR values of =21, = 16.5, =l10, and =9.5%. 

The time-temperature transients in Hesson's tests were integrated also by using 

the Baker-Just equation. 

The CE data, discussed in the Hearing, are represented by squares on the 

oxidation isotherms of 2500, 2400, 2300, and 2100°F. If the sample fractured 

on compression by CE's load standard, it was considered to have failed and is 

denoted with a filled square. Open squares denote CE's non-failed specimens. 

By the CE's load standard, only those samples with calculated ECR values 

> 17% failed. 

Based on the results given in Fig. 7 and the five sets of information shown 

in Fig. 8, one can conclude that no samples tested by slow compression at 

>135°C failed with zero ductility if equivalent cladding reacted (ECR), 

calculated on the basis of Baker-Just correlation, was less than 17%. 

Furthermore, all samples oxidized to <17% ECR (again calculated with Baker

Just correlation) survived direct quenching. 

In summary, the AEC Commissioners concluded that the very good 

consistency between the 17% limit, if calculated with the Baker-Just 

equation, and a wide variety of experiments supports adoption of this 

procedure [21J, and it was further stated: 

"There is relatively good agreement among the industrial participants 

as to what the limit on total oxidation should be.... The regulatory staff 

in their concluding statement compared various measures of oxidation 

and concluded that a 17 % total oxidation limit is satisfactory, if 

calculated by the Baker-Just equation... As argued by the regulatory 

staff, it appears that the 17% oxidation limit is within the Rittenhouse 

criteria. Thus a remarkable uniformity of opinion seems to exist with 

regard to the 17% oxidation limit [131." 

It is clear that the primary rationale of the 17% criterion is retention of 

cladding ductility at temperatures higher than 275°F (135°C, i.e., the 

saturation temperature during reflood). Of major importance in this proceeding 

is that the threshold ECR value of 17% is tied with the use of Baker-Just 

correlation. That is, the 17% ECR criterion is specific to Baker-Just 

correlation that must be used to determine the degree of total oxidation. If an 

oxidation correlation other than the Baker-Just equation (e.g., Cathcart-Pawel 

correlation) were used, the threshold ECR would have been less than 17%.

110
This means that use of a best-estimate correlation may not necessarily be 

conservative in evaluating post-quench cladding ductility. 

5.2 Other Embrittlement Criteria Proposed after the 1973 Hearing 

Few months after the 1973 Hearing, Pawel proposed a new criterion based 

on <95 % saturation of the average oxygen concentration in the beta phase [27]. 

However, such a criterion fails to recognize that in addition to a sufficiently low 

O concentration, a minimum thickness of beta layer is required to ensure 

adequate resistance to failure. Such criterion is less facilitated to handle, 

especially during non-isothermal LOCA transients, and it requires a computer 

code that can accurately calculate 0 diffusion under moving-phase-boundary 

conditions, a task more difficult than the calculation of a simple parabolic 

oxidation correlation. Nevertheless, many of such computer codes have been 

developed after the 1973 Hearing, e.g., those reported in Refs. 17 and 24. 

Sawatzky performed room-temperature tensile tests on specimens exposed 

to high-temperature spikes in steam [28]. Based on results of microhardness 

measurement, the distribution of 0 in the transformed beta (or prior beta) layer 

was found to be nonuniform, an observation confirmed subsequently by ANL 

investigators by Auger electron spectroscopy (Fig. 32-43, Ref. 17). In spite of 

total oxidation of only 16 %, a specimen with average 0 concentration >0.8 wt% 

in the prior beta exhibited very low strength and negligible elongation, whereas 

a specimen with 0 content <0.6 wt% in the prior beta retained some ductility. 

Based on this observation, Sawatzky proposed to replace the 1204'C PCT and 

the 17% ECR criteria by a unified criterion, that is, oxygen concentration in 

beta layer shall be <0.7 wt% over at least half of the cladding thickness. At 

temperatures >1280'C, Sawatzky's criterion is virtually identical to Pawel's 

criterion (see Fig. 9). 

Validity of the three criteria illustrated in Fig. 9 is, however, subject to 

variations in cladding wall thickness, because the time to reach the specified 

threshold state of material is strongly influenced by the clad wall thickness 

which may vary with fuel design and the degree of ballooning and burst. Thus, 

it was deemed desirable to develop a unified embrittlement criterion that would 

be valid independent of variations in wall thickness and oxidation temperature 

1171. 

5.3 One- vs. Two-Side Oxidation and Thermal-Shock Failure 

Grandjean et al. have reported results of extensive thermal-shock tests 

which were performed in TAGUIS facility [29,301. Hydrogen uptake in their 

short ring specimens was not excessive. In their investigation, ECR was 

calculated with PECLOX oxidation code [31], and failure-survival behavior was 

determined based on the result of gas-leakage check. The results of the tests 

were included in Fig. 6. The effect of one- vs, two-side oxidation on thermal-

_- 111
shock failure was the focus of investigation. As indicated in Fig. 8, such effect 

was considered negligible in establishing the 17% ECR limit in the 1973 Rule

Making Hearing. Interestingly, Grandjean et al.'s failure threshold for two-side 

oxidation appears to be slightly higher than the threshold for one-sided 

oxidation, i.e., =21 vs. =20% ECR. Nonetheless, this study provides an 

independent confirmation of the validity of the 17% ECR criterion relative to 

susceptibility to thermal-shock failure. 

Time to Embrittlement (sl

1000 1100 1200 1300 1400 1500W 

Oxidation Temperature (°C)

Figure 9. 

Comparison of 

current 

embrittlement 

criteria with those 

proposed by Pawel 

(Ref. 27) and 

Sawatzky (Ref. 28).

5.4 17% Oxidation Limit and Impact Failure at Small Hydrogen Uptake 

After the 1973 Hearing, ANL investigators conducted impact tests to 

provide an independent verification of the validity of 17% ECR threshold with 

respect to cladding resistance to impact failure [17]. Impact tests were 

performed at room temperature on non-pressurized open-ended Zircaloy-4 

tubes that were oxidized on two sides in steam at 1100-1400'C and cooled 

through the beta-to-alpha-prime transformation range at 5 or =10 0 °C/s. 

