spine:
5.455 mm
81 pages

OECD–IAEA Paks Fuel Project
Final Report

INTERNATIONAL ATOMIC ENERGY AGENCY
VIENNA

09-2210-PUB-1389-cover.indd 1

2010-06-29 10:12:33

OECD–IAEA PAKS FUEL PROJECTThe following States are Members of the International Atomic Energy Agency:
AFGHANISTANALBANIA
ALGERIAANGOLA
ARGENTINA
ARMENIA
AUSTRALIA
AUSTRIAAZERBAIJAN
BAHRAINBANGLADESH
BELARUSBELGIUM
BELIZEBENIN
BOLIVIABOSNIA AND HERZEGOVINABOTSWANABRAZIL
BULGARIA
BURKINA FASO
BURUNDICAMEROONCANADACENTRAL AFRICAN
REPUBLIC
CHAD
CHILE
CHINA
COLOMBIACONGOCOSTA RICA
CÔTE D’IVOIRECROATIACUBACYPRUSCZECH REPUBLIC
DEMOCRATIC REPUBLIC
OF THE CONGO
DENMARKDOMINICAN REPUBLICECUADOREGYPTEL SALVADORERITREAESTONIAETHIOPIAFINLANDFRANCEGABONGEORGIAGERMANYGHANA
GREECEGUATEMALAHAITI
HOLY SEE
HONDURASHUNGARYICELAND
INDIA
INDONESIA
IRAN, ISLAMIC REPUBLIC OFIRAQIRELANDISRAELITALY
JAMAICAJAPANJORDANKAZAKHSTANKENYAKOREA, REPUBLIC OFKUWAITKYRGYZSTANLATVIALEBANONLESOTHO
LIBERIA
LIBYAN ARAB JAMAHIRIYALIECHTENSTEIN
LITHUANIALUXEMBOURG
MADAGASCARMALAWI
MALAYSIAMALI
MALTA
MARSHALL ISLANDS
MAURITANIA
MAURITIUS
MEXICOMONACOMONGOLIAMONTENEGRO
MOROCCOMOZAMBIQUE
MYANMAR
NAMIBIANEPAL
NETHERLANDSNEW ZEALANDNICARAGUANIGERNIGERIANORWAYOMANPAKISTAN
PALAU
PANAMAPARAGUAYPERU
PHILIPPINESPOLANDPORTUGALQATARREPUBLIC OF MOLDOVAROMANIARUSSIAN FEDERATIONSAUDI ARABIASENEGALSERBIASEYCHELLESSIERRA LEONESINGAPORESLOVAKIASLOVENIA
SOUTH AFRICASPAIN
SRI LANKA
SUDAN
SWEDENSWITZERLANDSYRIAN ARAB REPUBLICTAJIKISTANTHAILANDTHE FORMER YUGOSLAVREPUBLIC OF MACEDONIATUNISIA
TURKEYUGANDAUKRAINE
UNITED ARAB EMIRATESUNITED KINGDOM OFGREAT BRITAIN ANDNORTHERN IRELAND
UNITED REPUBLICOF TANZANIAUNITED STATES OF AMERICAURUGUAYUZBEKISTANVENEZUELAVIETNAMYEMENZAMBIA
ZIMBABWE

The Agency’s Statute was approved on 23 October 1956 by the Conference on the Statute of the IAEA
held at United Nations Headquarters, New York; it entered into force on 29 July 1957. The Headquarters of the
Agency are situated in Vienna. Its principal objective is “to accelerate and enlarge the contribution of atomic
energy to peace, health and prosperity throughout the world’’.

OECD–IAEA
PAKS FUEL PROJECTFINAL REPORTINTERNATIONAL ATOMIC ENERGY AGENCY
VIENNA, 2009

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OECD–IAEA PAKS FUEL PROJECT
IAEA, VIENNA, 2010
© IAEA, 2010
Printed by the IAEA in Austria
May 2010

FOREWORD
It is important for nuclear power plant designers, operators and regulators to effectively use
lessons learned from events occurring at nuclear power plants since, in general, it is
impossible to reproduce the event using experimental facilities. In particular, evaluation of theevent using accident analysis codes is expected to contribute to improving understanding ofphenomena during the events and to facilitate the validation of computer codes throughsimulation analyses. The information presented in this publication will be of use in futurerevisions of safety guides on accident analysis.
During a fuel crud removal operation on the Paks-2 unit of the Paks nuclear power plant,
Hungary on 10 April 2003, several fuel assemblies were severely damaged. The assemblies
were being cleaned in a special tank under deep water in a service pit connected to the spent
fuel storage pool. The first sign of fuel failures was the detection of some fission gases
released from the cleaning tank. Later, visual inspection revealed that most of the 30 fuelassemblies suffered heavy oxidation and fragmentation. The first evaluation of the event
showed that the severe fuel damage had been caused by inadequate cooling.
The Paks-2 event was discussed in various committees of the OECD Nuclear Energy Agency(OECD/NEA) and of the International Atomic Energy Agency (IAEA). Recommendations
were made to undertake actions to improve the understanding of the incident sequence and of
the consequence this had on the fuel. It was considered that the Paks-2 event may constitute a
useful case for a comparative exercise on safety codes, in particular for models devised to
predict fuel damage and potential releases under abnormal cooling conditions and theanalyses of the Paks-2 event may provide information which is relevant for in-reactor andspent fuel storage safety evaluations.
The OECD-IAEA Paks Fuel Project was established in 2005 as a joint project between the
IAEA and the OECD/NEA. The IAEA provided financial support to the operating agent(Hungarian Academy of Sciences KFKI Atomic Energy Research Institute (AEKI)) andreviewed the progress of the project within the framework of Technical Cooperation Project,
RER9076, on Strengthening Safety and Reliability of Fuel and Material in Nuclear Power
Plants. Thirty organizations from sixteen countries participated in the project. The following
participants performed analyses of the event using numerical models based on a common
database:
•
•
•
•
•
•

AEKI (Hungary),
BME NTI (Hungary),
GRS (Germany),
IRSN (France),
IVS (Slovakia),
KI (Russian Federation),

•
•
•
•
•

SUEZ – TRACTEBEL ENGINEERING(Belgium),
NRC (United States of America),
VEIKI (Hungary),
VTT (Finland),VUJE (Slovakia).

The main results of calculations are summarized in the present report.
The IAEA and the OECD/NEA acknowledge the work of the participating experts and wish
to thank them for their valuable contributions to this publication. In particular, the work of
Z. Hozer in compiling the report is gratefully acknowledged, as well as N. Tricot forcompleting the final stages of preparation. The corresponding report has been issued as an
NEA/CSNI publication as NEA/CSNI/R(2008)2. The IAEA and the OECD/NEA officers
responsible for the preparation of this publication were Y. Makihara and C. Vitanza,
respectively.

1

EDITORIAL NOTE
The use of particular designations of countries or territories does not imply any judgement by the
publisher, the IAEA, as to the legal status of such countries or territories, of their authorities andinstitutions or of the delimitation of their boundaries.
The mention of names of specific companies or products (whether or not indicated as registered) does
not imply any intention to infringe proprietary rights, nor should it be construed as an endorsement
or recommendation on the part of the IAEA.

CONTENTS
EXECUTIVE SUMMARY........................................................................................................ 1
1

INTRODUCTION ........................................................................................................... 1\
1.1.
1.2.

2

DESCRIPTION OF THE INCIDENT.............................................................................. 5
2.1
2.2
2.3

2.4
2.5
2.6
3

Background......................................................................................................... 1
Structure.............................................................................................................. 4

Preliminaries....................................................................................................... 5
Operation of the cleaning tank............................................................................ 7
Chronology of the incident ............................................................................... 10
2.3.1 Formation of steam volume .................................................................. 11
2.3.2 Plastic deformation of cladding and high temperature oxidation ......... 122.3.3 Quench .................................................................................................. 13
State of damaged fuel ....................................................................................... 14
Post-incident investigations.............................................................................. 15
The recovery work............................................................................................ 16

MAJOR RESULTS OF ANALYSES............................................................................. 18
3.1

3.2

3.3

3.4

3.5

3.6

Participant SUEZ - TRACTEBEL ENGINEERING ....................................... 19
3.1.1 Code description ................................................................................... 19
3.1.2 Model description ................................................................................. 20
3.1.3 Chronology of main events................................................................... 22
3.1.4 Conclusions........................................................................................... 22
Participant VTT ................................................................................................ 23
3.2.1 Introduction and case description ......................................................... 23
3.2.2 Calculation results................................................................................. 26
3.2.3 Conclusions........................................................................................... 27
Participant GRS ................................................................................................ 28
3.3.1 Code description ................................................................................... 28
3.3.2 Files delivered for the comparison........................................................ 28
3.3.3 Nodalisation and input model ............................................................... 28
3.3.4 Results of calculation............................................................................ 29
Participant AEKI .............................................................................................. 30
3.4.1 Thermal hydraulic analysis................................................................... 30
3.4.2 Thermal hydraulic analysis with the RELAP code............................... 30
3.4.3 Thermal hydraulic analysis with the ATHLET code............................ 33
3.4.4 Fuel behaviour calculation.................................................................... 34
3.4.5 Calculation of activity release............................................................... 35
3.4.6 Summary of calculated results.............................................................. 36
Participant BME NTI........................................................................................ 36
3.5.1 The 3D CFD models of the cleaning tank ............................................ 36
3.5.2 Boundary conditions, parameters of the model .................................... 37
3.5.3 Calculational results.............................................................................. 37
Participant VEIKI ............................................................................................. 39
3.6.1 MELCOR PWR and BWR models of the cleaning tank ...................... 39
3.6.2 Safe envelope of operation ................................................................... 40
3.6.3 Transient calculations ........................................................................... 41
3.6.4 Conclusions........................................................................................... 42

