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NEA/CSNI/R(2006)5

28-Nov-2006___________________________________________________________________________________________English - Or. English

NUCLEAR ENERGY AGENCYCOMMITTEE ON THE SAFETY OF NUCLEAR INSTALLATIONSNEA/CSNI/R(2006)5
Unclassified
REVIEW OF HIGH BURN-UP RIA AND LOCA DATABASE AND CRITERIAEnglish - Or. English

JT03218687
Document complet disponible sur OLIS dans son format d'origine
Complete document available on OLIS in its original format

NEA/CSNI/R(2006)52

NEA/CSNI/R(2006)5
ORGANISATION FOR ECONOMIC CO-OPERATION AND DEVELOPMENT
Pursuant to Article 1 of the Convention signed in Paris on 14th December 1960, and which came into force on
30th September 1961, the Organisation for Economic Co-operation and Development (OECD) shall promote policies
designed:−to achieve the highest sustainable economic growth and employment and a rising standard of living inMember countries, while maintaining financial stability, and thus to contribute to the development of theworld economy;− to contribute to sound economic expansion in Member as well as non-member countries in the process of
economic development; and− to contribute to the expansion of world trade on a multilateral, non-discriminatory basis in accordance withinternational obligations.
The original Member countries of the OECD are Austria, Belgium, Canada, Denmark, France, Germany, Greece,
Iceland, Ireland, Italy, Luxembourg, the Netherlands, Norway, Portugal, Spain, Sweden, Switzerland, Turkey, the United
Kingdom and the United States. The following countries became Members subsequently through accession at the dates
indicated hereafter: Japan (28th April 1964), Finland (28th January 1969), Australia (7th June 1971), New Zealand (29th
May 1973), Mexico (18th May 1994), the Czech Republic (21st December 1995), Hungary (7th May 1996), Poland (22ndNovember 1996) and the Republic of Korea (12th December 1996). The Commission of the European Communities takes
part in the work of the OECD (Article 13 of the OECD Convention).

NUCLEAR ENERGY AGENCYThe OECD Nuclear Energy Agency (NEA) was established on 1st February 1958 under the name of the OEEC
European Nuclear Energy Agency. It received its present designation on 20th April 1972, when Japan became its first
non-European full Member. NEA membership today consists of 27 OECD Member countries: Australia, Austria, Belgium,
Canada, Czech Republic, Denmark, Finland, France, Germany, Greece, Hungary, Iceland, Ireland, Italy, Japan, Luxembourg,
Mexico, the Netherlands, Norway, Portugal, Republic of Korea, Spain, Sweden, Switzerland, Turkey, the United Kingdom
and the United States. The Commission of the European Communities also takes part in the work of the Agency.
The mission of the NEA is:−

to assist its Member countries in maintaining and further developing, through international co-operation, thescientific, technological and legal bases required for a safe, environmentally friendly and economical use of
nuclear energy for peaceful purposes, as well as− to provide authoritative assessments and to forge common understandings on key issues, as input togovernment decisions on nuclear energy policy and to broader OECD policy analyses in areas such as energy
and sustainable development.
Specific areas of competence of the NEA include safety and regulation of nuclear activities, radioactive waste
management, radiological protection, nuclear science, economic and technical analyses of the nuclear fuel cycle, nuclear law
and liability, and public information. The NEA Data Bank provides nuclear data and computer program services forparticipating countries.
In these and related tasks, the NEA works in close collaboration with the International Atomic Energy Agency in
Vienna, with which it has a Co-operation Agreement, as well as with other international organisations in the nuclear field.
© OECD 2000
Permission to reproduce a portion of this work for non-commercial purposes or classroom use should be obtained through the Centre
français d’exploitation du droit de copie (CCF), 20, rue des Grands-Augustins, 75006 Paris, France, Tel. (33-1) 44 07 47 70, Fax (33-1) 46
34 67 19, for every country except the United States. In the United States permission should be obtained through the Copyright ClearanceCenter, Customer Service, (508)750-8400, 222 Rosewood Drive, Danvers, MA 01923, USA, or CCC Online: http://www.copyright.com/.
All other applications for permission to reproduce or translate all or part of this book should be made to OECD Publications, 2, rue
André-Pascal, 75775 Paris Cedex 16, France.

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COMMITTEE ON THE SAFETY OF NUCLEAR INSTALLATIONSThe NEA Committee on the Safety of Nuclear Installations (CSNI) is an international committee made up of
scientists and engineers. It was set up in 1973 to develop and co-ordinate the activities of the Nuclear Energy
Agency concerning the technical aspects of the design, construction and operation of nuclear installations insofar
as they affect the safety of such installations. The Committee’s purpose is to foster international co-operation in
nuclear safety amongst the OECD Member countries.
CSNI constitutes a forum for the exchange of technical information and for collaboration between organisations
which can contribute, from their respective backgrounds in research, development, engineering or regulation, tothese activities and to the definition of its programme of work. It also reviews the state of knowledge on selected
topics of nuclear safety technology and safety assessment, including operating experience. It initiates andconducts programmes identified by these reviews and assessments in order to overcome discrepancies, develop
improvements and reach international consensus in different projects and International Standard Problems, andassists in the feedback of the results to participating organisations. Full use is also made of traditional methodsof co-operation, such as information exchanges, establishment of working groups and organisation ofconferences and specialist meeting.
The greater part of CSNI’s current programme of work is concerned with safety technology of water reactors.
The principal areas covered are operating experience and the human factor, reactor coolant system behaviour,
various aspects of reactor component integrity, the phenomenology of radioactive releases in reactor accidents
and their confinement, containment performance, risk assessment and severe accidents. The Committee alsostudies the safety of the fuel cycle, conducts periodic surveys of reactor safety research programmes and operates
an international mechanism for exchanging reports on nuclear power plant incidents.
In implementing its programme, CSNI establishes co-operative mechanisms with NEA’s Committee on Nuclear
Regulatory Activities (CNRA), responsible for the activities of the Agency concerning the regulation, licensing
and inspection of nuclear installations with regard to safety. It also co-operates with NEA’s Committee on
Radiation Protection and Public Health and NEA’s Radioactive Waste Management Committee on matters ofcommon interest.

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FOREWORD

This document is intended to provide regulators, their technical support organizations andindustry with a concise review of existing fuel experimental data at RIA and LOCA conditions andconsiderations on how these data affect fuel safety criteria at increasing burn-up. It mostly addresses
experimental results relevant to BWR and PWR fuel and it encompasses several contributions fromthe various experts that participated in the CSNI SEGFSM activities. It also covers the information
presented at the joint CSNI/CNRA Topical Discussion on high burn-up fuel issues that took place on
this subject in December 2004.
The document was assembled by C. Vitanza and M. Hrehor upon suggestion from the CSNI
chair. It was presented at the SEGFSM meeting in April 2005. Comments made by the SEGFSM
members have been incorporated in the final version of the document. The document content does notnecessarily embody the opinion of the organizations to which each Group member belongs.

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TABLE OF CONTENTSFOREWORD .............................................................................................................................................. 7
1. INTRODUCTION AND MAIN CONCLUSIONS .............................................................................. 11
2. THE CABRI RIA DATABASE............................................................................................................ 12
3. THE NSRR DATABASE ..................................................................................................................... 13
4. OTHER DATABASES......................................................................................................................... 14
5. RIA FAILURE THRESHOLDS ........................................................................................................... 15
6. COMPARISON OF FAILURE THRESHOLDS FOR THE HZP CASE ............................................ 16
7. LOCA DATABASE, DUCTILITY TESTS.......................................................................................... 17
8. LOCA DATABASE, QUENCH TESTS .............................................................................................. 21
9. LOCA SAFETY LIMIT........................................................................................................................ 23
10. PROVISIONAL BURN-UP DEPENDENT CRITERION FOR Zr-4 ................................................ 25
11. MAIN CONCLUSIONS ..................................................................................................................... 26
REFERENCES.......................................................................................................................................... 39

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1.

INTRODUCTION AND MAIN CONCLUSIONSThis paper follows a discussion that took place in a joint CSNI-CNRA session in December
2004. The objective was to review the database gathered through testing in RIA and LOCA conditions
and to assess the status of development of new criteria taking into account the effect of burn-up.
During the discussion, which was moderated and led by Mr. J.M. Conde of the Spanish CSN,
presentations were made by the following authors:•

K. Valtonen of STUK, Finland, on the outcome of typical RIA enthalpy depositions and LOCA
cladding temperature transients at high burn-up for BWR and PWR cores•

J. Papin of IRSN, France, on the database, main results and plans of the Cabri test programme
on RIA•T. Fuketa of JAERI, Japan, on the database, results and plans of the NSRR RIA programme andon the outcome and plans of the LOCA quench tests at JAERI•F. Eltawila of the USNRC on the status, main results and plans of the LOCA programme
conducted at the US Argonne National Laboratory

An overview of the correlations published in the literature for the RIA failure threshold vs. burnup was presented by C. Vitanza of the NEA secretariat. On LOCA he presented a provisional LOCA
criterion based on ductility tests done with pre-hydrided, non-irradiated cladding specimens. The
above presentations are available at the NEA secretariat.
In preparing this paper, it was decided to organise and present the database from various
laboratories in a consistent manner, in order to facilitate the comparison of data from different sources.
While recognising that valuable results have been obtained also by other laboratories, the focus of thispaper is on those tests carried out in NEA member countries which are directly applicable to PWR and
BWR fuels. For what concerns PWRs, most of the current data concern Zr-4 cladding. The limited
available data for Zirlo, M5 and E110 cladding are also presented here.
After the CSNI-CNRA discussion of December 2004, new results have been generated
especially in the LOCA area. These refer in particular to the ductility tests carried out at Argonne onhigh burn-up cladding – the first high burn-up ductility data produced so far. Because of these recent
results, the conclusion of this paper on LOCA is slightly different from the conclusion of the CSNICNRA discussion.
The conclusions are as follows:•

On RIA, there is a well-established testing method and a significant and relatively consistent
database from NSRR and Cabri tests, especially on high burn-up Zr-2 and Zr-4 cladding. It is
encouraging that several correlations have been proposed for the RIA fuel failure threshold.
Their predictions are compared and discussed in this paper for a representative PWR case.