Because the sample was oxidized on both OD and ID sides, hydrogen uptake 

was limited to <130 wppm. Therefore, microstructure and oxygen and 

hydrogen distributions in the specimens were similar to those of the ringJ-- 112

1000 -

1001-

10

900

Nm 

0.7% THROUGH HALF OF SHEATH

I•P, V#,I L.N.T' 1,111I ~iiPIUI'J ---- 02' 0.7% AND <95% SATURATION 

(PAWEL'S CRITERIA) 

T' 1204C -..... O2<17% REACTED 

(USAEC CRITERIA)

I - I I _ _ I I I

0O24,
compression specimens of Hobson [25,261 that were cooled fast through the 

beta-to-alpha-prime transformation range. 

It was found that slow-cooled specimens were more resistant to impact 

failure than fast-cooled specimens (Fig. 65, Ref. 171. Results obtained for slow

cooled specimens are summarized in Fig. 10. The ECR values in Fig. 10 were 

directly determined based on measured phase layer thickness, therefore, are 

considered more accurate than values calculated based on Baker-Just 

correlation. The results in Fig. 10 show that for cladding oxidized at <1315 0C 

to <17% ECR, a sufficient level of resistance to impact failure is retained at 

230C, i.e., failure impact energy of >0.8 J. 

0.8 1 1 1 1 1 1 1 1 . I I I I I I I • I I I - I - I 

Zircaloy-4 tube Impact test at 230

C 

0.7 OD 10.9 mm, WT 0.635 mm 13150C Figure 10. 

H < 130 wppm 0.6 after oxidation Failure impact 

CIO energy vs. 

M 0.5 oxidized in steam equivalent .,at 11 00°C >. 0 A cladding 

cm A 

0.4 reacted, from 

, tests at 23'C on 

S0.3 - undeformed ca L 14000A A Zircaloy-4 tube E oxidized on two * 0.2 oe 

. 0 1204oc . sides and "• ~cooled at 5°C/ s 

1L. 0.1 Chung & Kassner 1980 (2 oed a7). 

NUREG/CR-1344(Ref. 17). 

4 8 12 16 20 24 28 32 36 

Equivalent Cladding Reacted (%) 

5.5 17% Limit and Ring-Compression Ductilty at Small Hydrogen Uptake 

As shown in Fig. 8, the 17% threshold ECR was derived by indirect 

multistep procedure. Of particular importance in this procedure is the 

accuracy of two key factors, i.e., (1) temperature measurement in the 

experiments of Baker-Just and Hobson-Rittenhouse [25,26] and (2) definition of 

nil-ductility as given in Fig. 7. In consideration of this, ANL investigators 

performed independent compression tests at room temperature on short 

Zircaloy-4 ring specimens. Rings were sectioned from long tubes that were 

oxidized in steam at 1100-1400'C and cooled through the beta-to-alpha-prime 

transformation range at 5 or =100°C/s. Hydrogen uptake in the ring specimens 

was <130 wppm. This procedure reproduced the conditions of the ring

compression tests of Hobson. In the ANL compression tests, however, load-

"T- 113
deflection curves were obtained to better quantify the degree of remaining 

ductility and the magnitude of load that a ring can sustain. 

It was found that slow-cooled specimens retained more ductility than fast

cooled specimens under otherwise identical conditions (Fig. 67, Ref. 17]. Figure 

11 summarizes results obtained for a slow-cooling rate of =5°C/s, a rate 

probably more prototypic of a LOCA than fast cooling. The ECR values in the 

figure were determined based on measured phase layer thickness and time

temperature history. This result shows that for cladding oxidized at <1315 0C to 

<17% ECR, ductility is retained at 230C (i.e., relative diametral deflection 

>16%); no brittle failure was observed. This experiment provides an 

independent confirmation of the validity of the 17% oxidation limit for 

undeformed Zircaloy specimens that contain hydrogen <130 wppm. 

2.0 1 ' 1 . 1 1 1 , I -I -I - 1 I 

Z0rcaloy-4 Figure 11. 

OD 10.9 mm, WT 0.635 mm -F 

6.35-mm-long ring oxidized in steam 

cut from preoxidized tube 1315'C at 11 0ooC Total deflection 

1.5 H < 130wppm at230 C vs. 

E 12040C equivalent Slow ring compression 

r_ 0.043 mm/s at 230C 

o .oreacted, from 

o ring

14000 \e 0 compression 

00 tests on 

120.5 A 0Zircaloy-4 0s oxidized on two

Chung & Kassner 1980 A A sides and 

NUREG/CR-1344 A A cooled at50C/s 0 . L I I- L I . I - I l I . 1 

0.00 4 8 12 16 20 24 28 32 36 (Ref. 17). 

Equivalent Cladding Reacted (%) 

5.6 Resistance to Impact Failure at Large Hydrogen Uptake 

In addition to the impact tests on non-ruptured empty tubes, ANL 

investigators performed 0.15- and 0.3-J pendulum impact tests at 230C on 

pressurized Zircaloy-4 tubes that were burst, oxidized, cooled at =5°C/s, and 

survived quenching thermal shock [17]. The CSNI experts [18] considered that: 

"Ambient impact of 0.3 J were thought to be a reasonable approximation to post 

LOCA quench ambient impact loads." The results of the 0.3-J impact tests, 

summarized in Fig. 12, indicate that the 17%-ECR limit is adequate to prevent 

a burst-and-oxidized cladding from failure under 0.3-J impact at 230C, as long 

as peak cladding temperature remained _<1204°C. The ECR values in the figure 

were determined based on measured thickness of oxide, alpha, and beta phase

-.T1-- 114
layers, rather than calculated based on Baker-Just correlation, and hence, are 

considered more accurate. 