3.7

3.8

3.9

3.10

4

Participant KI/IRSN ......................................................................................... 43
3.7.1 Code description ................................................................................... 43
3.7.2 Thermal hydraulic phase of the transient.............................................. 43
3.7.3 SFD phase of the transient .................................................................... 44
Participant IVS ................................................................................................. 45
3.8.1 Code description ................................................................................... 45
3.8.2 Nodalisation scheme ............................................................................. 46
3.8.3 Main results........................................................................................... 47
Participant VUJE .............................................................................................. 48
3.9.1 Description of the input model ............................................................. 48
3.9.2 Initial and boundary conditions ............................................................ 49
3.9.3 Description of results ............................................................................ 50
Participant USNRC........................................................................................... 50
3.10.1 Code description ................................................................................... 51
3.10.2 Thermal hydraulic response.................................................................. 51
3.10.3 Fission Product release ......................................................................... 52

COMPARISON OF CALCULATED RESULTS .......................................................... 54
4.1
4.2
4.3
4.4

Thermal hydraulics ........................................................................................... 56
Fuel behaviour .................................................................................................. 57
Activity release ................................................................................................. 60
Plots of selected parameters ............................................................................. 61

5

OUTLOOK ..................................................................................................................... 73

6

CONCLUSIONS ............................................................................................................ 74

REFERENCES......................................................................................................................... 77
ABBREVIATIONS.................................................................................................................. 79
CONTRIBUTORS TO DRAFTING AND REVIEW ............................................................. 81

EXECUTIVE SUMMARYThe OECD-IAEA Paks Fuel Project aimed to support the understanding of fuel behaviour in
accident conditions on the basis of analyses of the Paks-2 event. Numerical simulation of themost relevant aspects of the event and comparison of the calculation results with the available
information was carried out between 2006 and 2007. A database was collected to provide
input data for the code calculations. The activities of the project covered the following threeareas:•

Thermal hydraulic calculations described the cooling conditions possibly establishedduring the incident.

•

Simulation of fuel behaviour described the oxidation and degradation mechanisms of fuelassemblies.

•

The release of fission products from the failed fuel rods was estimated and compared toavailable measured data.

The produced numerical results improved the understanding of the causes and mechanisms offuel failures during the Paks-2 incident and provided new information on the behaviour of
nuclear fuel under accident conditions.
1

INTRODUCTION

1.1

BACKGROUND

On the 10th April 2003 severe damage to fuel assemblies took place during an incident at
Unit 2 of Paks nuclear power plant in Hungary. The assemblies were being cleaned in a
special tank below the water level of the spent fuel storage pool in order to remove crudbuildup. That afternoon, the chemical cleaning of assemblies was completed and the fuel rods
were being cooled by circulation of spent fuel storage pool water. The first sign of fuel failure
was the detection of some fission gases released from the cleaning tank during that evening.
The cleaning tank cover locks were released after midnight and this operation was followed
by a sudden increase in activity concentrations. The visual inspection revealed that all 30 fuelassemblies were severely damaged. The first evaluation of the event showed that the severe
fuel damage happened due to inadequate coolant circulation within the cleaning tank. The
damaged fuel assemblies were removed from the cleaning tank in 2006 and are stored in
special canisters in the spent fuel storage pool of the Paks nuclear power plant.
The event was reported to the IAEA and classified as a Level 3 (severe) incident on theInternational Nuclear Event Scale (INES). The IAEA sent an expert review mission toinvestigate the incident in June 2003.
First information on the Paks-2 incidents was provided to the NEA Committee on the Safetyof Nuclear Installations (CSNI) in its June 2003 meeting, two months after the incident. At
that time the CSNI requested the NEA secretariat to co-ordinate actions with Hungarian andRussian parties aimed at assessing the extent to which possible fuel investigations would
benefit the nuclear community while helping to clarify the conditions experienced by the fuelduring the incident. As a first step in this direction, a meeting was organised in Budapest on
18 August 2003 under the auspices of the NEA. Visual examinations of the damaged
assemblies as well as analyses were presented on that occasion. The participants agreed thatinternational collaboration should be pursued to the extent feasible. Recommendations were1

made that the NEA and the IAEA try to organise a joint project in co-ordination with Russianand Hungarian organisations, provided that useful information could be gathered through suchinitiative.
The experts of the Hungarian Academy of Sciences KFKI Atomic Energy Research Institute(AEKI) and of the Russian Research Centre Kurchatov Institute (KI) prepared the first draftof the potential international project in 2004. An experts’ meeting of potential participants
was organised in November 2004 in Studsvik that represented a crucial point in the process oflaunching the project. The participants were requested to assess the technical validity of the
proposed project, suggest improvements if needed and provide a first indication of technicalinterest in the proposal. It was intended to organise the project in two phases:-

Phase 1 focussing on analytical activities.

-

Phase 2 focussing on the fuel examinations.

The analytical part of the project was intended to cover a database collection which would
have served as basis for code calculations, for the selection of the fuel to be examined and for
preparing the list of fuel examinations to be carried out. The fuel examination phase would
have involved poolside activities first, then the shipping from Paks to hot cells in Russia andthen the hot cells examinations.
Phase 1 of the project was launched in the framework of the OECD-IAEA Paks Fuel Projectin 2006 aimed to investigate fuel behaviour in accident conditions on the basis of analyses of
the Paks-2 event. The project was co-ordinated by the OECD NEA and by the IAEA. The
meetings were chaired by GRS, Germany. The tasks of the operating agent were fulfilled by
the Hungarian AEKI.
There was no formal agreement signed by all participants for the OECD-IAEA Paks Fuel
Project. The participation in the project was available through national nuclear regulatoryorganisations with the approval of OECD NEA and IAEA. During the two years duration of
the project 58 experts from 34 organisations of 16 countries were listed among theparticipants. All participants had access to the materials of the project, including the database
and the presentations given at the meetings. In the frame of the RER/9/076 on Strengthening
Safety and Reliability of Nuclear Fuel and Materials in Nuclear Power Plants project the
IAEA had made three contracts with AEKI in order to support financially the activities of the
Operating Agent.
The kick-off meeting of the project was held in January 2006 in Budapest. The presentations
on the earlier analytical work performed by some participants revealed that there were big
differences in the calculation results, since most of calculations were done just after the eventin 2003 and little information was provided to the analysts. Nevertheless, it was found in thethermal hydraulic calculations, that the by-pass flow would play an essential role for theprediction of uncovery time. The fuel properties analysis also showed that the results werestrongly dependent on the assumptions adopted by the analysts. In particular, decay heat andtreatment of zirconium-water reactions would give a strong impact on the results. It was also
revealed that the steam starvation effect may have stabilized the heat-up phase, which isstrongly dependent on the assumed gas leakage rate from the tank. It was agreed that new
calculations would be carried out by the participants using a common database. The scope of
calculations would cover thermal hydraulic, fuel behaviour and activity release phenomena:2

•Thermal hydraulic calculations to describe how the inadequate cooling conditions could
have established during the incident.

•

Fuel behaviour simulation to describe the oxidation and degradation mechanisms of fuelassemblies.

•

Activity release and transport calculations to simulate the release of fission products fromthe failed fuel rods.

After the first meeting the AEKI collected the database, established a website for the projectand prepared a proposal for the requirements for output data.
The database (Appendix A/I) included the following main parts:•

Design characteristics of WWER-440 fuel assemblies (main geometrical data, somemechanical properties, oxidation kinetics of E110 cladding, and integral data ofassemblies). The source of these data was mainly the open literature publications of the
WWER fuel supplier. Special experimental data on the high temperature behaviour of
E110 alloy, including oxidation in hydrogen rich steam were provided by the AEKI.

•

Operational data of damaged fuel assemblies (power histories of fuel assemblies, burnup,
fuel rod internal pressure, isotope inventories, decay heat and axial power distribution).
These data included the main parameters of the damaged fuel assemblies during theirnormal operation in the reactors. The parameters describing the fuel state before theincident were produced by the experts of Paks nuclear power plant and AEKI using the
KARATE, SCALE, ORIGEN, TIBSO and TRANSURANUS codes. During the projectsome control calculations on fuel behaviour were calculated by VTT, Finland using the
FRAPCON code (Appendix B/III).

•

Design characteristics of the cleaning tank were provided by the tank designer AREVA
NP GmbH.

•

Measured data during the incident: (temperature, water level measurements, cleaning tankoutlet flow rates).

•

Activity measurements (measured coolant activity concentrations, activity release throughthe chimney, flow rate of water make-up system, released activities). These data included
not only the incident, but the period of wet storage of damaged fuel until their removalfrom the service pit.

•

Reports (describing results of investigations, chronology).

•

Status of fuel (results of visuals observations).

The database and the materials of the meetings were available on a password protectedwebsite (Appendix A/II) for the participants who received individual user id and password toenter the site. The content of the website was several times updated during the project.
The list of requested output parameters (Appendix A/III) was specified in order to makepossible the comparison of the results of calculations carried out by different organisations.
The approach applied in this project was similar as in International Standard Problem (ISP)3projects. The total number of requested parameters was more than one-hundred. The maingroups of parameters were the followings:•

histories of thermal hydraulic parameters,•

profiles of thermal hydraulic parameters by the time the maximum cladding temperature
exceeds 800°C,•

histories of fuel parameters,•

profiles of fuel parameters before quenching and,•

histories of activities release from the fuel rods.

The requirements specified also the scope of written presentations of the results according towhich the participants prepared individual chapters on their work. These materials werecollected into Chapter 3 of the present report. More detailed description of each calculation isavailable in Appendix B.
The second meeting of the project was held in October 2006 in Budapest. The agenda
included presentations with preliminary results of thermal hydraulic analyses, fuel behaviour
analyses and activity release calculations. Most of the presented results had preliminary
character and it was agreed that the participants would continue the numerical calculations.
The participants of the second meeting agreed that the thermal hydraulic calculations should
be completed by the end of March 2007, and the fuel behaviour and activity calculations by
the end of April 2007.
The final meeting of the project was held in April 2007 in Budapest. Individual presentations
were given on the latest calculated results by the participants. The first comparison of selectedparameters was also discussed in details. The participants agreed on the table of contents of
the report and that the report will be published as a joint IAEA-OECD Project Report.
The preparation of fuel examination in Phase 2 of the project was discussed during eachmeeting. Experts of the Paks nuclear power plant regularly informed the project on the stateof fuel and the actions related to the removal of damaged fuel from the cleaning tank and its
further storage. Hungarian experts prepared a proposal on the selection of fuel assembly forhot cell examinations. The damaged fuel is stored in the Paks nuclear power plant today. The
hot cell examinations could be carried out in Dimitrovgrad, Russia. Until the time of thewriting of the present report, no formal proposal had been developed for the examination ofdamaged fuel in Phase 2.
1.2

STRUCTURE

This report consists of six sections. Section 2 provides a description of the incident. Section 3
sets out the major results of the analysis and Section 4 provides a comparison of calculated
results. An outlook for the future of the Paks Fuel Project is contained in Section 5. Section 6
contains the conclusions drawn from phase 1 of the project. The attached CD contains
annexes providing description of the database, website and output data for the project anddetailed descriptions of the various calculations.