•

On LOCA, there are two different test methods, one based on ductility determinations and theother based on “integral” quench tests. The LOCA database at high burn-up is limited to bothtesting methods. Ductility tests carried out with pre-hydrided non-irradiated cladding show apronounced hydrogen effect. Data for actual high burn-up specimens are being gathered invarious laboratories and will form the basis for a burn-up dependent LOCA limit. A provisionalburn-up dependent criterion is discussed in the paper.

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2.

THE CABRI RIA DATABASEThe CABRI database consists of 14 tests, which were carried out in the decade 1993-2002. The
test conditions and range of most important parameters and fuel variants are given in the following
table:
Coolant

Liquid sodium

Coolant initial temperature

280ºC

Coolant flow rate

4 m/s

Coolant pressure~3 barTest fuelRe-fabricated segments, ~56 cm active length

Fuel origin

PWR reactors

Fuel type

UO2 (10 tests) and MOX (4 rods)

Cladding typeZr-4 (11 tests), M5 (2 tests), ZIRLO (1 test)

Pulse width

Variable, typically 9, 30 and 75 ms

Fuel enthalpy (at peak position)~200 cal/g at 30 MWd/kg, ~100 cal/g at 60
MWd/kg for UO2

Axial power shape

Peaked at mid height, peak factor=~1.2-1.3

Burn-up range

30-75 MWd/kg

Oxide thickness

10-130 µm

Fuel cladding failure occurred in four tests, whereas no failure was registered in the remaining
ten tests. Of the four tests that resulted in fuel failure, three had UO2 fuel and one MOX fuel. All
failures were with Zr-4 cladding, whereas no failure occurred in the three tests with M5 or ZIRLO
cladding.The three UO2 failures occurred at enthalpy below 80cal/g and on fuel that had a burn-up of~60 MWd/kg and significant corrosion (80 to 130µm). However, several other CABRI tests were run
at comparable burn-up and level of corrosion, without producing failure. Oxide spalling appears to be
the distinctive element that separates the failed and non-failed fuel in CABRI UO2 tests, in that all
three UO2 fuel rods that failed had spalling, whereas the rods that didn’t fail had uniform non-spalled
oxide.
One MOX test out of four resulted in fuel failure. The failure occurred on a fuel rod that hadburn-up of 55 MWd/kg and moderate corrosion, 50µm (without oxide spalling). This is the only
CABRI failure that took place on a non-spalled cladding. However, failure in this case occurred at
113cal/g, which is a rather high enthalpy level, where a failure may not be surprising, considering that
this level is near the upper envelope of all CABRI data.
Failures were not associated to a particular pulse width. The three UO2 fuel rod failures occurredat all three pulse widths that have been used in CABRI, i.e. 9, 30 and 75 ms.
A summary of the main parameters and outcome of the CABRI tests are listed in Table 1. The
achieved enthalpy – or the failure enthalpy for the test fuel rods that failed – is plotted as a function ofburn-up in Fig. 1.

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NEA/CSNI/R(2006)5
3.

THE NSRR DATABASEThe NSRR database consists of 26 tests of PWR fuel, 16 tests of BWR fuel and 28 additional
tests with non-commercial or ATR (MOX) fuel. The tests were carried out in the period 1989-2003.
The test conditions and the range of most important parameters is given in the following table:
Coolant

Water

Coolant initial temperature

20-85ºC

Coolant flow rate

Stagnant

Coolant pressure

1 bar

Test fuelRe-fabricated segments, ~10 cm active length

Fuel origin

PWRs, BWRs, ATR and JMTR (test reactor)

Fuel type

UO2, MOX in 6 cases

Cladding typeZr-4, MDA (1 test), ZIRLO (1 test) for PWR, Zr-2
for BWR fuelPulse width

4-5 ms in most cases

Fuel enthalpy (at peak position)~200 cal/g at 30 MWd/kg, ~120 cal/g at 60 MWd/kg

Axial power shape

Uniform

Burn-up range

0-65 MWd/kg

Oxide thickness

0-60 µm

Failure was registered in 5 out of 26 tests with PWR fuel and in 5 out of 16 tests with BWR fuel.
The five BWR failures belong to one specific set of fuel. Seven failures were registered in the other 28
tests. Failures appear to start at the cladding outer surface and propagate in a brittle mode up to ~40%of the wall thickness and then continue to propagate in a ductile manner [1].
There is currently no clear-cut explanation as to why some fuel rods failed and others didn’t atapparently similar conditions. In particular, there is one set of BWR rods that failed at 61 MWd/kg - 5
out of 5 for enthalpy between 70 and 86 cal/g [2] -, whereas the same fuel in the 40-56 MWd/kg range
did not fail even at 130-145 cal/g.
The NSRR failures occurred at moderate enthalpy both for Zr-2 and Zr-4, at enthalpy typicallyin the range 60-80 cal/g. Failures occurred also for cladding that had low corrosion and hydrogen pickup and that had no spalling, suggesting that at low coolant temperature as in the NSRR, the cladding
may become brittle even for a small amount of absorbed hydrogen. None of the MOX fuel rods tested
in NSRR failed, but burn-up was low in these cases [3]. A ZIRLO cladding with UO2 fuel failed, but
enthalpy was very high for the burn-up level of that test (120 cal/g at 58 MWd/kg) [4].
A summary of the main parameters and outcome of the NSRR PWR and BWR tests arepresented in Table 2. The achieved enthalpy – or failure enthalpy for fuel that failed – is plotted versus
burn-up in Fig. 2. (Additional NSRR tests were run after this note was written).

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4.

OTHER DATABASES

RIA tests at cold coolant conditions were performed in the SPERT and PBF facilities (USA) inthe 1968-70 and 1978-80 time-periods respectively. RIA tests on VVER fuel were also carried out in
the Russian IGR, BIGR and Gidra reactors (1984-2000)1.
There were six fuel failures in the SPERT tests, three of which occurred for burn-up of 1-4
MWd/kg and for enthalpy higher than 190cal/g, which is not surprising. However, a fourth low burnup failure occurred at a lower enthalpy, i.e. 147cal/g, a point that will be briefly commented below.
Only two SPERT tests were made at appreciable burn-up, i.e. 32 MWd/kg, resulting in fuel failures at<143 cal/g and at 85 cal/g respectively.
Within the PBF experimental series there are only six tests which are relevant for fuel failure
threshold assessment. They were run with fuel having burn-up between 0 and 5 MWd/kg. No failure
was registered in five tests run at high enthalpy, i.e. 185 cal/g or higher, whereas failure occurred in
one test at 140 cal/g.
These databases can thus be summarised as follows:•

There were two failures, one in the SPERT and one in the PBF series, which occurred forpractically fresh fuel at enthalpies of 140 and 147cal/g. This contrasts with 11 other SPERT and
PBF tests, 18 NSRR tests and 2 CABRI tests – altogether 31 tests showing a higher failure level
than in the above two tests.

•

More importantly, there was a SPERT failure at 85 cal/g for fuel burn-up of 32 MWd/kg, which
is a very low failure enthalpy compared with NSRR and CABRI data at similar burn-up. A
peculiarity of this test (and of the other SPERT tests at the same burn-up level that failed below
143cal/g) is that corrosion was as large as 65µm, which is very high compared with modern
BWR fuel. The commercial BWR fuel tested in NSRR, for instance, had oxide thickness of 2030µm at 60 MWd/kg.

The SPERT fuel was non-commercial, specially fabricated BWR fuel (with reduced cladding
wall thickness in some cases). The burn-up accumulation didn’t take place in a power plant but in atest reactor. Details of the base irradiation, e.g. on water chemistry control, are not known. The fuelwas fabricated ~40 years ago, i.e. with the technology and quality available at that time, and as alreadymentioned, the corrosion was abnormally high compared with current BWR fuel. Bearing in mind theimportance of corrosion for failure occurrence, consideration should be given to the above remarks
when using these data for assessing failure thresholds of modern fuel.
VVER fuel failures in the IGR and BIGR tests occurred at very high enthalpy, i.e. >~160 cal/g
in pressurised fuel rods at zero burn-up and in the burn-up range 47 - 49 MWd/kg. At 60 MWd/kg,
conservative failure thresholds of 130 and 140 cal/g have been derived for pressurised and respectively
un-pressurised rods.
In accordance with [38], there is no concern with reduced failure enthalpy vs. burn-up in VVER
fuel as long as the cladding corrosion and hydriding remain low. Further, the specific fuel pellet
design, e.g. the pellet central hole also has an effect on cladding strain and rupture threshold.

1

See additional Russian references number [36, 37, 38]14

NEA/CSNI/R(2006)55.