70 

Zircaloy-4 tube filled with Figure 12. 

alumina pellets, 

60 0 OD 10.9 mm, WT 0.635 mm, Impact failure - burst and oxidized in steam threshold as 

50 H 20-2200 wppm function of 

after quench 

0 failed equivalent 

M 40 0 intact Impact test cladding reacted 

0 at 23°C and oxidation 

030 *0 o ternperature of - 1 burst, oxidized, 

0 0 slow-cooled, *•2o- 2-0 9 -oo° 0 •^ 8 0 

_e 0 ,______ and quenched 

.c-o ° o • Zircaloy-4 tube S0 O( 

- 113- O containing 20

.1 o Chung & Kassner 1980 wt i m 

0 NUREG/CR-1344 2200 wppm S'' , ' 1 1 hydrogen ffrom 

1000 1100 1200 1300 1400 1500 

Isothermal Oxidation Temperature (°C) Ref. 17). 

In contrast to two-side-oxidized non-pressurized non-ruptured tubes in 

which hydrogen uptake was small (<130 wppm), burst Zircaloy-4 tubes 

exhibited peculiar oxidation behavior near the burst opening. The inner

diameter (ID) surfaces of the top and bottom "necks," =30-mm away from the 

burst center, were exposed to hydrogen-rich stagnant steam-hydrogen mixture 

which is produced because of poor mixing of steam and hydrogen at the narrow 

gap between the alumina pellets and the ID surface of the necks. As a 

consequence, thick breakaway oxides formed at 900-1120'C [ 17], and hydrogen 

uptake as high as =2200 wppm was observed at the "necked" regions. 

Subsequently, JAERI investigators confirmed occurrence of the same 

phenomenon [32,331. 

The results from the same tests shown in Fig. 12 were converted to failure

survival map based on average hydrogen content of the impact-loaded local 

region and the thickness of transformed-beta layer containing <0.7 wt.% 

oxygen. This failure-survival map is shown in Fig. 13.. On the basis of the 

figure, ANL investigators proposed to replace the 1204'C PCT and 17% ECR 

criteria by a unified criterion which specifies that the thickness of transformed

beta layer containing <0.7 wt.% oxygen shall be >0.3 mm [17]. The criterion 

implicitly incorporates a limit in peak cladding temperature. This limiting 

temperature corresponds to the temperature at which oxygen solubility is 0.7 

wt.% in Zircaloy that contains 700-1200 wppm hydrogen. This temperature is 

believed to be between 1200 and 1250'C, although the exact data from

7- 115
applicable Zircaloy-O-H ternary diagrams are not. This criterion is not subject 

to variations in cladding wall thickness and oxidation temperature. 

0.6 I I I 1 1 I 1 1 1 1 1 

0 Zircaloy-4 tube filled with 

0 aumina OD pellets, Figure 13. 10.9 mm, WT 0.635 mm, 

-0 0 0 burst and oxidized in steam o Im pactfailure 

"0 0 0 0 0 0 

"0 0 000 0 o map as function 

0E of hydrogen 

SE O 0 00 0 content and 

0o. thickness of 0 0 of0 

cc __ _a_ - beta layer 

0. o containing <0.7 

.0 H 20-2200 wppm wt% oxygen; 

. after quench Zircaloy-4 tube 

0,- Impact test burst, oxidized, 

Sat23°c slow-cooled, 

& • - 0.11 0 • 0 • 0 and quenched - (from Ref. 17). Chung & Kassner 1980 

NUREGICR-1344 

"--To v , 0 12t 1600 2000 2400 

Hydrogen Content (wppm) 

The results in Fig. 13 show that for a given thickness and a given oxygen 

content in transformed-beta layer, resistance of cladding to impact failure is 

significantly reduced if hydrogen uptake exceeds =700 wppm. Such situation 

does not occur in non-pressurized, non-ruptured, two-side-oxidized Zircaloy 

cladding, such as those tested by Hobson [26] or discussed in Figs. 10 or 11. 

5.7 Ring-Compression Ductility at Large Hydrogen Uptake 

Investigators in ANL [17] and JAERI [32,33] conducted extensive tests on 

tube or ring specimens of Zircaloy-4 that contained high concentrations of 

hydrogen. In the former investigation, Zircaloy-4 tubes filled with alumina "pellets" were pressurized, heated, burst, oxidized, slow-cooled, and quenched 

with bottom-flooding water. Then, the tubes that survived the quenching 

thermal shock were compressed diametrically at 23'C [171. Such specimens 

contained H up to =2200 wppm. In the latter investigation, short rings, 

sectioned from tubes that were exposed to similar conditions, were compressed 

at 100 0C. The ring specimens contained H up to z1800 wppm. Typical 

distributions of oxide layer thickness, hydrogen concentration, and ring 

deflection to failure are shown in Fig. 14. The top and bottom "necks" that 

contained the highest concentration of hydrogen and the thinnest transformed

beta layer exhibited the lowest ductility.

S-116
However, ANL investigators observed that the rate of hydrogen generation, 

amount of hydrogen uptake, and hence, the degree of embrittlement of the 

necked regions are strongly influenced by the method of heating cladding tubes 

during LOCA-like transients, i.e., more uniform (indirect heating in JAERI) vs. 

less uniform (direct heating in ANL) heating [17]. This is schematically 

illustrated in Fig. 15.

Oxide Layer Thickness (micrometer)

30 

25

20-

15F

10o

5

uxiaizea in .team Oxide Layer 10590C 240 s Thickness 

at 100 

** . (mm) 

Ring Deflection Hydrogen 

at I000 Content 

Uetsuka et al. 

1981 I I - I I 

Hydrogen 

\Content 

\A. 

/ 'A, "-- 

-60 -45 -30 -15 0 15 30 45 60

Bottom Distance from Burst Center (mm)

Hydrogen 

Content 

(W ppm) 

oti'1 1500 

0°C

5 

4 

3 

2 

1

1000 

500

Figure 14. 

Distributions of 

hydrogen content, 

inner-diameter 

oxide layer 

thickness, and 

total deflection at 

100-C of ring 

specimens 

sectioned from 

burst region (from 

Uetsuka et al., 

Refs. 32 and 33).