4

2

DESCRIPTION OF THE INCIDENT

The Paks nuclear power plant has four WWER-440 type reactor units of 440 MW original
capacities per each. These units serve nearly 40% of the Hungarian electricity demand. The
four units were installed between 1983 and 1987.
On 10th of April, 2003 a serious incident (classified as Level 3 on the INES scale) caused
heavy damage of 30 fuel assemblies in the Unit 2 of the Paks nuclear power plant. During theprevious years, Units 1, 2 and 3 were burdened with considerable magnetite corrosion productdeposition. Parallel to the annual maintenance, the deposited fuel assemblies were cleaned ina special cleaning tank which was constructed by a large European company and was installed
into the so-called refuelling pit just beside the reactor.
The cleaning processes were performed on batches of 30 fuel assemblies. On 10th of April,
2003, the cleaning of the sixth batch was finished. During the cleaning process itself, thecooling of the fuel was ensured by high mass flow rate coolant circulation. However, after
finishing the 6th batch’s cleaning process the fuel was not removed from the cleaning tank
immediately since the crane was busy with other tasks. In the meantime, the coolant wascirculated by a submergible pump with much lower mass flow rate according to the service
regulations.
The later inspection has revealed that due to the special design of the cleaning tank and thecharacteristic of the fuel assemblies, the cooling of the submersible pump of lower mass flow
was insufficient. The coolant was not capable of removing the bulk of the decay heat and thisresulted in a complicated natural convection flow in the cleaning tank and the fuel assemblies.
The temperature stratification blocked the flow and therefore the coolant temperature reachedsaturation temperature in the upper part of the cleaning tank. Then the steam-formation
pushed the main volume of the coolant out of the cleaning tank vessel due to thedisadvantageous design of the instrument. This way, the fuel assemblies were left without
cooling for hours and heated up to above 1000°C, which resulted in severe damage andoxidation of the fuel assemblies’ structure.
The radioactive noble gases that escaped from the damaged fuel assemblies were released into
the environment through the reactor hall stack. However, the impact on the environment was
negligible thanks to the small released quantity.
In this chapter the main technical causes, the course of the incident, and the steps of therecovery work are set out.
2.1

Preliminaries

The first steps towards the incident could be the chemical decontamination processes of thesteam generators which were performed in large number in the 90’s. At that time corrosionerosion caused problems in the feed water distributors of the steam generators which resulted
in the series of distributor replacements (see Figure 2.1). During these operations, themaintenance staff had to work inside the steam generator secondary side. For radiologicalprotection of the staff, chemical decontamination of the SGs’ primary side were adopted.
(This method was used only on Units 1, 2 and 3. In case of Unit 4, the so called physicalshielding technique subsequently used based on methods developed abroad.)5

Figure 2.1. The old (left) and the new (right) distributors in the steam generator secondaryside (courtesy of GMO CAD group of the Paks nuclear power plant)
The widely used chemical decontaminations resulted in unacceptable aftermath: thedissolution of corrosion products accelerated to a large measure and the coolant transported
the dissolved corrosion products into the reactor core. This yielded magnetite deposition on
the surfaces of the fuel assemblies. In certain positions, these magnetite depositions blocked
the part of the flow area, therefore the mass flow through these assemblies decreased and the
maximum outlet temperature increased.
Significant deposits were experienced first time in the Unit 2 of Paks nuclear power plant in
1996. Until 1998 the quantity of the depositions was significant enough to require thereplacement of the entire reactor core fuel with fresh assemblies for Unit 2. By 2000, the troubles
with depositions affected Units 1, 2 and 3. In February of 2003, during the cycle, an entire reactor
core fuel replacement had to be done in Unit 3. (In contrast, Unit 4, which used the physicalshielding method, did not have any deposition problems.)
To comply with the Technical Operating Regulations and other specifications, the Paks
nuclear power plant adopted the following measures: the power of the concerned units waslimited in accordance with the limitations for the fuel assembly outlet temperatures, thedropping times of the absorber elements were verified, and the increased difficulties
concerning the visual investigations of the reactors’ inner components were taken into
account.
The validity of safety analyses performed earlier was reviewed. The Russian firm Hydropress
performed the analysis of the operation with magnetite deposits burdened fuel assemblies [1]for both normal and incidental conditions of the reactor units. These investigations haveshown that the deformations of the fuel assembly shrouds might be attained much easier if the
shroud temperature exceeds 355°C.
The Hydropress report suggested new operational limits: the relative mass flow of theassemblies could not fall under 86% of the average value, because at lower mass flow rates,
the shroud temperatures might exceed the problematic temperature value of 355°C. For
controlling the mass flow limitations, the modification of the VERONA reactor core
monitoring system was also necessary. In addition, the Hydropress suggested to plan therefuelling with extreme caution.
As a consequence of the above drafted events, the plant management decided to clean theaffected fuel assemblies, first time on the turn of the millennium. In the years 2000 and 2001,
the chemical cleaning of a total sum of 170 fuel assemblies was performed successfully in an
instrument designed by Siemens KWU, which was capable for housing 7 fuel assemblies
simultaneously [2]. In accordance with the preliminary expectations, this instrument operated
effectively. (From the viewpoint of the events discussed below, it is important to mention that6the cleaned fuel assemblies were removed from the core two years before the cleaning process
started and they had low residual heat. Furthermore, this equipment would have been able toensure the cooling of fuel assemblies of much higher thermal power. This cleaning tank had
inlet junction in the bottom and outlet junction on the top of the vessel. The safety criteria
specified by the Paks nuclear power plant experts were fulfilled for this equipment as it wasproved by the main constructor Hydropress.)
In November of 2002, the Paks nuclear power plant commissioned a reputed WesternEuropean nuclear company to design, construct and operate a new chemical cleaning
equipment of larger dimensions. With this new cleaning tank the operators were able toperform the chemical cleaning of 30 fuel assemblies simultaneously with the same
permanganate-oxalic-acid method. The instrument arrived to Hungary in the beginning of2003 and was placed into the refuelling pit beside the reactor of Unit 2 in the Phase 1 of thePaks nuclear power plant (see Figure 2.2). The cleaning of the first batch of 30 fuelassemblies was started on 20th of March. Prior to the incident, five batches were cleaned
successfully. The 1st, 2nd, and 4th batches consisted of fuel assemblies removed a few years
earlier, but the 3rd and 5th batches contained fuel assemblies with significant residual power(removed from the reactor on those days).
2.2

Operation of the cleaning tank

The AMDA (acronym of the German phrase: Automatic Mobile Decontaminating Instrument)
cleaning tank was installed into the refuelling pit of Unit 2 (see Figure 2.2). The refuelling pit
was filled up with borated water and so it was in a common water space with the spent fuelpool. The fuel assemblies were loaded into and removed from the cleaning tank by therefuelling machine of the Paks nuclear power plant. During the normal operational mode
when the chemical cleaning was in progress (operational mode ‘C’), the oxalic acid solution –which played both the role of cleaning and cooling medium – circulated in a closed loop witha mass flow rate of 200-250 t/h (see Fig. 2.4). The heat exchanger built in the AMDA system
ensured the removal of the fuel assemblies’ decay heat.

Figure 2.2. The cleaning tank’s position in Unit 2

Figure 2.3. Flow scheme of thecleaning tank7

Cleaning technology(operational mode „C’)

Cleaning technology(operational mode „C’)

Water
level

Water
level

Submersible
pump(operationalmode „B’)

Refuellingpit

200-250 t/h

Refuellingpit

Submersible
pump(operationalmode „B’)

20 t/hCleaning
tank

Cleaningtank

Figure 2.4. Coolant flow during the ‘C’ Figure 2.5. Coolant flow during theoperational mode (Valves V1, V2 are ‘B’ operational mode (Valves V1, V2
closed, V3, V4 are opened) [3]are opened, V3, V4 are closed) [3]Figure 2.3 shows the flow pattern of the oxalic acid solution coolant inside the cleaning tankduring the cleaning process: the coolant flowed down through a pipe in the gap between the
inner and outer vessels of the double mantled cleaning tank. Then it entered the inner domain
at the bottom. The coolant flowed up inside the fuel assemblies and down among them. The
coolant left the inner volume via the outlet nozzles in the bottom part of the cleaning tank, at asimilar level, as the inlet nozzles. (This design solution was applied in order to avoid theaccumulation of the crud on the lower plate of the tank, since such problems were observed
during the operation of the small cleaning tank with seven fuel assemblies.) After that, themedium of mass flow rate of 200-250 t/h entered the cleaning module of the AMDA system(see Figure 2.4), where the dissolved magnetite and heat were removed from the solute.
After finishing the cleaning process, the system was changed over to the so-called ‘B’operational mode, which was rather different from the above described ‘C’ mode (seeFigure 2.5): a submersible pump of 20 t/h mass flow capacity circulated the coolant of therefuelling pit through the cleaning tank in an open loop. During the cleaning tank opening
process – using the reactor hall crane – and during the loading and removing of the fuelassemblies, the AMDA system was cooled in the ‘B’ operational mode.
The subsequent visual surveys after the incident showed clearly that oxidation in hightemperature steam took place in the cleaning tank for several hours which resulted in theembrittlement of the fuel assembly shrouds and the fuel pin claddings. At the opening of thecleaning tank, the injection of cold coolant caused the breaking up of the embrittled shrouds
and fuel pin claddings. The phenomenon of the high temperature oxidation indicates that the8cooling of the fuel assemblies was insufficient. The investigations revealed two possible
reasons for it:
1. Development of significant by-pass flow through the bottom perforations of the fuelassembly shrouds.
In the WWER-440 type reactors, two kinds of fuel assemblies are used: the normal‘working’ assemblies and the so-called follower assemblies. The follower assemblies
are positioned under the 37 absorber elements and attached to them. During the reactor
operation, these follower assemblies move together with the absorber elements and are
positioned partly or entirely in the reactor core. Otherwise, they can be found below thereactor core. The other 312 fuel assemblies are working assemblies, which have fix
positions in the reactor core.
The two fuel assembly types differ from each other in their head and leg parts. From theviewpoint of the incident, the most important difference is that the working assemblies’
shrouds are perforated by 12 and 12 holes of 9 mm diameter in the top and bottom parts.
The 6th batch – which suffered the incident – consisted of 11 working and 19 follower
assemblies. The subsequent investigations pointed out that with the rather low mass
flow rate ensured by the submersible pump in the ‘B’ operational mode, the bulk of thecoolant might flow to the outlet through the bottom perforations of the working
assemblies without flowing along and cooling the fuel assemblies. The special flow
pattern which might cause this phenomenon is discussed in section 3.2.
2. Displacement of one or more fuel assemblies.
The bottom positioning and support plate was formed such a manner that both theworking and follower assemblies were able to engage into the 30 seats on it (seeFigure 2.7). According to the different formation of the two assembly types, the seats
contain three flanges with different diameters. Unfortunately, the seats were notfurnished with such instruments which would have been able to signal the proper
engagement of the fuel assemblies into the seats. Furthermore, the proper positioning of
the assemblies was assured by only one, the so-called upper positioning plate.
Consequently, there was a real possibility that one or more of the 6th batch fuelassemblies were put inaccurately into their seat positions. In case of an entirely
dispositioned fuel assembly, it could result in a by-pass cross section of 13-20 cm2(see Fig. 2.7).