RIA FAILURE THRESHOLDS

Data renditions, correlations or models have been proposed by different authors for the RIA
failure threshold. They are illustrated in Fig.3 and briefly described below.
The Japanese NSC threshold was introduced in 1998 [5] as a data rendition for the NSRR,CABRI, SPERT and PBF failures plotted vs. burn-up. The threshold is placed below all fuel failure
data, except REPNa-1. No distinction is made between different fuels or between hot and cold RIA.
The NSC criteria also consider a fuel coolability limit, which is considerably higher than the failure
threshold (and which thus constitutes the actual RIA limit).
The fuel failure threshold proposed in [6] was developed primarily on the basis of CABRI data,
but it was well shown to also reproduce the NSRR data. It makes a distinction between ductile andbrittle cladding, with the latter having appreciably lower failure enthalpy. Oxide spalling (for Zr-4)
and low coolant temperature (for Zr-2 or Zr-4 cladding) are considered as important causes ofcladding brittleness. Burn-up and corrosion affect the threshold. There is no difference between UO2and MOX fuel or between BWR and PWR fuel. The differences between hot and cold coolant
conditions are the lower initial enthalpy and the lower cladding ductility at cold conditions (for Zr-2 or
Zr-4 cladding).
The KAERI correlation given in [7] is a fit of the CABRI, NSRR and SPERT failure data (nonfailed data were not considered). Burn-up and corrosion are the most important parameters, whereas
there is only a weak dependency on pulse width.
The fuel failure threshold proposed by the Swiss HSK makes a distinction between hot and cold
coolant conditions and between UO2 and MOX fuel [8]. The thresholds are based on the calculations
performed with the FALCON code. For the HZP PWR case, the HSK threshold is practically the sameas the one considered by EPRI. It is expressed vs. burn-up and is valid for corrosion up to 130µm. The
CZP (Cold Zero Power) BWR case and the MOX fuel case have significantly lower thresholds thanthe HZP (Hot Zero Power) UO2 case. The HSK criteria also have a fuel coolability limit, which isappreciably higher than the failure threshold.
The threshold proposed by SKI and other Swedish organisations [9] is derived on the basis of
FRAPCON-SCANAIR calculations. The calculations give different thresholds depending on transientconditions such as pulse width and coolant temperature. For PWR HZP transients, the thresholddepends on both burn-up and corrosion.
The Battelle threshold is based on calculations with the upgraded FRAPTRAN code [10]. As in
the previous case, the threshold depends on pulse and temperature conditions. Burn-up and especiallycorrosion affect the threshold level.
The threshold proposed in the 2004 Research Information Letter (RIL) is expressed as afunction of the corrosion level [11]. The threshold constitutes the lower envelope of the fuel failuresfrom the NSRR, CABRI (except REPNa-1), PBF and SPERT data. The RIL threshold envisages aslight difference between hot and cold RIA conditions.
IRSN note: For IRSN, the existing database (mainly CABRI and NSRR tests) pointed out theinadequacy of the existing (i.e. zero burn-up) criteria for high burn-up fuel. Waiting for theestablishment of suitable ones, the IRSN’s current approach to justify the safety of new fuel15

NEA/CSNI/R(2006)5
managements or cladding materials is based on a comparison of PWRs RIA transient and rod
characteristics, compared to the existing data base (including CABRI REP Na1). A temporary “safetydomain”, based on peak fuel enthalpy, pulse width, oxide thickness and oxide spalling (which isconsidered as a major cause of cladding brittleness) has been used for safety assessment.
6.

COMPARISON OF FAILURE THRESHOLDS FOR THE HZP CASE

The failure thresholds outlined above have been compared with each other for a PWR HZP case.
In order to have a common basis for the comparison, the same oxide thickness vs. burn-up is assumed
for all thresholds. The result of the comparison is shown in Fig.3. The oxide thickness considered isplotted versus burn-up in the upper part of the same figure.
All correlations predict a decrease of failure threshold with burn-up. In most cases the decrease
is of roughly 100-150 cal/g, but in the HSK (EPRI) case the decrease is smaller, approximately 50
cal/g.
At low burn-up, i.e. up to ~30 MWd/kg, most thresholds are in the range of 170-200 cal/g. The
RIL and NSC thresholds are however appreciably lower than that, typically from 100 to 150 cal/g
between 0 and 30 MWd/kg. The NSC correlation threshold [5] is a data rendition that also includes theold SPERT data, particularly the failure data point at 32 MWd/kg and ∆H=85 cal/g (i.e. H=103 cal/g
for a HZP transient). None of the other correlations, except the RIL threshold, accounts for thisparticular data point.
At high burn-up, the thresholds range from 60 to 100 cal/g in most cases. For the correlations
that acknowledge the oxide thickness dependence, a lower corrosion would result in a higher failure
threshold. The HSK correlation is the one that gives, for large corrosion at high burn-up, the highest
threshold for HZP RIA. However, no HZP tests have been carried out so far beyond enthalpy of 100
cal/g with highly corroded fuel (>80 µm) at high burn-up. This is partly because it is difficult toachieve a high enthalpy at high burn-up due to fissile depletion in the fuel. The CIP0-1 test in CABRI,
for instance, achieved 92cal/g at 75MWd/kg. (A NSRR test with 78MWd/kg fuel and 81µm corrosion
was carried out while this report was being completed. The fuel failed at 61cal/g). As a final remark on
predicted thresholds, no correlation accounts for the reported REPNa-1 failure enthalpy.
The failure thresholds are generally above the calculated enthalpies for RIA cases in actual
LWR cores. The shaded areas shown in Fig.4 give the conservative upper envelope of enthalpy
calculated for HZP RIA in PWRs [12] at low and high burn-up. As seen, there is a generally good
margin at low burn-up, whereas at high burn-up, the calculated enthalpies are very close or slightly
exceed the lower failure thresholds. Since the failure threshold is expected to increase for ductile andlow corroded claddings, limiting corrosion and retaining ductility (i.e. avoiding spalling) are importantmeasures to maintain a sufficiently large failure margin at high burn-up.

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7.

LOCA DATABASE, DUCTILITY TESTS
Hobson’s data, non-irradiated Zr-4

In most countries, the current LOCA safety limits are based on ductility tests. The most relevant
results were produced by Hobson and Rittenhouse [13a, b] using test specimens cut from Zircaloy-4
cladding tubes. Prior to cutting, the open tubes were subjected to two-side steam oxidation at various
temperature levels and for a varied amount of time. Two-side oxidation was used in order to be
representative of ballooned and burst tubes. After the high temperature oxidation phase, the sample
residual ductility was determined by ring compression tests. The ring compression results wereseparated into two categories, i.e. those showing some residual ductility and those showing brittle
behaviour. As expected, the “brittle” category consisted of those specimens that experienced longer
exposures and/or higher temperature in the steam oxidation phase. The Hobson and Rittenhouse test
conditions and specimen characteristics were as follows:
Test type
Slow compression of ring specimens
Crosshead speed
2.5 mm/min
Maximum deflection
3.8 mmTest temperature
23 - 150oC
Criterion for "failure"
Zero ductility
Specimen typeUn-irradiated Zircaloy-4 rings cut from oxidised tubes
Ring length
6.35 mm (1/4 in.)Ring outer diameter
10.72 mmRing thickness
0.686 mm
Specimen conditions
Tubes oxidised in high temperature and steam environment
Oxidation temperature
Constant, in the range 900 - 1300oC
Oxidation time
From 120 to 3600 s
Steam pressure
AtmosphericSurface conditions
Two-side oxidation
The outcome of the Hobson-Rittenhouse tests
and evaluations was that zero-ductility at 135°C isreached when the fractional thickness of combined[oxide +α -phase] layers relative to the wall thickness
attains the value of 0.44. As said earlier, this occurs
when the temperature is sufficiently high and thetime-duration is sufficiently long. It was found thatthe temperature-time limit for zero-ductility could be
expressed by the Baker-Just correlation for oxidation,
where the Equivalent Cladding Reacted (ECRBJ) was
set equal to 17%. This limit is thus given by2(1762/W)ּ exp (-Q/TK)ּ √τ = 172

Schematic of the ring compression test set-up(courtesy of F. Eltawila, USNRC)

where τ is time in sec. and TK is temperature in°K.

/1/

The Baker-Just equation for two-side oxidation is: ECR= (1762/W)·exp (-Q/TK)ּ √τ , where ECR is thepercentage of metal wall thickness that has oxidized, W is clad thickness (mm), τ is the oxidation time (s) and
TK is the oxidation temperature in°K. Q =11450°K, although the value 11500 has also been used.

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The additional database from ductility tests is presented in the following.
Hungarian data (AEKI), non irradiated Zr-4 and E-110
Ring compression tests have been carried out in VVER countries, including comparative
assessments of the Zr-4 and VVER E110 alloy. Hungarian data include approximately 50 ring
compression tests performed on Zr-4 and E-110 specimens oxidised at temperatures between 900 and
1200°C for time intervals ranging from 10 to 10000 seconds [14, 15]. The results are presented inTable 3. By setting the zero-ductility limit in correspondence with a crosshead displacement limit at a
rupture of ∆l/d<8%, and by plotting the brittle and ductile data on a time – temperature diagram (i.e.
Logt vs. 1/TK), one obtains the plot shown in Fig. 5a and 5b. As one can observe; the Zr-4 data arereasonably well separated by the 17% Baker-Just (BJ) limit, whereas for the E110 data; the limit is~100°C lower and corresponds to a ECRBJ of 8%.
Czech data (UJP), non irradiated Zr-4 and E-110, including pre-hydriding
Similar experiments were made in the Czech Republic [16, 17]. Also in this case there was acomparison between Zr-4 and E-110, but including a simulation of base irradiation corrosion made bypre-oxidising some of the specimens in autoclave. The data are shown in Table 4. With a ductilebrittle transition set at displacement ∆l/d≤ 8%, the results plotted in Fig. 6 are obtained (for asreceived samples). One can again observe that the Zr-4 brittle and ductile data are delimited by the
17% BJ line, whereas for E-110 the limit is ~8%. Specimen pre-oxidation resulted in a considerable
drop of the ductility limit. This is clearly shown by the data in Table 4. According to the received
information, however, pre-oxidizing in autoclave led to much higher hydrogen uptake than in actual
reactor conditions.
US data (ANL) for non-irradiated Zr-4, ZIRLO and M5 (including pre-hydriding)
The ring compression tests performed at the Argonne National Laboratory (ANL) include nonirradiated Zr-4, ZIRLO and M5 cladding [18-21]. A total of ~60 tests were made at temperatures of1000, 1100 and 1200°C. The results are listed in Table 5. Ductility was measured using the so-called“offset strain” as an indicator. Specimens with <2% offset strain were classified as brittle. The
outcome of the ANL test is depicted in Fig.7a, b, c. The data show practically no difference between
the three alloy types in terms of zero-ductility limit 3.
One can observe from Fig.7 that the ECRBJ=17% remains a valid and actually conservative limit
at 1000 and 1100°C, whereas there is evidence that brittleness (at RT) can occur somewhat below that
limit for the 1200°C case. At 1200°C, in fact, the ANL data show a ductile-brittle transition after ~100
seconds (ECRBJ=12%). However, further ANL tests - as well as Hobson’s tests - showed that ductility
increases significantly when ring compression is carried out at 100ºC or 135ºC instead of RT, even for
ECRBJ >17% at 1200ºC [31].