0

Top

The effect of hydrogen uptake on post-quench ductility, determined either 

from diametral-compression test of burst-and-oxidized tubes at 23°C [17] or 

compression at 100'C of ring specimens sectioned from burst-and-oxidized 

tubes [32,33], is summarized in Fig. 16. At hydrogen-uptake >700 wppm, 

significant embrittlement of cladding is evident, even if total oxidation is <17% 

(see Fig. 14). Similar dependencies of plastic deflection on beta-layer oxygen 

content and total hydrogen content have been also reported in Fig. 88, Ref. 17 

and Fig. 89, Ref. 17, respectively. These results show that post-quench 

ductility of Zircaloy is strongly influenced by not only oxidation but also

_7:- 117

I
hydrogen uptake. This is shown in Fig. 17. Apparently, the important effect of 

hydrogen uptake on post-quench ductility was not well realized at the time of 

1973 Hearing.

C>

U 

'C 

I

0. 

U.' 

0 

0

RUPTURE From Chung & Kassner 1980 RUPTURE 

CENTER CENTER 

Hydrogen Generation Rate Hydrogen Uptake Rate

200 400 600 800 1000 1200 1400 

Hydrogen Content (wppm)

Figure 15. 

Effect of heating 

method (uniform 

vs. nonuniform 

heating) on 

hydrogen 

generation and 

uptake near burst 

opening (from 

Ref. 17)

Figure 16. 

Effect of 

hydrogen uptake 

on diametral 

deflection of 

burst, oxidized, 

and quenched 

Zircaloy-4 tube or 

sectioned ring.

1600 1800 2000

Essentially similar observation has also been reported by Komatsu et al. 

[34,35]. They reported that the load to initial ring cracking is strongly 

influenced by total oxidation and hydrogen uptake. For oxidation temperatures 

>12600C in which the oxygen content in the beta layer exceeds -0.7 wt.% in 

short period of time, the embrittling effect of oxygen appears to be predominant 

(see Fig. 18). The "zero-ductility" region denoted in Fig. 18 appears to have

_r- 118

I 

U.' z 

U.' 

z 

U.' 

0

UNIFORMI (% ! 

HEATING ,, \ 

NO I NII 

H I \ t .1 \ 

/ "-I 

/N ,.--,, IF0R\

14 

o 12 

S10 

0 o 

00

I I I I I I I I I I I I I I I I I I 

Zircaloy-4 Tube Burst and 

Oxidized in Steam 0m 

. Chung & Kassner 1980; Tube Diameter Compressed at 23*C 

to Fragmentation or to Pellect Contact 

"-. m Uetsuka et al. 1982, 

Ring Cut from Oxidized Tube 

Compressed at 1 00C 

-e . *U1 • ' 0 • o 

mm, a a111J1116S
been determined based on a threshold load to initial cracking rather than based 

on ductility consideration. As such, this "zero-ductility" threshold differs 

significantly from that defined by Hobson [25,261. 

-° Chung & Kassner 1980 

Figure 17. o for fresh Zircaloy-4 

burst and oxidized in Post-quench ductility 

4 steam near 1 atm. pressure shown as function of 

o oxidation (beta-layer -"centerline oxygen content) 

2 and total hydrogen 

content, from diametral 

compression test at 23 0C _ on burst, oxidized, slow

0 cooled, and quenched 

Zircaloy-4 tubes (for 

.. 4. • •database see Ref. 17). 

40 Total/ 

Oxidation (%) Z u 7 

30 -- w/,. 0• Region Fqr 8 300 

o/ 0 0 Applied load at initial 

20 - 51 cracking of ring as 

o " _ / function of total oxidation 

temperature, from 

10 (3 0"-'-• Komatsu et aL, Refs. 34 " r~~("5g--..- o0 " and 35. 

Komatsu et al. o ad3 

Refs. 34 & 35 ( )Load at causing Initial Crack 

0 1100 1200 1300 

Exposure Temperature (0C) 

5.8 17% Oxidation Criterion - Summary 

It is clear that the primary rationale of the 17% ECR criterion is retention 

of cladding ductility at temperatures higher than 275°F (135°C), i.e., the

_T_ 119
saturation temperature during reflood. The threshold ECR value of 17% is tied 

with the use of Baker-Just correlation. If a best-estimate correlation other than 

Baker-Just equation (e.g., Cathcart-Pawel correlation) were used, the threshold 

ECR would have been < 17%. 

Investigations conducted after the 1973 Rule-Making Hearing showed that 

for oxidation temperatures <1204'C, the 17% oxidation limit (as calculated with 

Baker-Just correlation) is adequate to ensure survival of fully constrained or 

unconstrained cladding under quenching thermal shock. It was also shown 

that the 17% limit (ECR determined on the basis of measured phase layer 

thickness) is adequate to ensure retention of ductility and resistance to 0.3-J 

impact failure in non-irradiated, non-ruptured, two-side-oxidized Zircaloy 

cladding in which hydrogen uptake during a LOCA-like transient is small. 

However, the 17% limit appears to be inadequate to ensure post-quench 

ductility for hydrogen uptake >700 wppm. Such level of large hydrogen uptake 

could occur in some types of fuel rods during normal operation, especially at 

high burnup, or during a LOCA-like transient in localized regions in a ballooned 

and ruptured node. 

6. 12040C (22000F) Peak Cladding Temperature Criterion 

6.1 Selection of 1204'C Criterion in 1973 Hearing 

From the results of posttest metallographic analysis of the slow-ring

compression specimens, Hobson 126] observed a good correlation between zero 

ductility temperature (ZDT) and fractional thickness of transformed-beta layer 

(or the sum of oxide plus alpha layer thickness) as long as the specimen was 

oxidized at -<2200°F (12040 C) (see Fig. 7). However, in spite of comparable 

thickness of transformed beta layer, specimens oxidized at 2400°F (1315 0C) 

were far more brittle. This observation was explained on the basis of excessive 

solid-solution hardening of transformed-beta phase at high oxygen 

concentrations. For mechanical properties near room temperature the critical 

concentration of oxygen in the transformed-beta was estimated to be =0.7 wt%. 

Above this concentration, transformed beta phase becomes brittle near room 

temperature. Because of the solubility limit of oxygen in the beta phase, this 

high 0 concentration cannot be reached at 2200'F (1204'C) but can be reached 

at 2400°F (1315 0C). Hobson concluded that: "embrittlement is not simply a 

function of the extent of oxidation alone, but is related in yet another way to the 

exposure temperature." 