Figure 2.6. Working fuel assembly and its leg part [3]9Figure 2.7. Possible by-pass orifices from displacements
For evaluating the design deficiencies of the cleaning tank it must be mentioned that no
instrumentation (such as thermocouples, pressure transducers or water level measurements)
were installed inside the tank. Therefore, the operating staff could not recognize theinsufficiency of the cooling and the increase of the inner temperature.
2.3

Chronology of the incident

Unit 2 of the Paks nuclear power plant was shut down for refuelling on 29 March 2003. The
cleaning operation was carried out in refuelling pit of the spent fuel storage pool of unit No. 2.
Five batches of fuel assemblies with crud were cleaned without any indication of fuel failure.
Later investigation pointed out that the intermediate cooling mode was applied only for ashort time in most of the five cases and the tank was open very soon after the completion ofcleaning.
On the 10th April batch No. 6 was cleaned, the total decay heat of the 30 assemblies was
241 kW [4,5]. The chemical cleaning was completed at 16.00 and the intermediate cooling(operational mode ‘B’ in Fig. 2.5) was started at 16.40 (Fig. 2.8). The opening of the tank was
postponed, because the crane was used for other operations.
Course of the incident

Release ofradioactive
noble gas,
MBq/10i

Cleaning finished 16:00
Submergible pump starts 16:40
Cleaning circuit vacatedWater level of spent fuel pool rises 19:20
Increase of Kr-85 activity in the cleaning system 21:50

Cover could not removed
Near to spent fuel pool 12 mSv/h 02:45
Tank cover open, radioactive gas release 02:15

Ventilation starts 23:45
The CS evacuates the reactor hall 23:30

Figure 2.8. The course of the incident in time [5,6]10After 19.00 a slight increase of water level was recorded at the pressurizer. The change wasonly a few cm, but considering the common surface of the system it meant 4 m3 water
volume. As the total volume of the tank was 6 m3, the 4 m3 volume indicated that the mainpart of the fuel rods were in dry conditions after this elevation change.
The first indication of fuel failure was detected by the analyser device of the air letdown
system at 21.50 as activity increase of 85Kr. The operational team decided to open the tank tounderstand the cause of activity release.
The hydraulic interlock of the cover was opened at 2.15 after midnight. The cover removed
from its original position without lifting up and a large gas bubble emerged from the tank.
The detectors indicated high noble gas activity in the reactor hall and in the ventilation
chimney. At 4.20h an attempt was made to remove the cover, but the rope of the crane wasbroken. The tank remained in partially open position. The first activity concentration
measurements showed high (107-108 Bq/l) coolant activity in the spent fuel storage pool forseveral fission product isotopes. The cover was lifted on 16 April.
Thermal hydraulic calculations pointed out that the main reason for the insufficient cooling in
the cleaning tank was a design problem: both inlet and outlet junctions were located in thelower part of the tank and a by-pass flow could be formed between them [4]. Due to the low
flow rate of the intermediate cooling pump and because of the low Reynolds number in thebundle, the by-pass flow through the perforations in the assembly shroud and at the bottom of
the imperfectly seated assemblies became much more significant, than it was during thecleaning operation with high flow rate (Fig. 4.). The numerical analysis showed that theheating of the water led to saturation state in the top of the tank. The production of 4 m3 steam
could take place rather quickly, considering the high decay heat of the assemblies, and theformation of water level in the cleaning tank took only some minutes after saturation wasreached. The fuel behaviour calculations proved that the loss of fuel integrity in 5 hours after
the initiation of intermediate cooling mode was a result of fuel rod ballooning and burst [7].
The fuel rods could heat up above 800 ºC by that time and the high internal pressure in somerods resulted in plastic deformation and burst. The calculation of zirconium oxidation pointed
out that the volume of produced hydrogen could be much larger, than the volume of thecleaning tank. For this reason most of the hydrogen should have been released through the air
letdown valve. Since the produced hydrogen could be released from the tank only with sometime delay, the high temperature oxidation of fuel assemblies took place in hydrogen richsteam [8].
As there were no measurements inside of the cleaning tank the reconstruction of the event
involved many uncertainties. The following scenario (see Fig. 2.8) was put together from themosaics of numerical analysis, high temperature experimental work and observations of thefuel state [9, 10]. The Paks-2 incident contained three main phases.
2.3.1

Formation of steam volume

After the initiation of intermediate cooling mode with 20 t/h flow rate part of the coolant
could by-pass the fuel assemblies and did not take part in the removal of decay heat. Twopaths were identified for the by-pass flow (Fig. 2.7):•

The working type WWER-440 fuel assemblies have perforation holes in the shroud.

•

The perfect seating of the assemblies was not controlled and some gaps could exist
between the bottom of the assemblies and the lower plate.

11

The development of the thermal stratification and the increasing by-pass flow rate were notdetectable from the outlet temperature measurement signal. The above described process
conduced to the boiling of the coolant in two hours (see Figure 2.9.c). Then, the generated
steam pushed out the water from the bulk volume of the cleaning tank between 19:00 and
19:20. This process was shown by the increase of pressurizer level of the unit. However, it
was revealed only by the post-incident investigations.
The flow through the assemblies was not enough to remove the 241 kW decay heat and thetemperature in the upper part of the tank started to increase. When the saturation temperature
was reached the formation of a steam volume took place in short time. The subsequent heatup period lasted about 2.5 hours.

a

b

c

d

d

Figure 2.9. Flow path of coolant in the cleaning tank during normal cleaning operation (a),at the beginning of intermediate cooling mode (b) and the steps of steam volume formation(c,d,e)2.3.2

Plastic deformation of cladding and high temperature oxidation

Continuous temperature increase started in the tank, when the upper part of the fuel rods was
not cooled by water. The heat removal from the tank to the surrounding water was very low,
because the vessel was isolated by the double wall system. The temperature increase led to theincrease of pressure inside of the fuel rods. At 800-900 ºC the internal pressure could reach30-40 bars. In this range of pressure and temperature plastic deformation of the cladding cantake place and the ballooning can lead to bursts and activity release from the fuel. It is very
probable that this type of fuel failure was responsible for the activity release measured by the85
Kr detectors. (Very long ballooned areas were found in the later visual inspection of thefuel.) The first fuel rod ruptures were detected at 21:50, when the 85Kr measurement of the
AMDA system showed an unexpected jump in the signal (see Fig. 2.8) and a few minutes
later the noble gas detectors of the unit were also alerted.
The further increase of temperature accelerated the oxidation of zirconium components. The
maximum cladding temperature reached 1200-1300 ºC. The oxidation produced highhydrogen content in the stagnant steam volume. Most of the hydrogen could escape throughthe air letdown valve, but the high hydrogen concentration resulted in significant hydrogen
uptake by the cladding and shroud. This period lasted for 7 hours. The upper part of theassemblies became oxidised and picked-up large amounts of hydrogen. The bottom part wascooled by water and suffered no significant changes.
The temperature distribution in the vessel was non-uniform. Probably there were significant
differences between the temperature of peripheral and central fuel assemblies. Large variationof axial temperature can be supposed, since the bottom part of the fuel rods was cooled by

12

water. For these reasons the plastic deformation at 800-900 ºC and the intense oxidation
above 900 ºC could take place at the same time in the fuel assemblies.
As a result of hydrogen production in the Zr oxidation process the upper part of the cleaningtank was filled with steam-gas mixture. Probably most of the produced hydrogen was released
through the air letdown valve, but there was no reliable information on the operation of thisdevice and on the released gas flow rate.
2.3.3

QuenchThe severe fuel damage took place during the quench phase that was accompanied by large
activity release. No signs of temperature escalation or hydrogen production were found,probably because of the high degree of oxidation before quench. The quench was a
combination of top and bottom injection and the flow path had complex three-dimensional
picture.

a

b

c

d

Figure 2.10. The main phases of quenching: intermediate cooling mode with formation of bypass and 7 hours of oxidation (a), opening of the cover, gas release to the pool and water
injection from the bottom (b), cold water injection from the top (c), stable final state (d).
The possible quench scenario included the following steps:a) The opening of the hydraulic locks created a small gap between the cover and the vessel.
b) Some gas was released through the gap from tank into the spent fuel storage pool and thepressure in the tank decreased (Fig. 2.10b).
c) Water injection through the inlet line could reach higher elevations than before theopening of the cover due to low pressure.
d) Intense steam production took place when the cold water evaporated on the hot surface ofthe fuel. Probably this process resulted in vertical upward movements of the assemblies
that were recognised by the signs of interactions between the inner surface of the coverand the assembly heads. These signs were found after the removal of the cover.
e) The pressure increase in the tank led to the rising of the cover and increase of the gap
between cover and vessel.
f) The large gap size facilitated water penetration from the surrounding pool and so large
amount of water quenched the fuel assemblies from the top (Fig. 2.10c).
g) Finally the tank was filled up by water and stable circulation of coolant was establishedwith the pump (Fig. 2.10d).