3

When the offset strain is plotted as a function of time for each of the three temperature levels, one can actuallyinfer that ZIRLO and M5 behave generally better than Zr-4 in that they retain more ductility – on theassumption that offset strain is a reliable indicator of ductility. This has been shown for instance in [22].

18NEA/CSNI/R(2006)5

Ring compression tests (at 135ºC) of pre-hydrided, non-irradiated samples were also carried out
at ANL. The main outcome was that [31]:−Ductile-brittle transition occurs at ECRBJ = 6.5% for H≥600ppm (15x15 Zr-4)−Ductile-brittle transition occurs at ECRBJ = 9.5% for H≥400ppm (15x15 Zr-4)−For 17x17 Zr-4, ductile-brittle transition occurs at ECRBJ <13% for H≥ 300ppm
French data (CEA) for non-irradiated Zr-4 and M5 cladding, including pre-hydriding
In addition to the Czech and ANL data described above, the French CEA has also carried out
experiments with non-irradiated, pre-hydrided cladding [23]. Both Zr-4 and M5 cladding were tested
using various methods to deduct the post-oxidation residual ductility, including the ring compression
method. Hydrogen concentration was 600ppm in Zr-4 and 200ppm in M5, reflecting the different levelof corrosion expected for these two alloys at high burn-up. The outcome of the CEA tests is shown in
Fig. 8. As-received, Zr-4 and M5 exhibited similar behaviour. Hydrogen addition had a marked effecton Zr-4 and a more moderate effect on M5 due to the lower H-content. Zero-ductility occurred at
ECRBJ between 5 and 10% for Zr-4 with 600ppm hydrogen4 and between 10-15% for M5 with
200ppm hydrogen. One can infer from this that the zero-ductility ECR drops roughly by a factor of 2.5
at 600 ppm (17 →~7%) and by a factor of ~1.3 (17→13%) at 200ppm hydrogen. This can be
expressed as follows:
ECRBJ(zero-ductility)= 17% ⋅[1-H/1000] (from pre-hydrided Zr-4 and M5 samples, H≤600ppm). /2/
US data (ANL) for high burn-up Zr-4 cladding
While this note was being compiled, the NRC published new ring compression results for highburn-up Zr-4 cladding [24]. These are the first ductility data obtained from actual fuel cladding
material. Ring compression was performed at 135ºC. The fuel was retrieved from the H.B. Robinson
reactor at high burn-up (~64 MWd/kg) and had a base-irradiation oxide layer of ~70µm and H-content
of ~700ppm. The cladding was de-fuelled and subjected to two-side oxidation, according to the timetemperature scheme shown in the upper left diagram of Fig. 95. The corresponding effective
temperature and time6 are given in the table on the same figure, together with the residual ductility
results. The data, plotted on the usual time-temperature diagram as shown at the bottom-right part of
Fig. 9, indicate that the ductile-brittle transition is somewhat below the 17% line, i.e. at about 11% BJ,
which in terms of temperature means a ~80ºC lower claddind temperature limit at high burn-up. It
should be noticed that these first ANL tests do not account for oxide cracking occurring in theballooning region, which might reduce the oxide protective effect. Future ANL tests intend to address
this point. Based on these data the LOCA limit decrease would be:
ECRBJ(zero-ductility)= 17% ⋅[1-0.5⋅H/1000](Based on H.B. Robinson cladding, H=~700ppm)/2/

4

Fig 8 plots the results in term of measured ECR and not the conventional Baker-Just ECR. Further, the timeduration of the transient is not provided.

5

The upper right diagram shows that the burn-up effect is comparable with the ductility change from 135°C to
RT for non-irradiated specimens.

6

The effective temperature and time correspond to a constant temperature transient, equivalent to the actual
transient in terms of final ECR.

19

NEA/CSNI/R(2006)5
Russian data for Zr1%Nb cladding (fresh and high burn-up)7
Studies such as those conducted at the Russian Kurchatov Institute on the understanding of hightemperature oxidation of Zr-Nb alloys have given important contributions to the understanding of the
VVER E-110 alloy as compared to other Zr-1%Nb alloys [32]. As stated in [33]; the following generalobservations may be made on the basis of Russian experimental data:•

The irradiation inhibits the breakaway oxidation tendency on the outer surface of the E110
cladding;•

The tendency towards oxide spalling during high temperature oxidation is supplemented by thetendency towards an increase in oxidation rate on the cladding inner surface. In accordance withthese data, it may be assumed that the contamination of the cladding inner surface by fission
products is responsible for these effects.
The database characterising the residual ductility micro-hardness and hydrogen content in theoxidized irradiated cladding as a function of ECR has shown that:•

In accordance with the postulated relationship between the oxygen concentration in the prior βphase, micro-hardness, and the oxygen induced zero ductility threshold, this corresponds to
8.3% measured ECR (i.e. 13% calculated with a Russian conservative correlation);•

The combination of the oxygen induced and hydrogen induced embrittlement of the E110
irradiated cladding leads to the reduction of the zero ductility threshold down to 6.5% measured
ECR at the fast/fast combination of heating and cooling rates.

The investigations performed demonstrated that the oxidation behaviour of niobium-bearing
alloys is very sensitive not only to alloying components but also to impurity composition. The analysisof the E110 problems in this context has shown that different methods are used for PWR and VVER
zirconium alloy production. In fact, sponge Zirconium is used to produce the PWR alloys Zr-4, M5
and ZIRLO, whereas a mixture of iodide and electrolytic Zirconium is used for the fabrication of the
VVER E110 alloy. This difference leads to differences in impurity composition.
Ring compression tests carried out after steam oxidation at 1100ºC on modified E110 cladding
produced from sponge Zr confirmed that the zero ductility threshold increased up to 19% ECR [33],due to low hydrogen uptake. The post-oxidation zero-ductility threshold of sponge E110 alloy wascomparable to the ANL data. These results have prompted the Russian fuel manufacturer to movetowards sponge zirconium alloy, as well as to modify the finish of the cladding surface (two-side
polish and no etching).
The performed research enabled to make the conclusion that the oxidation behaviour and residual
ductility of the E110 alloy are very sensitive to cladding micro-chemical composition and surface
finishing. The use of sponge zirconium for the cladding fabrication results in a significant reduction ofoxidation rate especially in the range 900-1000ºC and in an increase of the zero ductility threshold.
However, an increase of oxidation rate of sponge vs. standard E110 has been registered at 1200°C.

7

The text of this subsection was provided by experts of the Kurchatov Institute.

20

NEA/CSNI/R(2006)5
Summary of ductility tests−ANL results show no significant difference between as received Zr-4, ZIRLO and M5−CEA data show the same for Zr-4 and M5−AEKI and UJP tests show a lower E-110 limit compared with Zr-4 (ECRBJ =8% for E-110 vs.
17% for Zr-4; i.e. a ~100ºC lower tolerable temperature)−UJP, CEA and ANL non-irradiated samples exhibit a significant hydrogen effect. ANL data at
135ºC show a ductile-brittle transition at ECRBJ =6-7% for H≥ 600ppm (15x15 Zr-4) and at<13% for H≥ 300ppm (17x17 Zr-4)−However, preliminary ANL high burn-up data show that the ductile-brittle transition occurs at alevel that is only moderately below 17% (at 11±2%). The reason for the difference between
these and pre-hydrided cladding results is currently being debated. The tests do not simulate theoxide cracking occurring in the ballooning area.
8.

LOCA DATABASE, QUENCH TESTSJAERI tests with as received Zr-4 (zero burn-up)

The experimental basis for the LOCA criteria inJapan is constituted by quench tests conducted on
cladding tube segments. These tests were performed
at JAERI in the 80s [25]. The tubes were ballooned,
ruptured, subjected to high temperature two-sideoxidation and then quenched. During quenching
there was full constraint, i.e. axial contraction wasprevented during cooling. Full constraint represented
a bounding case for the actual fuel assembly
condition, since a reliable assessment of fuel-gridinteraction and of consequent loading was difficult to
make.

Load cell

Restraint load
control system
Grabbing device

SteamoutletQuartztube
Test rod
Infrared
furnace

Approximately 50 experiments were carried out
Equipped on
on non-irradiated Zr-4 cladding tubes, which served
tensile testingmachineas a basis for the Japanese LOCA criteria. The results
were divided into two categories, depending onwhether the tube had survived the quenching or not.
Flooding water inlet
Steam inlet
Fig. 10 provides a plot of the experimental results,
showing the data points for the failed and non-failed
tubes on a time vs. temperature diagram. As one can
Schematic of the quench test set-up(courtesy of T. Fuketa, JAERI)
observe, the Baker-Just 15 % ECR line is a conservative rendition of the results. As most phenomena
under LOCA conditions, including rupture, secondary hydriding from inner surface and axial loading8
during the quench, are take into account in the quench tests, the 15% ECR criteria was adopted as the
LOCA failure and coolability criterion in Japan.
8Considerations have however been expressed regarding the possible presence of additional transversal loads.