During the 1973 Rule-Making Hearing, AEC Staff endorsed Hobson's 

conclusion and wrote: "The staff recognizes the importance of oxygen 

concentration in the beta phase in determining the load bearing ability of 

Zircaloy cladding, and the implication from the recent compression tests that 

this may not be satisfactorily characterized above 2200°F by a ZDT as a

120
function of remaining beta fraction only. We therefore believe that peak 

cladding temperatures should be limited to 2200'F [20]." 

Subsequently, it was also concluded that: 

"Additional metallurgical and slow compression mechanical tests on other 

quenched samples from the ORNL experiments indicated that an important 

consideration was the amount and distribution of oxygen in the nominally 

ductile prior-beta phase. However, these factors could not be correlated as 

functions of time and temperature in the same manner as the (combined oxide 

and alpha layer) penetration. In particular, the slow compression tests 

indicated a greater degradation in cladding ductility at higher temperatures 

than would be expected from considerations of (combined oxide and alpha 

layer) penetration alone. It was on this basis that the staff previously 

suggested a 22000F maximum cladding temperature... What was observed in 

the slow compression tests was that 6 samples exposed at 2400'F for only two 

minutes and with relatively high values of Fw (Fw being fractional thickness of 

prior beta, all greater than 0.65) all fractured with nil ductility... Only when 

brittle failure was detected at high Fw in the slow compression tests did the 

suspicion arise that ductility was a function of both Fw and the exposure 

temperature... As the temperature rises above 2200-2300'F, solid solution 

hardening in the beta phase appears to contribute significantly to formation of 

a brittle structure. That is, brittle failure occurs even though alpha incursions 

are not observed, and the fraction of remaining beta is greater than that 

observed in lower temperature tests. This is confirmed by examination of the 

six samples from the ORNL exposed at 2400'F for two minutes (Exhibit 1126)... 

From the foregoing, there is ample evidence that load bearing ability and 

ductility decrease with increasing exposure temperatures, even for transients 

with comparable Fw. Increased solubility of oxygen in the prior-beta phase has 

been discussed as a contributing factor... The staff believes that because of high 

temperature degradation ... phenomena (... strongly suggested by the 

experimental evidence cited), the suggested 22000F limit should be imposed 

[211." 

Then it was added: 

"The situation is complicated by the fact that not all of the prior beta phase 

is equally strong or ductile, since these properties depend on the amount of 

dissolved oxygen. This fact has been suspected for some time... From the phase 

diagram, given by both Scatena and Westinghouse, it is obvious that it is 

possible for the beta phase zirconium to take on a higher oxygen content at 

2600'F than at 2000'F. Furthermore, since the diffusion rate depends 

exponentially upon temperature, one might expect a greater incursion of oxygen 

into the beta phase for a given thickness of oxide and stabilized alpha phase at 

higher temperatures... Others (than Hobson) have also observed that the 

resistance to rupture depends upon the temperature at which oxidation occurs

_7 -121
as well as the extent of oxidation... To recapitulate, measures of Zircaloy 

oxidation, whether by percent, XT, or Fw, are largely or wholly determined from 

the brittle layers of zirconium oxide or stabilized alpha phase, while the 

ductility and strength of oxidized zirconium depend upon the condition and the 

thickness of the prior beta phase... Thus a criterion based solely on the extent 

of total oxidation is not enough, and some additional criterion is needed to 

assure that the prior beta phase is not too brittle. The specification of a 

maximum temperature of 2200'F will accomplish this adequately. The data 

cited in exhibit 1113 would not support a choice of a less conservative 

limit [13]." 

Few months after the Hearing, Pawel [27] explained Hobson's observation 

based on data that indicate oxygen solubility in the beta Zr at 2200-2400'F 

(1204-13160 C) is =0.7 wt.%. The 0 solubility in beta Zircaloy is significantly 

influenced by not only temperature but also the concentration of hydrogen, a 

strong beta stabilizer. Nevertheless, Pawel endorsed that: "...the above 

reasoning easily explains why the mechanical or load bearing properties of the 

oxidized specimens should not be a unique function of the extent of (total) 

oxidation." Consequently, Pawel proposed to replace the peak cladding 

temperature (PCT) criterion by a new criterion that specifies the average oxygen 

concentration in the beta phase shall be less than 0.7 wt% [see Fig. 9]. 

6.2 1204°C Limit vs. In-Pile Test Results 

In 1970s, high-temperature oxidation and embrittlement behaviors were 

investigated extensively in TREAT and PBF test reactors. During the TREAT

FRF2 test, a seven-rod cluster was oxidized at 2400'F (-z13150 C) [36]. 

According to hardness measurements, all rods contained portions that 

possessed no ductility at room temperature. Three rods were broken 

accidentally during handling in ORNL hot cell, which indicates the degree of 

brittleness of a badly embrittled rod and the magnitude of a typical load during 

handling in hot cell (see Fig. 19). 

Some fuel rods tested in the Power Burst Facility (PBF) were also known to 

have failed during handling or posttest examination in hot cell. This 

information is summarized in Fig. 20 [37]. Total oxidation of several failed rods 

were <17 %. Of particular interest is Rod IE-019 of Test IE 5, because 

ballooning and burst occurred in the rod before exposure to temperatures 

>1,100°C. In spite of the fact that ECR was only =12%, the rod broke into pieces 

after exposure to an "equivalent" oxidation temperature of -1262°C. Most 

likely, actual peak temperature was higher than this equivalent temperature. 

Rod A-0021 also ruptured before entering high temperature transient; this 

caused ingress of steam to the rod interior. The rod failed after exposure to 

=1307°C, although ECR was only =6%. Hydrogen uptake in the two rods was 

excessive because of exposure to stagnant steam near the rupture opening.

2/_7 122
Figure 19.

Fuel pellet released through 

fragmented cladding section 

of Rod 11, TREAT-FRF 2 

Test (from Ref. 37). 

It is not clear if the failure behavior of Rods IE-019 and A-0021 is predicted based on Pawel's criterion (Fig. 9). However, because the exposure 

temperatures of the rods exceeded =1262°C, the thickness of beta layer that 

contained <0.7 wt.% oxygen should have been zero or close to zero. However, because oxygen solubility in beta is influenced by hydrogen and because 

accurate peak temperatures reached in the rods are not well known, it is 

difficult to calculate accurately the thickness of beta layer that contains 0 •0.7 

wt.%. Therefore, it is not clear if the failure behavior of the two rods is 

consistent with the criterion shown in Fig. 13. 