13The thermal and mechanical loads resulted in the fragmentation of many fuel rods. Several
assemblies suffered full cross section break, many parts of shroud plates were missing withsigns of brittle failure. The high noble gas activity release indicated that many fuel rods losttheir integrity during quench.
2.4

State of damaged fuelAfter the incident detailed visual examination was carried out with the help of video cameras.
The examination indicated that most of the fuel assemblies suffered damage. Brittle failureand fragmentation of fuel assemblies was observed. Above the upper plate several assembly
heads were broken, standing in inclined position (Fig. 2.11). One assembly header was found
far from its original place. Many assemblies were broken and fragmented below the upper
plate, too. Some assemblies were fractured in their entirety. Fuel rod fragments and shroud
pieces accumulated on the lower plate between the assemblies. Many fuel rod pieces andfragments of assembly shroud were spread in the tank. Some fuel pellets fell out of fuel rods,
their form remained mainly intact. Heavy oxidation of the zirconium components was
identified. Less oxidation was found in the periphery than in the centre. The bottom part of
the fuel remained intact.
The visual investigations have also shown that the fuel assemblies positioned closer to thevertical axis suffered heavier damage, in some cases long parts were simply broken out fromthem. Thanks to the better position for the radiative heat transfer, the outermost assemblies
suffered less heavy damage. The broken fuel pins, shrouds and fallen down fuel pellets
formed a heap of debris on the bottom positioning and support plate.

broken headers

missing
headermissing
section offuelassemblydamaged fuelassemblies

accumulated
debris (fuel rod
segments, shroud
pieces)cleaning tankintact lower
section offuelassemblies

water of service pit

Figure 2.11. Distribution of damaged fuel in the cleaning tank
TABLE 2.1. ACTIVITY RELEASE FROM DAMAGED FUEL DURING THE INCIDENT
Isotope
133
Xe
131
I
106
Ru
134
Cs
137
Cs
140
Ba
144
Ce

14

Released activity
6.6e+14Bq
5.9e+14Bq
8.7e+12Bq
4.2e+13Bq
3.8e+13Bq
1.8e+14Bq
7.2e+13Bq

Relative release
1.19%
1.41%
0.03%
0.74%
0.53%
0.16%
0.06%

There were no signs of melting or formation of zirconium-steel eutectics on the surface ofstainless steel components. This fact indicates that the maximum temperature during theincident remained below ≈1400 oC.
The activity concentrations in the coolant and the release through the chimney are regularly
measured and such data were available after the incident, too. The incident happened two
weeks after reactor shutdown, for this reason the release of isotopes with short half-life wasvery low. Integrating the activity concentrations over time and coolant volume in the pool,and summarising the release through the chimney in time, the total activity release from thefuel was determined for several isotopes. Most of the activity remained in the water, since theincident took place under 13 m water level, only the noble gases were released through thechimney. The integrated activity release was compared to the calculated inventories and therelease rate was determined. In case of gaseous and volatile isotopes the release rate wasroughly 1% (the precise data are given in Table 2.1). The release rate of non-volatile isotopes
was much less. The ≈1% iodine, cesium and noble gas release indicated that the temperaturesin the cleaning tank could not be very high, otherwise larger release should have beenrecorded. Considering these release rates the maximum temperature was estimated about
1200-1300 oC. This temperature range can explain as well that the local oxidation reached100% in some positions.
The hot cell examination of the damaged fuel could not be carried out at the Paks nuclear
power plant, since the power plant does not have the necessary equipment and facilities forthe detailed investigation of irradiated fuel.
The very brittle state of the damaged fuel was observed during the removal operations.
Several fuel assemblies and fuel rods were fragmented when the damaged fuel was removed
from the cleaning tank and placed in the containers.
2.5

Post-incident investigations

The cleaning tank cover was removed on 16 April [5]. Until that time, the operators of the AMDA
system suspected only leakage from a few fuel rods. After the removal of the cover it has became
clear that almost all fuel assemblies suffered heavy damage. The characteristics of the damagereferred to that the insufficiency of the cooling would be the main cause. As to the extent of
the damage, the classification of the incident was changed to INES-3 from the previous ratingof INES-2 which classification was based on the radioactive release [11].

The subsequent criticality calculations [12] showed that the multiplication factor of thedamaged system was 0.595 by 20 g/kg boric acid concentration, so the system was in a deep
subcritical state following the incident. The sophisticated 3D reactor-physical analyses
pointed out that ensuring a minimal boric acid concentration of 16 g/kg, there had not been an
arrangement of the fissile material that could have become critical.
The subsequent calculations [4] have also revealed that the geometrical arrangement which
formed during the incident and the boric acid concentration level at the time of the incident
made the criticality impossible. The maximum value of the multiplication factor could be 0.66
during the incident. The radioactive release data also exclude the progress of uncontrolled
chain reaction.
The calculations based on measurement data showed that the extra doses will not exceed thevalue of 0.13 μSv for the most affected population group. This value corresponds to the dose
which accumulates from the natural background for 1.5 hours. On the other hand, the annual15limitation of the artificial dose for the population nearby the nuclear power plant is 90 μSv,
while the natural background caused annual doses are 2300-3000 μSv/person. On the basis ofthese values it can be stated that the extra doses from the incident are negligible.
As a consequence of the incident, a significant radioactive contamination occurred in only one
case: the face of an nuclear power plant worker was contaminated in such a way that it caused
an extra dose much below the annual dose limitation (0.059 mSv external and 0.55 mSv
internal doses [5]). Personal injury did not occur during the incident.
Between 16-25 June 2005, an examiner group mandated by the International Atomic Energy
Agency visited the Paks nuclear power plant in accordance with the Hungarian government’s
wish, in order to help to reveal objectively the causes of the incident. The investigation of thegroup pointed out that the main reason of the incident was the insufficiency of the cleaningtank cooling system and, among others, revealed the below listed design deficiencies [13]:•The submersible pump of the ‘B’ operational mode was under sized and also did nothave any redundancy or reserve.

•

The by-pass flows of the fuel assemblies were not taken into account during thethermal hydraulic design of the cleaning tank.

•

The further possibility for developing by-pass flow through orifices that could comebecause of the possible fuel displacements were realized but were not handled
effectively.

•

The measurement instrumentation, the parameter trend observation and the alarmsystem were not designed sufficiently.

The examiner group recommended several correcting measures for both the power plant andthe Hungarian regulatory. Among others they made suggestions connected to the licensing of
new technologies and methods, the improvement of the safety culture, the education of thestaff, development of emergency plans and the feedback of operational experiences. A strong
recommendation was that the safety of an important operating activity should not be given toa contractor without the supervision of the operator. The international examiner group also
proposed to strengthen the communication between the operator and the regulatory.
The performance of these recommendations was investigated by a new examiner group in
February 2005, and 71% was classified as ‘performed’ and the remaining as ‘adequately
progressing’.
2.6

The recovery work

For the recovery of the incident, the Paks nuclear power plant established a working group(Recovery Project) which is charged with the removal of the damaged fuel. This group waspreviously charged with the normalization of the state of the system, and the preparation forand licensing of the recovery work.
During the normalization of the system’s state the following main steps were made:•Separation of the refuelling pit with the damaged cleaning tank and the spent fuel poolfrom the reactor.

•

Increase of the boric acid concentration up to 20 g/kg in the refuelling pit.

•

Development of the safety borating system of the cleaning tank.

16•Construction of an independent cooling system of the cleaning tank.

•

Separation of the refuelling pit from the spent fuel pool.

•

Installation of redundant temperature, coolant level and neutron measurement
instrumentation in order to make the refuelling pit an independently operated system.

•

Visual exploration of the state and geometry of the damaged fuel assemblies and thecleaning tank in deep details.

In the summer of 2003, the Russian TVEL company was chosen for the execution of thetechnical tasks concerning the recovery of the damaged system.
During the removal, the pieces of the damaged fuel were picked up by manipulators which areoperated by a staff standing on the so-called working and lifting platforms which werepositioned into the refuelling pit. The picked up pieces were placed into special casks andcontainers which presumably will be stored in the spent fuel pool for five years.

Figure 2.12. Instruments installed into the service pit during the recovery work [14]The article [23] describes the technical equipments which were used during the recovery
work. Fig. 2.12 shows the different equipments which were installed into the refuelling pit.
The so-called middle flange and the protective flange (coloured with white and red in Fig.
2.12) were placed onto the split plane of the cleaning tank. The middle flange holds thetemperature and neutron measuring equipments which are placed into the cleaning tank. The
protective flange was a horizontal plate that was held by the consoles of the refuelling pit wall andit closed the domain between the split plane of the cleaning tank and the pit wall. Between thebottom and upper part of the refuelling pit the flow domain was connected by the TV shaft.
Among others, the protective flange contained a mechanical filter, and holding structures for
positioning 17 casks or containers constructed for storing the removed fuel assemblies andfuel pieces.