21

NEA/CSNI/R(2006)5
JAERI tests with non-irradiated, pre-hydrided cladding
Quench tests with pre-hydrided cladding have been carried out at JAERI with the purpose ofdetermining the separate effect of hydrogen on quench failure. Tests carried out with H-content up to
600-800 ppm and full constraint show that the quench failure limit decreases substantially, typicallyfrom 17 to 7% [26]. This is consistent with the CEA ductility tests with pre-hydrided cladding. On theother hand, JAERI results show that the hydrogen effect is significantly smaller when the constraint islimited to 50 kg [27]. The difference between the full constraint and the 50 kg constraint is shown in
Fig. 11. As for the case with a 50 kg constraint, it should be noticed that the scale is very sensitive andthat all failure data points (as well as the two lines drawn in the figure) are above the BJ 17% limit.
JAERI and ANL tests with high burn-up cladding
Quench tests with high burn-up BWR and PWR fuel have recently been carried out at JAERI
and ANL. At JAERI, three Zr-4 cladding tubes were cut from a PWR fuel rod having a burn-up of
48MWd/kg and moderate corrosion, i.e. 15-25µm. The hydrogen content was in the range of 120210ppm. The tubes de-fuelled and filled with alumina pellets. They were then pressurized andsubjected to steam oxidation at temperature between 1030 and 1180ºC for a time ranging from 120 to2195 seconds. Ballooning and burst occurred during the heat up phase. After holding at hightemperature, the tubes were quenched while maintaining an axial tensile load of about 50 kg. Two of
the tubes failed with a transversal crack during quenching, whereas the other specimens survived thequenching [27]. The choice of a 50 kg axial load in JAERI tests is based on evaluations and is still
discussed [28, 29].
At ANL, two quench tests were done on a fuelled BWR rod segment having burn-up of 56
MWd/kg and 10µm oxide (~70 ppm H). There was no constraint on the segment during the quench
test. The fuel segment was kept at 1204ºC for 5 minutes. The cladding did not fail during quench, but
the sample that was not embedded in epoxy failed during handling [30].
The results of the JAERI and ANL tests described above are summarized in Table 7 and in
Fig.12. The latter clearly shows that the JAERI results are very well reproduced by the 17% ECRBJ
limit, in that the two failed tests are above this limit whereas the 4 non-failed ones are below it. The
ANL data are just above the BJ 17% ECR line, but there was no failure during quenching in these twocases. One should however notice that these ANL tests were done without constraint - and that there
was subsequent failure during handling at cold conditions in one case.
Russian tests
Considerable amount of quench tests has been done on both non-irradiated and irradiated VVER
fuel in the Russian Federation [31]. These quench tests showed that the BJ 17% ECR limit holds also
for the E-110 alloy. More general information on VVER fuel safety criteria can also be found in [34].
Summary of quench tests– For as received Zr-4 samples, JAERI quench tests with full constraint give similar LOCA limit

as the ductility tests (ECRBJ =15% compared with 17%)
– Russian quench tests without constraint show that the BJ 17% limit holds also for the E-110
alloy, which contrasts with ductility test results (at RT)– JAERI quench tests with full constraint show a failure limit reduction by a factor of 2.5 (17→7%) due to pre-hydriding, which is consistent with ductility test results– JAERI (50 kg constraint) and ANL (no constraint) quench tests with high burn-up claddingshow no quench failure below 17% ECRBJ. Consensus on constraint level would be desirable.

22

NEA/CSNI/R(2006)5

oxidation temperature, OC

LOCA SAFETY LIMIT

1400 1300 1200

As shown in the previous sections, a straightline in a temperature-time chart (1/T vs. Log Time)
expresses the current LOCA limit. In most countries
this limit is defined by the so-called zero-ductility
boundary as derived from ring compression tests,
which are done at cold conditions after the hightemperature steam corrosion phase. The boundary
separating the ductile from the brittle cladding data
points defines the zero-ductility-based LOCA limit.
In other countries and notably in Japan thelimit is defined by a different criterion, i.e. bywhether fuel simulators survive or fail during
quenching after the high temperature corrosion
phase. The simulators are constituted of fuelsegments or cladding tubes filled with alumina
pellets. The boundary separating the non-failed
from the failed cladding data points defines thequenching-based LOCA limit.

1100

1000

900

800

10000

Zero ductility

Oxidation time, s

9.

Zero ductility
threshold for Zr-4
(Hobson- Rittenhouse
results)

1000

100

10

6

7

8

9

10000/TK, OK-1

It so happens that (at least for Zr-4) the two approaches lead to virtually coincidental limits atzero burn-up if the quenching tests are done under full constraint conditions. The Baker-Just (BJ)
equation, which in the past was used to describe high temperature corrosion, was also found to be aconvenient means of expressing both the zero-ductility limit derived from ring compression tests andthe failure limit derived from quench tests, apart from a very minor difference in a parameter setting(15 vs.17%). These limits are given by the two equations:
17=(1762/W)·exp(-Q/TK) √τ “Zero-ductility limit”/3/

15=(1762/W)·exp(-Q/TK) √τ/4/“Quench limit”

The graphic form of the BJ 17% limit is given by the line in the diagram shown above(for Hobson’s specimens, i.e. W=0.686mm). The BJ 15% limit is only ~20ºC below this line.
One should notice that the Zr-4 17% ECR criterion implies the use of the BJ equation, regardless of how good this correlation is for oxidation. The criterion doesn’t say that the actual cladding
oxidation should be below 17%, but that temperature and time condition should be below thoseleading to a calculated 17% ECR if the BJ correlation was used. Had one chosen another equation
instead of BJ to calculate ECR, the zero-ductility limit would not have been expressed by the 17%
ECR, but by some other value9 (while the temperature-time limit for zero ductility would obviously
remain what it is).
Using measured ECR or using different ECR correlations may bring confusion on what thesevarious ECRs mean in terms of temperature and time (these being the real LOCA parameters),
9

Had one chosen the Cathcart-Pawel equation, for instance, the limit would have been 13% - in order to obtainthe same time-temperature limit as the 17% BJ limit.

23

NEA/CSNI/R(2006)5
especially when different alloys having different oxidation and ductility properties are considered.
Further, at high burn-up, base irradiation ECR and high temperature ECR can be improperly mixed up.
The use of a commonly acknowledged reference, such as the BJ equation, is therefore advisable.
Alternatively, one can express the LOCA limit directly in terms of temperature and time – which areunambiguous parameters - instead of ECR. The two following tables give the main outcome of thezero and high burn-up investigations.
Summary of the LOCA tests
ALLOY

ZERO BURN-UP, AS RECEIVED

ZERO BURN-UP, PRE-

TYPE

HIGH BURN-UP

HYDRIDED
Ductility
Tests

Quench, fullQuench,constr.
0-50kg constr.

Zr-4

BJ 17%

BJ 15%--------

BJ ~7% for BJ ~7% for BJ 17% for Moderate burn- -------H≥ 600ppm H≥ 600ppm H≥ 600ppm up effect, limit
50 kg constr. at 11% BJ

ZIRLO

BJ 17%----------------------------------------------------------------M5

BJ 17%----------------

BJ ~13% for -------H≥ 200ppm--------------------------------

E-110

(∗)
BJ 17%--------

BJ 17%--------------------------------

BJ 17%

E-110sponge-

Quench, full Quench, 0- Ductility Tests Quench, full Quench,
constr.
50kg constr
constr. 0-50kg const.

--------

BJ 17%(0-50kg
constraint)

Ductility tests reported here were done by means of ring compression. The database consists mostly of tests
done at RT. ANL has observed a marked ductility enhancement for tests done at 100ºC and 135ºC(∗) BJ 13% from Russian data and 8% from [14-17]
Main conclusions of the LOCA tests

CONCLUSIONS- No burn-up effect
Limit is 17% regardless ofburn-up- Moderate burn-up effect
Limit is 11% BJ at high burnup

Worst case
- Large burn-up effect
Limit is 17%⋅[1-H/1000] BJ
H in ppm, max. H=600ppm

EXPERIMENTAL BASIS- Quench tests, 50 kg constraint, pre-hydriding to ~1000ppm (JAERI)- Quench tests, 50 kg constraints, high burn-up cladding (JAERI)
Note: Only Zr-4 has been tested under these conditions- High burn-up fuel, ANL HBR tests at 1200°C), without oxide cracking
Note: Only Zr-4 has been tested under these conditions- Ductility tests, pre-hydrided Zr-4 samples (UJP, CEA, ANL)
- Quench tests, full constraints, pre-hydriding to ~1000ppm (JAERI)
Note: Mostly Zr-4 data24

NEA/CSNI/R(2006)5

10.

PROVISIONAL BURN-UP DEPENDENT CRITERION FOR Zr-4

In [22] an attempt was made in order to show how existing data could form the basis for aprovisional burn-up dependent LOCA criterion. This was, among other things, aiming to provide an
initial answer to a question posed by the NEA Committee on Nuclear Regulatory Activities (CNRA)
to CSNI on the issue of burn-up dependent fuel criteria. The main points of the proposed approach arereviewed below.
Irradiation damage. There are indications that the irradiation damage as such does not haveimportant effects on post-LOCA ductility. Consequently, an irradiation dose is not considered as aburn-up effect in this analysis, neither for Zr-4 nor for other alloys.
Hydrogen pick-up. Data on the separate effect of hydrogen have been presented above. They
were obtained by ring compression tests and quench tests of non-irradiated, pre-hydrided samples. An
important hydrogen effect is shown by ring compression tests and by quench tests with full constraint,
up to a hydrogen content of 600ppm. Up to this level, the LOCA limit expressed in terms of ECRBJ is
reduced from 17 to ~7%, i.e. by a factor of 2.5. No further reduction is apparent for hydrogenconcentration exceeding 600ppm. This means that the LOCA limit expressed as the ECR during thehigh temperature oxidation is:
ECRLIMIT, BJ =17%⋅[1-H/1000] (Worst case. H<600ppm, for higher content use H=600).