As long as clad oxidation temperature was limited to <1204°C, a handling failure at measured ECR < 17% was not observed from the TREAT and PBF tests 

or the ANL 0.3-J impact tests (see Fig. 20). This observation clearly demonstrates the importance of the 1204'C PCT limit. That is, the 1204'C PCT 

and the 17% ECR limits are inseparable, and as such, constitute an integral 

criterion. 

6.3 Summary of 1204'C Criterion 

The 2200'F (1204'C) peak cladding temperature (PCT) criterion was 

selected on the basis of Hobson's slow-ring-compression tests that were 

performed at 25-150'C. Samples oxidized at 2400'F (1315'C) were far more 

brittle than samples oxidized at <2200'F (<1204'C) in spite of comparable level 

of total oxidation. This is because oxygen solid-solution hardening of the prior

beta phase is excessive at oxygen concentrations >0.7 wt%. 

The selection of the 1204'C criterion was subsequently justified by the 

observations from the ANL 0.3-J impact tests and the handling failure of rods

.. T- 123
tested in the Power Burst Facility. These results also take into account of the 

effect of large hydrogen uptake that occurred near the burst opening. Consideration of potential for runaway oxidation alone would have lead to a PCT limit somewhat higher than 2200'F (1204'C). In conjunction with the 17% oxidation criterion, the primary objective of the PCT criterion is to ensure adequate margin of protection against post-quench failure that may occur 

under hydraulic, impact, handling, and seismic loading. 

70 1 1 1 

* PBF, intact 

60 * B PBF non-breached rod, fail F 20. 

"" PBF breached rod, fail 

0 ANL 0.3-J impact, intact Comparison of 

a 50- 0 ANL 0.3-J impact, fail o 0data from hot

M 40] cell handling failure of 

- . Zircaloy rods " 30 0 9o- exposed to high 0 o temperature in "20 00 o Power Burst 5 0 0 8 UTA0004 . BSF-1 183 ~~~~~IE015Faiiy(e37 IE016 Facilit(Rf 37) 

0-121. ,E022 and 0.3-J 

tr 10 0 Q E021 W 20-4 impact tests in ""0 

0 

le • $ A 1E1 ANL (Ref. 17). ,, A0021 E0 19 

1000 1100 1200 1300 1400 1500 1600 1700 

Maximum Oxidation Temperature (°C) 

7. Conclusions 

I. In the 1973 Rule-Making Hearing, the U. S. Atomic Energy Commission 

(AEC) staff and commissioners were clearly reluctant to neglect the effect of 

mechanical constraints on the susceptibility of oxidized fuel cladding to thermal-shock fragmentation. Subsequent test results appear to justify this rationale. Results from unconstrained or partially constrained quench tests 

were considered only corroborative; their use for regulatory purposes was 

not accepted. 

2. The AEC staff and commissioners and OECD-CSNI specialists were of the 

opinion that retention of ductility was the best guarantee against potential fragmentation of fuel cladding under various types of not-so-well-quantified loading, such as thermal shock, hydraulic, and seismic forces, and the 

forces related with handling and transportation.

7-. -124
3. Primary rationale of the 17% oxidation criterion was retention of cladding 

ductility at temperatures higher than 275°F (135°C), i.e., the saturation 

temperature during reflood. The threshold equivalent cladding reacted 

(ECR) of 17% is tied with the use of Baker-Just correlation. If a best

estimate correlation other than Baker-Just equation (e.g., Cathcart-Pawel 

correlation) had been used, the threshold ECR would have been < 17%. 

4. Investigations conducted after the 1973 Rule-Making Hearing show that for 

oxidation temperatures <1204°C, the 17% oxidation limit (calculated with 

Baker-Just correlation) is adequate to ensure survival of unconstrained or 

fully constrained cladding under quenching thermal shock. It was also 

shown that the 17% limit (ECR determined on the basis of measured phase 

layer thickness) is adequate to ensure retention of ductility and resistance to 

0.3-J impact failure in non-irradiated non-ruptured two-side-oxidized 

Zircaloy cladding in which hydrogen uptake during a LOCA-like transient is 

small. 

5. However, the 17% ECR limit appears to be inadequate to ensure post

quench ductility at hydrogen concentrations >700 wppm. A major finding 

from tests performed after the 1973 Rule-Making Hearing shows that post

quench ductility is strongly influenced by not only oxidation but also 

hydrogen uptake. It seems that this effect of large hydrogen uptake was not 

known at the time of 1973 Hearing. 

6. By definition, an embrittlement criterion expressed in terms of ECR is 

subject to uncertainties because calculated ECR varies with variations in 

cladding wall thickness and the degree of ballooning. 

7. The 1204'C peak cladding temperature (PCT) limit was selected on the basis 

of slow-ring-compression tests that were performed at 25-150'C. Samples 

oxidized at 1315'C were far more brittle than samples oxidized at 1204'C in 

spite of comparable level of total oxidation. This is because oxygen solid

solution hardening of the prior-beta phase is excessive at oxygen 

concentrations >0.7wt%. Consideration of potential for runaway oxidation 

was a secondary factor in selecting the 1204'C limit. The 12040C limit was 

subsequently justified by the observations from impact tests and handling 

failure of fuel rods exposed to high temperatures in the Power Burst Facility. 

The 1204*C PCT and the 17% ECR limits are inseparable, and as such, 

constitute an integral criterion. 

8. The degree of oxygen saturation and the thickness of beta layer that 

contains oxygen concentrations _50.7 wt.% were important parameters used 

by investigators to develop new embrittlement criteria based on beta phase 

thickness rather than total oxidation. Such a criterion is not subject to 

inherent uncertainties associated with variations in cladding wall thickness 

and pre-LOCA oxidation. 