17For the time period of the removal process itself, the coolant level was decreased. Above thisdecreased level, the working and lifting platforms were placed (coloured by yellow and violeton the left side of Fig. 2.12). The working platform was hold by the reactor-podium, and it
could be set to different levels as it was needed. The bottom plate of this platform could be
rotated and a working aperture was constructed on it. The lifting platform was put onto therotating plate of the working platform. The lifting platform also had overlapping working
aperture. During the removal process, the working staff was standing on the lifting platform.
In case of emergency, the staff could exit on the ladder, or the platform could be lifted out by thereactor hall crane. The lifting platform was lifted out every time when the casks or containers
hold by the protective flange got filled and the refuelling machine transported them into thespent fuel cooling pond.
During the recovery work, first the upper positioning plate of the cleaning tank was cleaned
and removed. Then, the fuel assembly top parts were removed. The removing of the fuelassemblies followed a way from the inner parts to the outermost assemblies. For lifting out
the fuel assemblies, the leg parts had to be disengaged. After that, the proper seats on thebottom positioning and supporting plate were closed by special plugs. The leg parts were cut
off and the fuel pin batches were placed into the casks. After removing of the fuel assemblies,
the bottom positioning and supporting plate was cleaned. Then the elliptical bottom part of
the fuel tank was cleaned with vacuum-suction method. At the end, the cleaning tank was
chemically decontaminated and lifted out from the refuelling pit. Also the chemical
decontamination and the entire cleaning of the refuelling pit were necessary.
After the recovery work the damaged fuel is stored altogether in 24 wide and 44 narrowcontainers in the spent fuel pool of the unit 2 of Paks nuclear power plant.
3

MAJOR RESULTS OF ANALYSES

Ten organisations provided calculated data on their analyses of the Paks-2 event. Each partner
produced a short summary on the work performed and the results received with the following
content:•

short description of the code applied,•

short description of the input model, nodalisation scheme, specific options,•description of total decay power, power of assembly groups, axial power distribution,•

description of inlet flow (mass flow, temperature) and external (pool) temperature,•description of pressure boundary condition of cleaning tank.

It was requested to provide the chronology of main events considering the followings:•

saturation in the cleaning tank,•water level drop in the cleaning tank,•maximum cladding temperature exceeds 800°C,•

maximum cladding temperature exceeds 1200°C,•

start of intense Zr oxidation,•

failure of the first fuel rod due to ballooning,•

failure of the last fuel rod before quenching.

The short summary produced by each participating organisation is given in this chapter in thealphabetical order of the countries.

18

3.1

Participant SUEZ - TRACTEBEL ENGINEERING

3.1.1

Code description

Suez-Tractebel Engineering used the integral code MELCOR 1.8.5 to perform the calculationsin the framework of the ‘OECD-IAEA Paks Fuel Project’. MELCOR 1.8.5 is a severe accident
code developed by Sandia National Laboratories (SNL), which allows the user to perform
integral severe accident calculations. As a result, MELCOR 1.8.5 is not only capable tocalculate the thermal hydraulic behaviour of a facility, but also the possible degradation of thefuel assemblies and fission product releases.

Figure 3.1.1. Schematic representation of the nodalisation scheme of the Suez-Tractebel
Engineering MELCOR 1.8.5 model for the Paks Fuel cleaning tank.

19

3.1.2Model description

3.1.2.1 Nodalisation scheme
An overview of the nodalisation scheme can be found in Fig. 3.1.1. Control volumes 201, 202
and 203 represent the air-filled space between the inner and the outer vessel walls of thecleaning tank. The coolant enters the cleaning tank in control volume 101, from where it goes
through the fuel assemblies (represented by control volume 104) towards the cleaning tankhead (control volume 102) and the downcomer region (control volume 103).
In Fig 3.1.1, the flow paths between the different control volumes are indicated in blue. In
order to obtain a proper modeling of the transient phase, a fictive control volume has been
added (control volume 105). This control volume is situated between the downcomer region
and the outlet tube. The different bypass flow paths (holes and accidental bypass) inject in thiscontrol volume. The accidental bypass has a flow area of 0.012 m², while a total surface of
0.0083 m² has been attributed to the bypass flow through the holes at the bottom of the fuelassemblies.
3.1.2.2 Power distribution
Two axial zones (core and lower head zone) are defined for the core fuel assemblies. The firstzone, representing the lower head, contains three axial levels of the fuel assemblies. Twenty
axial levels have been modelled for the active core region. Based on the geometry in which
the fuel assemblies are placed inside the tank, the core has been divided in three concentric
radial rings. The tank core has been modelled as a BWR type reactor, in order to track thecontribution of the cladding and canister zirconium to hydrogen formation.
3.1.2.3 Axial level input
In the AEKI database, the axial power density profiles of all the assemblies, present in thecleaning tank at the moment of the incident, are given as a function of time. Since the axial
power density profiles of all the fuel assemblies are similar and since the evolution in time is
not significant, the axial power density profile of fuel assembly 1 at 30 hours is considered as
representative and is implemented in the model.
Ten axial power density profiles are considered in the AEKI database. In the MELCOR 1.8.5
model, 20 axial levels are considered. The additional power levels have been obtained bylinear interpolation. In TABLE 3.1.1, input data, required by MELCOR 1.8.5, are indicated.
3.1.2.4 Radial ring input
In TABLE 3.1.2, the input data for MELCOR 1.8.5 are indicated. The area of the radial rings isrepresented schematically in Fig. 3.1.2.
3.1.2.5 Inlet flow
The four inlet flow tubes have been modeled as flow paths only. Each of them has a crosssection of 9.2E-3 m². The mass flow rate has been fixed, throughout the entire calculation,and corresponds to a speed of 0.155 m/s (20,45 ton/h). After a series of preliminary
calculations, it has been chosen to set the inlet temperature to 60°C.

20

Active core region

Lower
plenumTABLE 3.1.1. MELCOR 1.8.5 AXIAL LEVEL INPUT.
Axial level
[-]
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23

Lowest point
[m]
0.359
0.521
0.683
0.756
0.91481.07361.2324
1.3912
1.55
1.70881.86762.02642.18522.344
2.50282.66162.82042.9792
3.138
3.29683.4556
3.61443.7732Height
[m]
0.162
0.162
0.073
0.15880.15880.15880.15880.15880.15880.15880.15880.15880.15880.15880.15880.15880.15880.15880.15880.15880.15880.15880.1588Corresponding HS
[-]
10231
10232
10233
10211
10212
10213
10214
10215
10216
10217
10218
10219
10220
10221
10222
10223
10224
10225
10226
10227
10228
10229
10230

Rel. power dens.
[-]
0
0
0
0.3841
0.48780.59140.69510.81410.9331
1.027
1.127
1.89
1.257
1.293
1.329
1.333
1.337
1.303
1.27
1.164
1.057
0.83730.6171

Ring 3
Ring 2

Ring 1Figure 3.1.2. Schematic representation of the arrangement of the fuel assemblies in the Paks
cleaning tank. The boundaries of the three concentric radial rings are indicated in the figure.

21

TABLE 3.1.2. MELCOR 1.8.5 RADIAL LEVEL INPUT
Radial ring
[-]
1
2
3
3.1.3

Cross section
[m²]
0.38
0.65
0.68

Heat structure
[-]
10290
10290
10290

Power per ring[kW]
42.95493.348105.336

Chronology of main events
TABLE 3.1.3. CHRONOLOGY OF MAIN EVENTS
Event

3.1.4

Timing [s]Saturation in the cleaning tank

8520

Water level drop in the cleaning tank

8520

Maximum cladding temperature exceeds 800°C

16250

Maximum cladding temperature exceeds 1200°C

20110

Start of intense Zr oxidation

16610

Failure of first fuel rod due to ballooning-

Failure of first fuel rod before quenching-

Conclusions

Suez-Tractebel Engineering performed two sets of calculations with regard to the ‘OECDIAEA Paks Fuel Project’. The first set (blind calculation) has been presented at the final
project meeting. The thermal hydraulic results of this calculation are in agreement with theobservations during the incident.
The second set of calculations (open calculation) has been performed after the final project
meeting. These calculations take into account the reflections made during and after the final
discussion with the other participants of the project. The open calculation presents not only agood agreement between the thermal hydraulic behaviour in the tank and the observations
during the incident, but also shows that the degradation of the fuel assemblies follows a clear
physical logic, which corresponds with the information available from the cleaning tankfacility:
(1)
(2)
(3)
(4)
(5)

Start of uncovery of the fuel assemblies (2h20’),
Start of fuel assembly degradation (4h35’),
Amount of core degradation (fission product releases),
Amount of core degradation (maximal cladding temperature – melting of the cladding
material),
Heat transferred out of the facility with respect to the calculations performed by the
University of Budapest.

It should be mentioned that certain elements of the thermal hydraulic behaviour of the TE
open calculation are in conflict with those of other participants of the program. The main22

differences lay in the reduction of the water level in the downcomer, more in particularwhether or not the outlet tube becomes uncovered, and the amount of heat which is evacuatedout of the cleaning tank by radiation.
However, it is difficult to state which of both types of calculations reflects best the actual
behaviour in the cleaning tank since only the information with regard to the fuel degradation(cladding temperature, fission product release) is available to evaluate the calculations. For
both types of calculations, the behaviour of the fuel and the amount of fuel degradation are
similar.
More detailed description of the Suez - Tractebel Engineering analysis is available inAppendix B/I.
3.2

Participant VTT

3.2.1

Introduction and case description

3.2.1.1 Paks-2 cleaning tank incident
The fuel performance simulation began from the instant when the cooling mode of thecleaning tank was initiated and the fuel rods were being cooled by circulation of storage poolwater. The cooling of the fuel assemblies was insufficient due to significant by-pass coolant
flow which was formed by fuel assembly perforations and some incorrectly seated fuelassemblies, which led to increasing fuel rod cladding temperatures. After 5 hours and 10
minutes an abrupt increase in Kr-85 counts was observed. After a 9.5 hour cooling period, theunderwater cleaning tank was opened, and the hot and heavily oxidised fuel rods werequenched by the cold refuelling pit water and were shattered.
3.2.1.2 Tools for Paks-2 cleaning tank incident analysis
Thermal hydraulic analysis of the Paks-2 fuel cleaning tank incident was performed with the
APROS-5.06 simulation code. Total of 30 fuel assemblies in the cleaning tank were divided
into six groups by their decay heats.
The transient fuel behaviour during the incident was analyzed with the fuel performance code
FRAPTRAN-1.3 with VTT modifications. The fuel behaviour analyses cover a representativerod from each of the groups into which the assemblies had been divided by their decay heatlevels, and the time period between the start of the cleaning tank cooling mode and the instantwhen the cleaning tank cover was opened. The steady-state fuel performance code
FRAPCON-3.3 was used to calculate the burnup dependent properties for the transient fuelperformance code FRAPTRAN.
The OECD-IAEA Paks fuel project database was used for the steady-state and transient fuelbehaviour calculations. The thermal boundary conditions for the transient analysis were
calculated with the APROS system code. The cladding oxidation and heat generation from themetal-water reaction were calculated with FRAPTRAN by using a conservative Baker-Justoxidation correlation.
A simple perl-script set was implemented to transfer initial and boundary condition data fromthe OECD-IAEA Paks Fuel Project database and from the APROS calculation results. A
second perl script set was implemented to transfer calculated heat generated by zirconium
oxidation with FRAPTRAN back to APROS. Total of three iteration runs were performed

23with APROS and FRAPTRAN in turns. The goal of this iterative solution method was toachieve maximum reaction heat generation by oxidation corresponding on the decay heat of
30 FAs in the cleaning tank.