/5/where H is in ppm. This should be taken as the worst case, compared with the more moderate burn-up
effect derived from actual high burn-up fuel cladding (but without oxide cracking).
Wall thickness reduction. The pre-existing corrosion reduces somewhat the cladding resistanceto LOCA because the metal wall thickness is reduced. This is accounted for by the factor [1ECRBASE/100] entering as multiplier in the previous equation, where ECRBASE is the oxidised metal as
percentage of wall thickness. This term is much weaker than the previous H pick-up term, and can beneglected in most cases.
As mentioned earlier, the ECR reduction can be expressed as a temperature limit reduction, i.e.
(for e.g. W=0.576mm):
T300s = Q / (8.04 - ∆ox) – 273.

/6/

where ∆ox= ln [1-H/1000]. The value T300s is the temperature corresponding to the time 300 seconds
on the LOCA limit line. In other words, T300s is the maximum tolerable temperature for a hightemperature oxidation phase lasting for 300 seconds, and this tolerable temperature decreases withcorrosion and Hydrogen content as given by Eq/6/.10 It can easily be shown that the amount of
temperature decrease vs. H (vs. burn-up) is practically invariant for any type of LOCA transient.
The predictions given by Eq./5/ and /6/, which are based on current data represent a “worst
case” for burn-up effect, are shown in Fig 13 and 14. The plots of Fig.13 relate to a relatively large
10This would be for an ideal transient where the temperature would be brought to a high value and kept constantfor a certain time period, such as in most ductility tests. There are means to “convert” from an actual LOCA
transient, where temperature varies, to an equivalent, constant temperature transient, as explained in [22]. For
any cladding thickness the term (7.49-lnW) should be used in /6/ instead of the term 8.04

25

NEA/CSNI/R(2006)5
corrosion vs. burn-up (also shown in the figures) and to a relatively large hydrogen pick up fraction,i.e. 10ppm H/µm oxide. By comparison, also the “total oxide” criterion – where the LOCA limit is set
as: ECRLIMIT, BJ = 17% - ECRBASE is also plotted. As one can notice, the “worst case” results in an ECR
decrease from 17 to 6% BJ at high burn-up (for large corrosion). This corresponds to a temperature
limit decrease by ~160ºC, which mostly occurs at 30-50 MWd/kg, as the plot of Fig 13 shows. In otherwords, the “worst case” criterion would imply a ~160ºC lower tolerable temperature at high burn-up
compared with zero burn-up. The effect of corrosion is shown in Fig.14. Not surprisingly, the LOCA
limit vs. burn-up does not decrease much when corrosion is small.
11.

MAIN CONCLUSIONS•On RIA there is a well-established testing method and a significant and relatively consistent
database from NSRR and Cabri tests, especially on high burn-up Zr-2 and Zr-4 cladding. As
observed in the CSNI-CNRA discussion and in this note, it is questionable whether other very
old data can be considered representative for modern claddings. It is encouraging that several
correlations have been proposed for the RIA fuel failure threshold. Their predictions arecompared and discussed in this paper for a representative PWR case.

•

On LOCA there are two different test methods, one based on ductility determinations and theother based on “integral” quench tests. The LOCA database at high burn-up is limited for bothtesting methods. Ductility tests carried out with pre-hydrided non-irradiated cladding show apronounced hydrogen effect. Tests with actual high burn-up specimens - but without oxide
cracking - show a moderate burn-up effect. However, this should be taken with caution, because
recent separate effect tests done at CEA indicate that cooling rate can affect the results.
Provisional burn-up dependent criteria are discussed in this paper. However, more high burn-up
data is needed in order to arrive to firmer conclusions on a burn-up dependent LOCA limit.

26

NEA/CSNI/R(2006)5

Table 1. The CABRI Database, Main Parameters
Test anddate
Na-1(11/93)Na-2(06/94)
Na-3(10/94)Na-4(07/95)Na-5(05/95)
Na 6(03/96)
Na 7(01/97)Na-8(07/97)Na-9(04/97)Na-10(07/98)Na-11(06/00)Na-12(12/00)CIP0-1(11/02)
CIP0-2(11/02)Rod and Burnup
GRA 564 MWd/kg
BR3
33 MWd/kg
GRA 5
53 MWd/kg
GRA 5
62 MWd/kg
GRA 5
64 MWd/kg
MOX
47 MWd/kg
MOX
55 MWd/kg
GRA 5
60 MWd/kg
MOX
28 MWd/kg
GRA 5
62 MWd/kg
M5
63 MWd/kg
MOX
65 MWd/kg
ZIRLO
75 MWd/kg
M5
77 MWd/kg

Pulse
9.5
9.1
9.5
75

ms

31

95
(at 1.2 s)
105(at 0.4 s)
125 at 0.66s
165 at 1.2 s
125 at 0.48s
175 at 1.20s
106(at 0.4 s)
197 at 0.5 s
241 at 1.2 s
107(at 1.2 s)
104

62

106

32

98

28

89

9.5
35
40
75
34
31

Corrosion µm

Energy dep.
cal/g
110
(at 0.4 s)
211(at 0.4 s)
120)

80
spalled4(at 0.4 s 40
80
no spalling
20
35
50
130
lim. spalling< 20
80
spalling
15
80
no spalling80
no spalling
20

Results and observations
Brittle failure at HF = 30 cal/g.
Fuel dispersal (6 g)
No failure HMAX = 199 cal/g
Max. strain: 3,5%, FGR: 5.5%
No failure HMAX = 124 cal/g
Max. strain: 2% FGR: 13.7%
No failure HMAX = 85 cal/g
Max. strain: 0.4% FGR: 8.3%
No failure HMAX = 108 cal/g
Max. strain: 1% FGR: 15.1%
No failure HMAX = 133 cal/g
Max. strain: 3.2%, FGR: 21.6%
Failure at HF= 113 cal/g
Strong flow ejection
FailureHF≤ 82 cal/g,
HMAX = 98 cal/g, no fuel dispersal
No failure HMAX = 197 cal/g
Max. strain: 7.4%, FGR: ~34%
Failure at HF = 81 cal/g,
HMAX = 98 cal/g, no fuel dispersal
No failure HMAX = 93 cal/g
Max strain ~0.5%
No failure HMAX = 103 cal/g
No failure

HMAX = 90 cal/g

No failure

HMAX = 81 cal/g

Table 1.

Fig 1. CABRI DATABASE

Fuel Enthalpy, cal/g

250
200
150

FailedNon-failed

100
50
0
0

20

40

60

Burn-up, MWd/kg

27

80

100

NEA/CSNI/R(2006)5
Table 2. The NSRR Database, Main Parameters11
PWR fuel

Burn-upCorrosion

Enthalpy

FailureJMTR fuelBurn-upCorrosion

Enthalpy

Failure1989-2003

MWd/kgµm

cal/g

enthalpy

1989-2003

MWd/kgµm

cal/g(enthalpy)

MH-1

39

4

47

NF

JM-1

22

2

92

NF

MH-2

39

4

55

NF

JM-2

27

2

84

NF

MH-3

39

4

67

NF

JM-3

20

2

132

NF

GK-1

42

10

93

NF

JM-4

21

2

177

FAILURE

GK-2

42

10

90

NF

JM-5

26

2

167

FAILURE

OI-1

39

15

106

NF

JM-6

15

2

156

NF

OI-2

39

15

108

NF

JM-7

13

2

146

NF

HBO-1

50

43

73

60 cal/g

JM-8

20

2

160

NF

HBO-2

50

35

37

NF

JM-9

25

2

160

NF

HBO-3

50

23

74

NF

JM-10

21

2

200

NF

HBO-4

50

19

50

NF

JM-11

31

2

160

NF

HBO-5

44

60

80

77 cal/g

JM-12

38

2

180

FAILURE

HBO-6

49

30

85

NF

JM-13

38

2

150

NF

HBO-7

49

45

88

NF

JM-14

38

2

160

144 cal/g

TK-1

38

7

126

NF

JM-15

30

2

150

NF

TK-2

48

35

107

60 cal/g

JM-16

38

2

140

NF

22

2

160

NF

TK-3

50

10

99

NF

JMH-1

TK-4

50

15

98

NF

JMH-2

22

2

200

NF

TK-5

48

20

101

NF

JMH-3

30

2

220

205 cal/g

TK-6

38

15

125

NF

JMH-4

30

2

170

NF

TK-7

50

30

95

86 cal/g

JMH-5

30

2

220

189 cal/g

TK-8

50

10

65

NF

ATR-1

20

15

80

NF

TK-9

50

10

99

NF

ATR-2

20

15

110

NF

TK-10

46

10

86

NF

ATR-3

20

15

120

NF

OI-10 MDA

60

27

104

NF

ATR-4, -5

20

15

140

NF

ZIRLO

58

28

157

120 cal/g

ATR-6

30

-

85

NF

Burn-up
MWd/kgCorrosionµm

Enthalpy
cal/g

Failureenthalpy

BWR fuel
1999-2002

Burn-up
MWd/kgCorrosionµm

Enthalpy
cal/g

TS-1

26

6

55

NF

FK-4

56

22

140

NF

TS-2

26

6

66

NF

FK-5

56

22

70

NF

TS-3

26

6

88

NF

FK-6

61

25

131

70 cal/g

TS-4

26

6

89

NF

FK-7

61

25

129

62 cal/g

OI-11

BWR fuel
1989-1998

11(7)