_... -- 125
9. Post-quench ductility and toughness are determined primarily by the 

thickness and the mechanical properties of transformed-beta layer. The 

mechanical properties are strongly influenced by several factors such as: 

oxygen solubility in beta, concentrations of alpha- (tin and oxygen) and beta

stabilizing elements (niobium and hydrogen), the nature of beta-to-alpha

prime transformation, redistribution of oxygen, niobium, and hydrogen 

during the transformation, and precipitation of hydrides. Significantly large 

hydrogen uptake can occur in some types of fuel cladding, during normal 

operation to high burnup, during breakaway oxidation at <1 120'C, and, for 

localized regions near a rupture opening, during LOCA transients. Hydrogen 

uptake and its effect on the properties of transformed beta could differ 

significantly in Zircaloys and in niobium-containing alloys. Considering 

these factors, it is recommended to obtain a better understanding of the 

effects of more realistic hydrogen uptake and niobium addition on the 

properties of transformed-beta layer and post-quench ductility. 

Acknowledgments 

The authors would like to express thanks to C. Grandjean, R. 0. Meyer, 

and S. Basu for helpful discussions. 

References 

1. Grandjean, C. et al., "Oxidation and quenching experiments with high 

burnup cladding under LOCA conditions," Proc. 26th Water Reactor Safety 

Information Meeting, Bethesda, USA, 26-28 October, 1998. 

2. Mardon, J.P. et al., "The M5 fuel rod cladding," Proc. ENS TOP'FUEL 99, 

Avignon, France, 13-15 September, 1999. 

3. Nagase, F. et al., "Experiments on high burnup fuel behavior under LOCA 

conditions at JAERI," Proc. ANS Topical Meeting on LWR fuel Performance, 

Park City, USA, 10-13 April, 2000. 

4. Ozawa, M. et al., "Behavior of irradiated zircaloy4 fuel cladding under 

simulated LOCA conditions," Proc. 12th Symp. on Zirconium in the Nuclear 

Industry, Toronto, Canada, 15-18 June, 1998. 

5. Aomi, M. et al., "Behavior of irradiated BWR fuel cladding tubes under 

simulated LOCA conditions," Proc. ANS Topical Meeting on LWR fuel 

Performance, Park City, USA, 10-13 April, 2000. 

6. Chung, H.M. et al., 'Test plan for high-bumup fuel cladding behavior under 

loss-of-coolant accident conditions," Proc. 24th Water Reactor Safety 

Information Meeting, Bethesda, USA, 21-23 October, 1996. 

7. Report of Advisory Task Force on Power Reactor Emergency Cooling, TID

24226, 1967.

,_2._ 126
8. General Design Criteria for Nuclear Power Plants, U.S. Code of Federal 

Regulations, Title 10, Part 50, Appendix A, 20 February, 1971, amended. 

9. Interim Acceptance Criteria for Emergency Core-Cooling Systems for Light

Water Power Reactors, U.S. Federal Register 36 (125), pp. 12247-12250, 29 

June, 1971. 

10. Cottrell, W.B., "ECCS rule-making hearing," Nucl. Safety 15 (1974) 30-55. 

11. New acceptance criteria for emergency core-cooling systems of light-water

cooled nuclear power reactors, Nucl. Safety 15 (1974) 173-184. 

12. Acceptance Criteria for Emergency Core Cooling Systems for Light-Water 

Nuclear Power Reactors, U.S. Code of Federal Regulations, Title 10, Part 50, 

Section 46, 4 January, 1974, Amended. 

13 Atomic Energy Commission Rule-Making Hearing, Opinion of the 

Commission, Docket RM-50-1, 28 December, 1973. 

14. Hesson, J.C. et al., Laboratory Simulations of Cladding-Steam Reactions 

Following Loss-of-Coolant Accidents in Water-Cooled Power Reactors, ANL

7609, January 1970. 

15. Parker, G.W. et al., "Release of fission products from reactor fuels during 

transient accidents simulated in TREAT," Proc. Intl. Symp. Fission Product 

Release and Transport under Accident Conditions, Oak Ridge, USA, 5-7 

April, 1965. 

16. Fujishiro, T. et al., "Light water reactor fuel response during reactivity 

initiated accident experiments," NUREG/CR-0269, August 1978. 

17. Chung, H.M. and Kassner, T.F., Embrittlement Criteria for Zircaloy Fuel 

Cladding Applicable to Accident Situations in Light-Water Reactors, 

NUREG/CR-1344, January 1980. 

18. Parsons, P.D. et al., The Deformation, Oxidation and Embrittlement of PWR 

Fuel Cladding in a Loss-of-Coolant Accident: A State-of-the-Art Report, 

CSNI Report 129, December 1986. 

19. Zuzek, E. et al., 'Tne H-Zr (hydrogen-zirconium) system," Bulletin of Alloy 

Phase Diagrams, 11 (1990) 385-395. 

20. Atomic Energy Commission Rule-Making Hearing, Supplemental Testimony 

of the Regulatory Staff Docket RM-50-1, 26 October, 1972. 

21. Atomic Energy Commission Rule-Making Hearing, Concluding Statement of 

the Regulatory Staff, Docket RM-50-1, 16 April, 1973. 

22. Uetsuka, H. et al., "Failure-bearing capability of oxidized Zircaloy-4 cladding 

under simulated loss-of-coolant condition," J. Nucl. Sci. Tech. 20 (1983) 

941-950.

127
23. Reocreux, M. and Scott de Martinville, E., "A study of fuel behavior in PWR 

design basis accident: An analysis of results from the PHEBUS and EDGAR 

experiments," Nucl. Eng. Design 124 (1990) 363-378. 

24. Suzuki, M. and Kawasaki, S., "Development of computer code PRECIP-II for 

calculation of Zr-steam reaction," J. Nucl. Sci. Tech. 17 (1980) 291. 

25. Hobson, D.O., and Rittenhouse, P. L., "Embrittlement of Zircaloy Clad Fuel 

Rods by Steam During LOCA Transients," ORNL-4758, Oak Ridge National 

Laboratory, January 1972. 

26. Hobson, D.O., "Ductile-brittle behavior of Zircaloy fuel cladding," Proc. ANS 

Topical Mtg. on Water Reactor Safety, Salt Lake City, 26 March, 1973. 

27. Pawel, R.E., "Oxygen diffusion in beta Zircaloy during steam oxidation," J. 

Nucl. Mater. 50 (1974) 247-258. 