3.2.1.2.1 APROS simulation model
The main APROS simulation model parameters:−

Thermal hydraulic phenomenon modelled with 6-equation model.

−

Radiation heat transfer between solid surfaces and solid surface and gas.

−

The fuel assemblies were divided into three concentric rings which was taken into accountin defining the viewing factors for radiation heat transfer modelling.

−

Constant heat transfer coefficients were used on outer wall of the vessel (vessel wall 1280
W/m2/K, vessel cover 1450 W/m2/K) when calculating the heat losses to the storage pool.

−

The total decay heat of the assemblies was 241 kW.

−

The fuel rods were axially divided into 23 or 24 nodes depending on the type of the fuelassembly.

−

Downcomer side of the vessel was divided into 12 nodes.

−

The flow area of the perforations was 83.3 cm2 and an extra by-pass area was 120 cm2.

−

Centre of the cleaning tank was located 10 meters under water. The end of the outlet pipe
was located at the top of the spent fuel pool.

−

Inlet temperature was 30 oC and inlet mass flow was 5.8 kg/s.

−

Initial cleaning tank temperature was 58 oC.

3.2.1.2.2 FRAPCON and FRAPTRAN fuel performance codes
FRAPCON is a single LWR fuel rod fuel performance code designed for steady-statecalculations and to generate initial input conditions for FRAPTRAN.
The main phenomena and output parameters with FRAPCON-3 include:1) radial burnup-dependent heat conduction through the fuel and cladding,2) cladding elastic, thermal, creep, and plastic deformations,3) fuel cladding mechanical interaction (PCMI),4) fission gas release (FGR),
5) fuel rod internal gas pressure,6) radial heat transfer between fuel and cladding,7) cladding oxidation, and
8) heat transfer from the cladding to the coolant. The code contains necessary material
properties, water properties, and heat transfer correlations.

24

The cladding elastic-plastic deformation model based on finite element (FE) approach havebeen implemented in FRAPCON and FRAPTRAN at VTT.
FRAPTRAN is a transient fuel performance code designed to calculate the response of asingle LWR fuel rod to transient power and/or coolant conditions.
The phenomena modelled by FRAPTRAN include:1) radial heat conduction,2) heat transfer from cladding to coolant,3) elastic-plastic cladding deformation,4) cladding high temperature oxidation, and5) fuel rod gas pressure.
The E110 (Zr-1%Nb) material properties and the clad rupture model for E110 clad material
have been implemented in FRAPTRAN at VTT.
The standard fuel rods and the follower type fuel rods were divided in 23 and 24 axial evenly
spaced nodes, respectively.
3.2.1.3 Initial and boundary conditions
The fuel assembly and rod design parameters the WWER-440 type fuel construction is basedon the fuel design information that is provided in OECD-IAEA Fuel project database.
The LHGR profiles at the beginning of the cooling mode are presented in Figure 3.2.1. The
decay heat profiles of FA groups 1-2 and 5-6 are cosine shaped while the decay heat power
profiles of Group 3 and 4 are more peaked towards the upper end of the rod.
The fuel assemblies were divided into six groups according to their levels of decay heat. The
resulting division of the fuel assemblies in six groups is given in Table 3.2.1.

Figure 3.2.1. Relative decay heat profiles at the beginning of transient.

25

TABLE 3.2.1. FUEL ASSEMBLY GROUPS

3.2.2

Fuel assemblies

Group

Average decay heat
[kW/m]

1-6
7-11
12
13-18
19-24
25-30

1
2
3
4
5
6

2.96
3.72
3.51
2.89
3.81
3.76

Calculation results

Calculated maximum fuel pellet and cladding outer surface temperatures of the FA groups areclose to each other at the end of the simulation. The maximum fuel pellet and cladding outersurface temperatures during the transient are between 1866 and 2043 oC, and between 1628
and 1582 oC, respectively. The reaction energy release is highest in fuel assembly group 3 due
to peaked decay heat profile and the highest clad outer surface temperature. The calculated
maximum cladding temperature without oxidation modelling was 1155 oC.

Figure 3.2.2. Thermal hydraulic and fuel behaviour calculation results.

26

The FRAPTRAN transient calculations suggest that the metal-coolant chemical energy
produced is quite significant in the Paks-2 cleaning tank incident. The calculated chemical
reaction heat is produced from oxidation of the fuel rod cladding only, the fuel assembly
shroud, spacer grids etc. were not considered. FRAPTRAN calculations with the Baker-Justoxidation correlation predict that a significant chemical reaction between the zirconium
cladding and steam starts after 3 h 50 min and reaches its maximum power of 441 kW at 5hours 57 minutes into the cooling mode. The cladding burst is predicted to occur first in fuelassembly 12 after 4h 34 min from the beginning of simulation, 36 min before the measured
activity increase.
By the end of transient, 46.7% of zirconium has been oxidised and 15.1 kg of hydrogenproduced.
TABLE 3.2.2. MAIN EVENTS OF SIMULATION IN CHRONOLOGY ORDER.

3.2.3

Time [s]Comment2h:1minSaturation in the cooling tank.

2h:20minWater level drop in the cleaning tank.

3h:50minStart of high-temperature Zr oxidation.

4h:7minMaximum cladding temperature exceeds 800 oC.

4h:34minFailure of the first fuel rod due to ballooning (FA 12).

5h:5min

Failure of the last fuel rod before quenching (FAs 1-11).

5h:13minMaximum cladding temperature exceeds 1200 oC.

Conclusions

The incident in the cleaning tank of the Paks-2 nuclear power plant was analyzed with the
APROS-5.06 simulation code and the FRAPTRAN-1.3 transient fuel performance code with
FRAPCON-3.3 initialisation.
The voiding of the cleaning tank was well predicted by APROS. Before the voiding, it waspredicted that 35% of the coolant flow has occurred through the perforations and 55% had
flown through the by-pass formed by incorrect seating of fuel assemblies. As a result of this,
only 10% of the inlet flow cooled the fuel assemblies. Inadequate cooling led to voiding of
the coolant in the cleaning tank and to overheating of the fuel rods.
Fuel failure was envisaged for all six fuel assembly groups. The cladding burst was estimatedto happen at first in the fuel assembly number 12 at 4 h 34 min from the beginning of thesimulation, rods in all other groups were predicted to lose their integrity within eight minutes.
The cladding burst was estimated to happen 36 minutes earlier than it was detected. This
result indicates that cladding temperatures and/or cladding oxidation were slightly
overestimated.

27

The top and middle parts of the fuel rods were estimated to deform and become almost fully
oxidised while the bottom parts of the assemblies would seem to remain intact. In addition, asignificant amount of hydrogen is expected to be produced by zirconium oxidation.
More detailed description of the VTT analyses is available in Appendices B/II and B/III.
This research project was funded by Fortum Nuclear Services Ltd., Finland.
3.3

Participant GRS

The first analyses of the Paks incident with ATHLET-CD were already performed in 2003. In
the frame of the OECD-IAEA Paks fuel project the data basis and code models have beensuccessively improved as well as weaknesses and errors in the code detected and eliminated.
3.3.1

Code description

The system code ATHLET-CD (Analysis of Thermal hydraulics of LEaks and Transientswith Core Degradation) is designed to describe the reactor coolant system thermal hydraulicresponse, core damage progression, fission products and aerosol behaviour during severe
accidents, to calculate the source term for containment analyses, and to evaluate accident
management measures. The development and validation of ATHLET-CD are sponsored by
the German Federal Ministry of Economics and Technology (BMWi).
3.3.2

Files delivered for the comparison

The calculation described in this report has been finished on 19 May. The computing time of
the 9 h transient was 35 h. The integral calculation covers the thermal hydraulics, the fuelbehaviour, and the activity release. The calculation results were post processed to create thetables to be delivered for the benchmark. According to the specification a set of 6 tables were
produced:
(1)
(2)
(3)
(4)
(5)
(6)

Histories of thermal hydraulic parameters (29 variables as function of time).
Profiles of thermal hydraulic parameters by the time the maximum cladding temperature
exceeds 800°C and at t = 32400s (12 variables as function of elevation).
Histories of fuel parameters (23 variables as function of time).
Profiles of fuel parameters before quenching (24 variables as function of elevation).
Histories of release from the fuel rods (5 variables as function of time).
Profiles of axial power distribution (6 variables as function of elevation).