Failureenthalpy

TS-5

26

6

98

NF

FK-8

61

25

65

NF

FK-1

45

16

130

NF

FK-9

61

25

90

86 cal/g

FK-2

45

18

70

NF

FK-10

61

25

103

80 cal/g

FK-3

41

24

145

NF

FK-12

61

25

89

72 cal/g

A fuel rod (JMN-1, 22 MWd/kg, 2µm oxide) failed at 150 cal/g. This rod, however, had a reduced gap, 100µm (as fabricated) instead of
the normal design value of 190µm like all other rods in the JM series

28

NEA/CSNI/R(2006)5

FIG. 2 NSRR DATABASE
NSRR, PWR FUEL

Fuel Enthalpy, cal/g

250
200
150

Failed100

Non-failed

50
0
0

20

40

60

80

100

Burn-up, MWd/kg

NSRR, PW R+BW R FUEL

Fuel Enthalpy, cal/g

250
200
150

FailedNon-f ailed

100
50
0
0

20

40

60

80

100

Bu r n-u p, M Wd/k g

NSRR, ENTIRE DATABASE

Fuel Enthalpy, cal/g

250
200
150

Failed100

Non-f ailed

50
0
0

20

40

60

Burn-up, M Wd/k g

29

80

100

Oxide thickness, microns

NEA/CSNI/R(2006)5

250

Fuel Enthalpy [cal/g]

200

160

Oxidation vs. burn-up

140
120
100
80
60
40
20
0
0

20

40

60

80

100

Burn-up, Gwd/t

150
100
50
0
0

20

40

60

80

Burnup [GW/t]Our, ductile
HSK & EPRI

Our, brittle
US RIL 0401

KAERI
PNL Battelle

NSC Japan
SKI & others

Fig.3. Comparison of various RIA failure thresholds for a PWR HZP case with relatively large oxidation.
The assumed oxidation vs. burn-up is shown in the upper right figure.

250

Fuel Enthalpy [cal/g]

200
150
100
50
0
0

20

40

60

80

Burnup [GW/t]Our, ductile
HSK & EPRI

Our, brittle
US RIL 0401

KAERI
PNL Battelle

NSC Japan
SKI & others

Fig.4. Most thresholds are within the upper shaded area. The two lower shaded zones represent the envelope ofconservative calculations for actual PWR cores [12]. As seen, the core enthalpy can exceed the lower failure
thresholds – i.e. for cases with large oxidation and brittle cladding.

30

NEA/CSNI/R(2006)5

Table 3. LOCA, HUNGARIAN DATABASE (AEKI) Ref. [14, 15]
Zr-4 0.625mm
No.

Temp.ºC

Time s.

ECR [%]

1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24

900
900
900
900
1000
1000
1000
1000
1000
1000
1000
1100
1100
1100
1100
1100
1100
1200
1200
1200
1200
1200
1200

300
1000
5000
11000
87
464
2600
3300
4090
7270
11360
27
102
398
900
1500
3000
10
40
163
367
790
1100

2.3
3.5
5.9
7
2.9
6
12.3
15.2
20.1
43.6
77.3
2.8
5.4
10.1
15.2
19.5
26.8
3.5
5.8
10.5
15.4
21.9
25.7

Zr 1%Nb 0.650mm

H cont.
[ppm]

2.6
2.3
1.6
0.4
1.2
0.9
0.6
8.0
997.
1854.
110.
0.6
1.4
2.6
1.6
2.1
5.5
1.6
0.7
0.9
1.4
4.9
1.1

∆ l/d [%]

Temp.ºC

Time s.

ECR [%]∆ l/d [%]

49.63
40.7
15.4
13.7
45.34
19.1
15.4
5.83
4.33
2.7
1.5
52.7
39.9
20.7
10
6.4
4.65
38.17
17.07
8.46
5.42
3.75
3.75

H cont.
[ppm]

900
900
900
900
900
900
1000
1000
1000
1000
1000
1000
1100
1100
1100
1100
1100
1100
1200
1200
1200
1200
1200
1200

350
1000
3000
7000
11000
14000
100
700
1200
1800
3600
6000
19
133
704
1500
2400
5000
7
49
167
380
646
1205

1.61
3.59
8.26
13.11
18.05
18.58
1.94
5.94
8.97
16.08
22.81
29.56
1.56
4.55
11.16
16.62
21.58
31.2
2.19
4.93
9.93
14.58
19.19
26.41

7.9
356
1325
2359
2896
2629
1.3
907
1812
3135
3274
2330
17.6
598
907
678
704
920
4.4
11.0
785
611
549
580.3

61.7
15.5
3.8
2.0
2.3
2.0
54.2
4.6
3.6
2.0
1.6
1.6
56.7
3.5
3.5
2.58
2.6
1.0
41.7
9.1
3.6
3.5
3.1
2.36

Fig. 5a . AEKI DATA, Zr-4
1200

1000001100

1000

Fig. 5b. AEKI DATA, E-110

900 oC

10000

1100

1000

900 oC

10000
Brittle1000

DuctileExpon. (17% B-J)

100
10
1

Oxidation Time, s

Oxidation Time, s

1200

100000

1000
100
10
1

6

7

8

9

6

1/T( oK) x10000

7

8

9

1/T(oK) x10000

Fig.5. Temperature-time plot of the Hungarian data. The Zr-4 data fit with the BJ 17% criterion (solidline), whereas the E-110 limit is at BJ 8% (dashed line), i.e. at~100ºC lower temperature.

31

NEA/CSNI/R(2006)5
Table 4. CZECH DATABASE (UJP) Ref [16,17]
Zr-4 0.71 mm
E-110 0.71mm
Oxidation
Temp.
ºC
1200
1200
1200
1200
1200
1200
1200
1100
1100
1100
1100
1100
1100
1100
1100
1000
1000
1000
1000
1000
1000
1000
1000
800
800
800
800
800
800

Displacement, ∆L/D, %

Time,
sec.

Asreceived
20.17
20.17
20.17
9.93
7.76
5.12
3.41

0
90
180
270
360
600
720
0
90
180
360
720
1800
3600
180
0
360
900
1800
1800
3600
7200
10800
360
1800
3600
10800
18000
25200

2 µm
oxide
6.05
5.74
3.76
3.29
3.38

10 µm
oxide
4.35
4.66
2.55
1.98
2.30

3.58
19.94
19.94
19.94
5.7
0.55
0.68
0.04
6.46
20.17
20.17
16.14
10.86
10.24
6.05
1.86
2.02
20.17
20.17
20.17
20.17
19.24
20.17

1.87
19.94
1.5
1.3
0.4
0.3
0
0.4
3.1
20.17
20.17
11.48
6.67
8.38
4.81
2.79
3.10
20.17
20.17
20.17
20.17

19.94
19.94
19.94
4.58
0.66
20.17
17.07
19.24
18.31
12.73
2.02
0.00
20.17
20.17
20.17
20.17
20.17

Displacement, ∆L/D, %8%

Asreceived
3.4
2.2
2.5
1.8
2.0
1.6
2.0

2 µm
oxide
23.6
4.9
3.5
3.8
2.3

10 µm
oxide
4.9
3.1
2.2
2.4
2.1

1.4
23.6
23.6
23.6
0

1.1
23.6
0
0
0
0
0
0

23.6
23.6
23.6
22.3
23.6
4.0
1.6
1.8
23.6
23.6
23.6
23.6
23.6
23.6

23.6
23.6
13.4
3.4
3.8
2.4
1.4
2.4
23.6
23.6
23.6
23.6
10.0
5.8

23.6
23.6
0
0
0
23.6
2.9
4.5
3.8
3.3
1.8
2.5
23.6
23.6
23.6
23.6
23.6
23.6

20.17

Fig. 6a. UJP DATA, Zr-4 (0.71 mm)
Fig. 6b. UJP DATA, E-110 (0.71 mm)
800 o C

1200 1100 1000

100000Oxidation Time, s

Oxidation Time, s

10000
1000
100
10

800 oC

1200 1100 1000

100000
10000
1000
100
10
1

1
6

7

8

9

6

10

7

8

9

10

1/T(oK) x10000

1/T(oK) x10000

Fig.6. Temperature-time plot of the Czech data. As for the AEKI data, the Zr-4 data fit with the BJ
17% criterion (solid line), whereas the E-110 limit is at BJ 8% (dashed line), or at ~100°C lower
temperature.

32

NEA/CSNI/R(2006)5
Table 5. ANL DATA BASE Ref. [18-21] and [35](As-received samples, ring compression at RT. Ring compression temperature effectand pre-hydriding effect are summarised at the bottom of the table)
Zr-4 0.57 mm
Time, sOff. strain, %
210
46
840
29
1900
7.5
2430
5.1
3360
3.2

ZIRLO 0.57 mm
M5 0.61 mm
T, C
Time, s Off. strain, % Time, sOff. strain, %
1000
210
52
180
48
1000
840
25
780
51
1000
1900
20
1700
40
1000
2430
4.7
2100
19
1000
3360
2.9
2980
3.2
1000
3560
3560
2.5
1100
67
58
67
52
60
57
1100
266
20
266
22
220
39
1100
600
5.4
600
5.1
520
7.5
1100
770
5
770
3.5
690
4.0
1100
1065
4.8
1065
3.3
930
1.8
1100
1100
1100
1100
3.2
1200
25
37
25
56
20
56
1200
65
4.0
65
1200
100
1.0
100
2.8
85
3.4
1200
165
1.0
165
1.2
140
1.8
1200
240
0.8
240
1.9
190
1.7
1200
280
1.0
280
1.8
250
1.4
1200
400
0.5
400
1.2
350
2.0
• Significant ductility enhancement at 100-135C vs. RT, even for ECRBJ >17%
• Pre-hydrided unirradiated 17x17 + 15x15 Zr-4 samples, oxidized at 1204C; ringcompression at 135C
o Ductile-brittle transition at ECRBJ = 6.5% for H≥600ppm (15x15 Zr-4)
o Ductile-brittle transition at ECRBJ = 9.5% for H≥400ppm (15x15 Zr-4)
o Pre-hydrided 17x17 better than 15x15 , but transition at ECRBJ <13% for H≥300ppm
Ref [35]

Fig.7b. ANL, ZIRLO 0.57mm

1200 1100 1000 oC

100000Oxidation Time, s

Oxidation Time, s

Fig.7a. ANL, Zr-4 0.57mm

10000
1000
100
10
1
6

7

8

1/T(K) x10000

91200 1100 1000 o C

100000
10000
Brittle1000
DuctileBrittle100(17% B-J)
Expon.
10

Expon. (17% B-J)

DuctileFig.7c. ANL M5 0.61mm

1
6

7

8

9

1/T(K) x10000

Oxidation Time, so1200 1100 1000 oC

100000
10000

Brittle1000

Ductile100

Expon. (17% B-J)

10
1
6

7

8

9

1/T(K) x10000

Fig.7. Argonne RT ductility data for Zr-4, ZIRLO and M5 cladding. The data generally fit with the 17% BJ
limit (solid lines), but there is a tendency to lower ductility at 1200ºC. However, ANL ring compression
data at 135ºC exhibit ductility beyond the 17% line also at 1200ºC.