28. Sawatzky, A., "Proposed criterion for the oxygen embrittlement of Zircaloy-4 

fuel cladding," Proc. 4th Symp. on Zirconium in the Nuclear Industry, 

Stratford-on-Avon, UK, 27-29 June, 1978. 

29. Grandjean, C., "Oxidation and quenching experiments under simulated 

LOCA conditions with high bumup clad material," Proc. 24th Water Reactor 

Safety Information Meeting, Bethesda, USA, 21-23 October, 1996. 

30. Grandjean, C. et al., "High burnup U02 fuel LOCA calculations to evaluate 

the possible impact of fuel relocation after burst," ANL Program Review 

Meeting, Rockville (USA), 22 October, 1999. 

31. Hofmann, P. et al., PECLOX: A Computer Model for the Calculation of the 

Internal and the External Zircaloy Cladding Oxidation, KFK-4422 Part 2, 

October 1988. 

32. Uetsuka, H. et al., "Zircaloy-4 cladding embrittlement due to inner surface 

oxidation under simulated loss-of-coolant condition," J. Nucl. Sci. Tech. 18 

(1981) 705-717. 

33. Uetsuka, H. et al., "Embrittlement of Zircaloy-4 due to oxidation in 

environment of stagnant steam," J. Nucl. Sci. Tech. 19 (1982), 158-165. 

34. Komatsu, K., "The effects of oxidation temperature and slow cooldown on 

ductile-brittle behavior of Zircaloy fuel cladding," Proc. CSNI Specialists' 

meeting on the behavior of water reactor fuel elements under accident 

conditions, Spaatind, Norway, 13-16 September, 1976. 

35. Komatsu, K. et al., "Load-bearing capability in deformed and oxidized 

Zircaloy cladding," Proc. CSNI Specialist Mtg. on Safety Aspects of Fuel 

Behavior in Off-Normal and Accident Conditions, Espoo, Finland, 1-4 

September, 1980. 

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Nucl. Tech. 11 (1971) 502-520.

128
37. Haggag, F. M., Zircaloy-Cladding-Embrittlement Criteria: Comparison of In

Pile and Out-of-Pile Results, NUREG/CR-2757, July 1982.

1'-- 129
NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER (2-89) (Assigned by NRC, Add Vol., Supp., Rev., NRCM 1102, B DATA SHEET and Addendum Numbers, if any.) 

(See instructions on the reverse) NUREG/CR-6744 

2. TITLE AND SUBTITLE LA-UR-00-5079 

Phenomenon Identification and Ranking Tables (PIRTs) for Loss-of-Coolant Accidents 

in Pressurized and Boiling Water Reactors Containing High Bumup Fuel 3. DATE REPORT PUBLISHED MONTH YEAR 

December 2001 

4. FIN OR GRANT NUMBER 

W6245 

5. AUTHOR(S) 6. TYPE OF REPORT 

B.E. Boyack, A.T. Motta, K.L. Peddicord, C.A. Alexander, J.G.M. Andersen, 

J.A. Blaisdell, B.M. Dunn, D. Ebeling-Koning, T. Fuketa, G. Hache, L.E. Hochreiter, Technical 

S.E. Jensen, S. Langenbuch, F.J. Moody, M.E. Nissley, K. Ohkawa, G. Potts, 7. PERIOD COVERED (Inclusive Dates) 

J. Rashid, R.J. Rohrer, J.S. Tulenko, K. Valtonen, N. Waeckel, W. Wiesenack 

04100 to 11/00 

8. PERFORMING ORGANIZATION - NAME AND ADDRESS (If NRC, provide Division, Office or Region, U.S. NuclearRegulatory Commission, and marling address; if contractor, 

provide name and mailing address.) 

Los Alamos National Laboratory 

Los Alamos, NM 87545

9. SPONSORING ORGANIZATION - NAME AND ADDRESS (if NRC, type "Same as above',- if contractor, provide NRC Division, Office or Region, U.S. Nuclear Regulatory Conmmisson, and mailing address) 

Division of Systems Analysis and Regulatory Effectiveness 

Office of Nuclear Regulatory Research 

U.S. Nuclear Regulatory Commission 

Washington, DC 20555-0001

10. SUPPLEMENTARY NOTES

H. Scott, NRC Project Manager

11. ABSTRACT (200 words or less) 

In the United States, cladding embrittlement criteria and related evaluation models are used to address loss-of-coolant 

accidents. The embrittlement criteria are a peak cladding temperature of 1204 C (2200 F) and an equivalent oxidation of 17% 

of the cladding wall thickness calculated with the Baker-Just correlation. Evaluation models address ballooning, rupture, flow 

blockage, and oxidation kinetics. The NRC is performing research with respect to high bumup fuel to acquire and develop the 

requisite understanding of the performance of high bumup fuel under accident conditions. It is also conducting research to 

determine if current embrittlement criteria and evaluation models are adequate for high-bumup fuel or if modifications are needed. To support these efforts, the NRC has commissioned the formation of a Phenomena Identification and Ranking Table 

(PIRT) panel to identify and rank the phenomena occurring during selected transient and accident scenarios in both pressurized 

water reactors and boiling water reactors containing high bumup fuel. Because the PIRT identifies and ranks phenomena for 

importance, currently existing experimental data, planned experiments, computational tools (codes), and code-calculated results 

can be screened to determine applicability and adequacy using the PIRT results. This PIRT identifies and ranks phenomena for 

loss-of-coolant accidents in both pressurized and boiling water reactors.

12. KEY WORDS/DESCRIPTORS (List words or phrases that will assist researchers in locating the report.) 13. AVAILABILITY STATEMENT 

cladding oxidation embrittlement rankings and rationales unlimited 

14. SECURITY CLASSIFICATION 

LOCA PWR BWR break-size primary evaluation criterion (This Page) 

unclassified 

expert elicitation plant transient analysis (This Repot) 

unclassified 

15. NUMBER OF PAGES 

16. PRICE

NRC FORM 335 (2-89)
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NUCLEAR REGULATORY COMMISSION 

WASHINGTON, DC 20555-0001 

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PENALTY FOR PRIVATE USE, $300