33

NEA/CSNI/R(2006)5

Fig.8. CEA data on ring compression tests of Zr-4 and M5 (RT) after steam oxidation at 1200ºC [23]. The data
show similar behaviour in terms of measured ECR (time not reported). The data also show a pronounced effectof pre-hydriding for both alloys.
60
670 ± 210 wppm H, Tmax = 1140°C
600 ± 90 wppm H, Tmax = 1185°C
780 ± 70 wppm H, Tmax = 1204°C

Offset Strain, %50

750 ± 100 wppm H, Tmax = 1206°C
545 ± 80 wppm H, Tmax = 1206°C

40

ANL RT data points for
non-irradiated Zr-4 (RT)30

ANL high burnup HBR data fit(at 135°C)

20

10
0
0

2

4

6

8

10

12

Predicted (CP) ECR, %Table 6 Ref [24, 31]
Equivalenttime, seconds

Ductility
at 350ºC

4% [3%CP]

1100

48

Ductile6.5%[5%CP]

1140

80

Ductile6.5%[5%CP]

1140

80

Ductile9% [7%CP]

1163

120 → 9% BJ

Ductile13%[10%CP]1193

180 → 13% BJ

Brittle10000
Oxidation equiv. time, s

SAMPLE Equivalent temBJ ECR,% perature, ºC

ANL DUCTILITY TESTS ON HIGH BURN-UP
CLADDING (H.B. Robinson fuel)

17%

1000

Brittle11%100

Ductile10

Expon.
(17% B-J)

1
6

7

8

9

1/T(K) x10000 (equivalent T)

Fig.9. ANL ductility data from H.B Robinson fuel at high burn-up [24, 33]. The ECRBJ was calculated here
from the temperature history (top left figure). Equivalent temperature and time give the same ECRBJ for aconstant temperature transient. The results show a ductile-brittle transition somewhat below the 17% line,
i.e. for a ECRBJ of ~11%. The upper right figure shows that the burn-up effect is comparable with theductility change from 135°C to RT.No oxide cracking is accounted for in these tests.

34

NEA/CSNI/R(2006)5

Fig.10. JAERI quench test database on as-received Zr-4 tubes subjected to high temperature oxidation andquench, under full constraint condition. Based on these data, the failure limit in Japan was set at ECRBJ=15%[25].

1573 K 1473 K 1373 K 1273 K

1573 K 1473 K 1373 K 1273 K

As-received(Failed)○ Non-hydrided(Survived)

Pre-hydrided(Failed)△ Pre-hydrided(Survived)● Non-hydrided(Failed)As-received(Survived)▲ Pre-hydrided(Failed)Pre-hydrided(Survived)

8

Threshold for non-hydrided
cladding tubes

Oxidationamount in
ECR

6

7.5

17%

4

4

14%

17%
7%

(a) Non-hydrided(b) Pre-hydrided

2
6

6.5

7

7.5

8

8.5 6

6.5

7

7.5

8

(1/ Oxidation temperature (K))x104

8.5

Ln (Oxidation Time (s))

Ln (Oxidation time (s))

10Unirradiated (100 – 350 ppm)
Survived
Fractured
48GWd/t (115 – 205 ppm)
Survived
Fractured
Unirradiated (400 – 800 ppm)
Survived
Fractured

6.5

5.5

4.5
6.5

7.0

7.5 ｘ 10-4

Reciprocal Temperature (1/K)

Fig.11. JAERI quench database on pre-hydrided cladding [26, 27].The figure above is for full constraint,
whereas that on the right hand side is for a 50 kg constraint. Under full constraint conditions, pre-hydriding
causes a failure limit drop from 17 to 7% ECRBJ. The data for a 50 kg constraint, instead, show a significantly
smaller pre-hydriding effect. (Note: the scale on the figure to the right is very sensitive and the 17% limit isbelow the two lines shown in the diagram).

35

NEA/CSNI/R(2006)5
Table 7. JAERI and ANL DATABASE: Quench tests with high burn-up cladding Ref [27, 30](JAERI tests are with 50 kg constraint, whereas there was no constraints in the ANL tests)
Fuel
Type

Burn-up, Base Oxide
MWd/kg thickness,
µm
44
20

H
content
ppm
170

Oxid.
Temper.
ºC
1176

Oxidation
Time, sResult486

Failure

JAERI

44

25

210

1178

120

JAERI

39

18

140

1154

200

41

18

140

1172

363

41

15

120

1030

2195

39

15

120

1177

543

No
failureNo
failureNo
failureNo
failureFailure

BWR

56

10

70

1204

300

ANL

BWR

56

10

70

1204

300

No
failureNo
failure

PWR
A3-1
PWR
A1-2
PWR
BL-3
PWR
BI-3
PWR
BI-5
PWR
BL-7

JAERI
JAERI
JAERI
JAERI

ANL

JAERI AND ANL QUENCH TESTSOxidation Time, s

10000010000
JAERI, Failed
1000

JAERI, Non-Failed

100

ANL, Handling Fail.
Expon. (17% B-J)

10
1
6

7

8

9

1/T(K), x10000

Fig.12. Plot of the JAERI and ANL data on quench tests of high burn-up cladding [27, 30]. The data fit well withthe 17% ECRBJ limit. One should notice that there are two coincidental data points in the ANL case. The ANL
tests were in fact not meant to provide a failure limit, and were run without constraints. In one case the sample
failed during post-test handling. These results are shown here for completeness.

36

NEA/CSNI/R(2006)5

C ase of large oxidation

Reduction of LOCAlimit with burn-up

25
Limit ECR(B-J), %Oxide thickness, micr.

160

20

140
120
100
80
60
40
20
0

15

0

"Worst case"20

40

Total ECR

10

60

80

100

Burn-up, M W d/kg

5
0
0

20

40

60

80

Reduction of LOCA limit with burn-up

100

1200
1100
T300, Deg.C

Burn-up, MWd/kg

1000
900

"Worst case"

800

Total ECR

700
600
500
400
0

20

40

60

80

100

Burn-up, MWd/kg

Fig.13. Calculated reduction of LOCA limit vs. burn-up in the “worst case” (Eq. /5/ and /6/). The limit isexpressed in terms of ECR on the figure above, and as temperature limit (for 300s transient) on the right handside figure. The reference corrosion vs. burn-up is given in the top-right figure. The “total ECR” criterion is
also shown for comparison.

Large andsmall oxidationcases

Reduction of LOCA limit with burn-up

160

T300, Deg.C

1100
1000
900

Large oxidation

800

Small oxidation

700
600
500

Oxide thickness, micr.

1200

140
120
100
80
60
40
20
0

400
0

20

40

60

80

0

100

20

40

60

80

100

Burn-up, MWd/kg

Burn-up, MWd/kg

Fig.14. Comparison of the “worst case” burn-up effect (Eq/6/) for cladding with a large and small baseirradiation corrosion. One can notice that the temperature limit decreases by ~160 ºC for large corrosion,
mostly in the 30-50 MWd/kg range. For low corrosion, the decrease is only ~60ºC at 50 MWd/kg.

37

NEA/CSNI/R(2006)538

NEA/CSNI/R(2006)5

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NEA/CSNI/R(2006)5[31] Yu. K. Bibilashvili et al: Thermomechanical Properties of Oxidised Zirconium-Based Alloy
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and Results”, Nuclear Safety, Vol.37, No.4, 1996.
[37] L. Yegorova, V. Asmolov, G. Abyshov, V. Malofeev, A. Avvakumov, E. Kaplar, K. Lioutov,
A. Shestopalov, A. Bortash, L. Maiorov, K. Mikitiouk, V. Palvanov, V. Smirnov, A. Gotyachev,
V. Prokhorov, and A. Vurim: “Data Base on the Behaviour of High Burnup Fuel Rods with Zr1%Nb
Cladding and UO2 Fuel (VVER Type) under Reactivity Accident Conditions”, NUREG/IA-0156(IPSN99/08-02, NSI/RRC KI 2179) Vol.1, 2, 1999.
[38] L. Yegorova, O. Nechaeva, N. Jouravkova, A. Salatov, V. Smirnov, A. Goryachev, V. Ustinenko,
I. Smirnov: “Experimental Study of Narrow Pulse Effects on the Behaviour of High Burnup Fuel Rods
with Zr-1%Nb Cladding and UO2 Fuel (VVER) under Reactivity Accident Conditions”, NUREG report(IRSN/DRAM 2005-275, NSI RRC KI 3230), Vol. 1, 2, 2006 (to be published).

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