The transient data consist of 2957 records with a time increment of approximately 10 s. The
axial profiles consist of 22 records.
3.3.3

Nodalisation and input model

The nodalisation scheme is presented in Fig 3.3.1. Besides the bottom and top region with
inlet and air letdown, the thermal hydraulic system is modelled by 6 parallel flow channels
with 24 axial nodes. Four flow channels simulate the flow within the assemblies. The
surrounding fluid is modelled by two radial interconnected channels. The by-pass flows are
simulated by two flow paths through the penetrations from the fixed assemblies to thesurrounding fluid as well as one flow path due to incorrect seating. The cleaning tank outlet
flow is modelled during the first 3 h by a pipe with constant pressure boundary condition(2.3 bar) and during the later 6 h by a controlled mass flow to limit the flow oscillations in the28system. The inlet flow rate and temperature are 6 kg/s and 30 °C. The initial temperature in
the tank was set to 58 °C.
Heat losses to the water pool are simulated only over the double-walled cylinder barrel (waterpool temperature 30 °C). The 30 fuel assemblies are modelled by 4 representative
components, each two for the fixed assemblies and for the followers. The rods and the
shrouds are subdivided axially in 20 nodes with equidistant spacing of 0.122 m in the heatedregion and one node with 0.320 m and 0.240 m length for the lower and upper unheated partrespectively. The number of assemblies per component is 3, 6, 9, and 12 and the bundle
power is 7101, 8932, 6847, and 8707 W/assembly. The total decay power is 241 kW.
3.3.4

Results of calculation

The mass flow through the air letdown decreases from 2.0 to 0.4 g/s over the time from 4 to
8 h. The average hydrogen generation rate is limited to 0.33 g/s, the total hydrogen generationis 4.63 kg. At the end of the calculation the maximum temperature is 1380 °C at almost
stationary conditions (heat up rate 23 K/h). The temperatures of the outer most assemblies are
350 K lower, i.e. the calculated temperatures lie in the expected range.
The chronology of the main event is in the order of appearance:
TABLE 3.3.1. CHRONOLOGY OF THE MAIN EVENTSEventTimingBeginnig of saturation in the cleaning tank

7060 sBeginning of water level drop in the cleaning tank

8640 sBeginning of core super heating

9140 sMaximum cladding temperature exceeds 800 °C

14690 sStart of intense Zr oxidation

15840 sRod failure in ROD2

16036 sRod failure in ROD3

16434 sRod failure in ROD1

16966 sRod failure in ROD4

19842 sMaximum cladding temperature exceeds 1200 °C

21240 s

The timing of the water level drop is well captured by the calculation, but the release isestimated too early by 40 min.
The maximum oxide layer thicknesses are in the range of 100 to 500 µm except the cold wallsof the outer most shrouds. The total relative oxidation of all claddings and shrouds is only
5.3% due to the steam starvation in the cleaning tank.
The relative release (released mass / inventory) of fission products and of the fuel at the endof the calculation is 7.68E-3 for Xe and Cs, 6.16E-3 for I, 3.05E-4 for Te, and 8.51E-7 for U.

29

The end state of the simulation is depicted in Figure 1. The colours indicate the temperatures(yellow > 1300 °C, red > 600 °C, blue saturation) and the void (white = 1, blue = 0) in thesystem. The water level is at the same height in the whole system at about 0.7 m elevation.
Furthermore the ballooning of the cladding and the oxidation profile is visualized. Most
oxidation layer are too thin for proper display, only in ROD3 (FOLLOW-1) the oxide layer isvisible just below the maximum ballooning.
The conclusion is that the Paks incident was well predicted and the most processes were
simulated adequately by the code. The estimated maximum temperatures are sensitive to thehydraulic simulation, the air letdown flow and the heat losses to the water pool.

Figure 3.3.1. Nodalisation scheme and end state of simulation with ATHLET-CD.
More detailed description of the GRS analysis is available in Appendix B/IV.
3.4

Participant AEKI

3.4.1

Thermal hydraulic analysis

3.4.2

Thermal hydraulic analysis with the RELAP code

3.4.2.1

Code description

The RELAP5/MOD3.3 code has been developed for best-estimate transient simulation oflight water reactor coolant systems during postulated accidents. The code models the coupledbehavior of the reactor coolant system and the core for loss-of-coolant accidents andoperational transients such as anticipated transient without scram, loss of offsite power, lossof feedwater, and loss of flow. A generic modeling approach is used that permits simulating avariety of thermal hydraulic systems. Control system and secondary system components areincluded to permit modeling of plant controls, turbines, condensers, and secondary feedwater
systems.

30

3.4.2.2

Nodalisation

Two different nodalisations have been developed for the analysis. A more detailed one
lumped the 30 fuel assemblies into 4 groups, but non-physical oscillations were produced bythis one as soon as the upper plenum turned to two-phase conditions. Therefore a simple
nodalisation was applied, as shown by Fig. 3.4.1, where all assemblies were represented by asingle flow channel. The main parts of the cleaning tank are modelled as follows:
Vol 4, Vol 6 and Vol 590 = pool water volumeJun 5 = pump for cooling mode A
Jun 7 = pump for cooling mode B
Vol 58, Vol 60 and Vol 62 = single flow channel representing 30 fuel assemblies
Vol 570 and Vol 500 (1 –4) = upper part of tankVol 60 (2 – 6) = 5 nodes for active part of fuel rods (5 × 0,50 m). In the calculation
presented the active part of the fuel rods was divided to 20 nodes: Vol 60 (2 – 21), i.e.
20 × 0,125 m.
Hs 100 – 106 = average heat slabs
Hs 110 – 116 = hot pins

Figure 3.4.1. Nodalisation scheme for RELAP5 code.

31

The following by-pass areas were considered:incorrect seating of assemblies = 120 cm2 Jun (Vol 54 - Vol 576)
perforation (below active part) = 84 cm2 Jun (Vol 58 - Vol 578)
perforation (above active part) = 84 cm2 Jun (Vol 62 - Vol 572)
3.4.2.3

Boundary conditions

Cleaning tank inlet flow rate was imposed to represent the different cooling modes, with a
5 min. interruption in forced flow due to transition from one mode to the other:time < -300 sec
47.20 kg/s,-300 < time < 0 sec
0.00 kg/s
time > 0 sec
5.83 kg/s
Initial tank temperature 57 ºC.
Pool temperature 30 ºC.
Total fuel assembly power 241 kW.
3.4.2.4

Results

The analysis starts from stabilised conditions in the cooling mode C, with high forced flow
through the tank. A 5 min. period was assumed before effective switch-over to the low flowrate of the cooling mode B was established, where flow rate to the tank was zero. In thisperiod natural circulation was established in the tank with upward flow through theassemblies, downward in the tank and back via the perforations and by-pass area due toincorrect seating.
With the low flow rate of cooling mode B the flow rate entering the heated part of theassemblies is continuously decreasing, while the by-pass flow rate is continuously increasing.
As a consequence, the coolant temperatures start to rise. Due to the rather fast decrease of theflow rate, the saturation temperature is reached at the assembly outlet already at 2830 s. In
spite of that, it is not before 10560 s that significant level decrease can be observed in thecleaning tank, which is much later then the phenomenon observed during the incident, where
about 2 h 20 min after change-over to cooling mode B the level of the pool rose by ~7 cm incca. 20 min. At about 11600 s the level finally stabilises at an elevation of 1.6 m. Alreadybefore the level stabilisation fuel temperatures start to increase and by the end of thecalculation the maximum temperature reaches 600 oC.
Results for the main parameters are presented in Figs. 4.1 to 4.15 as the parameter AEKI.1.
(Additional figures can be found in Appendix B/V.) Due to the simplified nodalisation theinlet flow rates in Figs. 4.8-9 and 4.10-11 were estimated in the following way:
Inlet flow rate for working assembly = [Jun (54 – 58)- 19*Jun(58-60)/30]/11
Inlet flow rate for follower assembly = Jun (58 – 60)/30
The sequence of the events can be summarised as follows:< -300 s 47,2 kg/s forced flow
< 0 s natural circulation is established via the perforations and by-pass> 0 s 5,83 kg/s forced flow, heat-up to saturation
2830 s void at outlet of assemblies
10560 s vessel liquid level decreases
10730 s fuel cladding temperature increases
13650 s end of calculation

32

3.4.2.5

Conclusions

Saturated conditions are reached too early in the transient due to rather fast decrease of thefuel assembly low rate. Since elevation heads are correctly calculated in RELAP, it isassumed that this is a consequence of inappropriate friction losses in by-pass areas and /orfuel assemblies. On the other hand, the accumulation time of steam in the upper head isoverestimated that may be due to non-physical oscillations caused by transition from low tohigher void fractions in the condensation models.
3.4.3

Thermal hydraulic analysis with the ATHLET code

In order to understand the physical processes taken place in the cleaning tank during the Paks2 incident, in April/May 2003 a series of calculations were performed by AEKI for different
batches of fuel assemblies cleaned in the tank. The thermal hydraulic analyses were
performed by the ATHLET code. In what follows, the ‘best estimate’ calculation results arepresented for Batch No. 6, which is of interest for the present project.
3.4.3.1

Code description

ATHLET is a one dimensional two fluid code for the thermal hydraulic analysis of reactor
transients and accidents. The physical model of the code is based on the system of six integral
balance equations. Application of finite volume method for spatial discretization allows arather coarse nodalisation of modelled objects. The system of governing equations is closed
by state equations, interfacial transfer conditions and constitutive equations. Non-condensable
gas as well as boron dilution can be simulated by ATHLET.
3.4.3.2

Nodalisation

The geometrical model included the following elements as represented by Fig. 3.4.2:
•
•
•
•
•

Simplified container geometry with lower and upper plena and outlet piping. Inlet
piping was only modelled by imposed pump flow rates.
Container wall, fuel assemblies (FA) including shrouds, support plates modelled asheat structures.
FAs grouped into 5 parallel channels with 20 axial nodes
Tank volume outside the assemblies represented by a single channel with 20 axial
nodes
Shroud perforation of working FA at bottom and top (A= 7.634 cm2)

V-UP
V-AV1-OU

V-TART-UP
V-INTJ21

V-AV2-OU

V-FO1-OU

V-FO2-OU

V-FO3-OU

V-FO1

V-FO2

V-FO3

V-FO1-IN

V-FO2-IN

V-FO3-IN

V-INTJ22

P-KI

V-TART

V-TART-LP
V-LP-SJP1

V-LP-SJP

V-AV1

V-AV2

V-INTJ11

V-INTJ12

V-AV1-IN

V-AV2-IN

V-LP

P-BE

Figure 3.4.2. ATHLET nodalisation of the cleaning tank.

33

3.4.3.3

Initial and boundary conditions

Initial and boundary conditions were selected on the basis of Batch No. 6 data:
•
•
•
•
•
•
•
•
•
3.4.3.4

Total decay heat: 241 kW
19 follower FAs represented by 3 groups with different power
11 working FAs represented by 2 groups with different power
Axial power profile: from decay heat calculations for Batch No. 6
Pump flow rate: 170 t/h in cooling mode C and 21 t/h in cooling mode B, withinstantaneous switch-over from mode C to B
Cleaning tank initial temperature: 57°C, uniform
Pool temperature: 30°C
Degassing line not modelled explicitly, only constant outflow of 0.0015 kg/s assumed.
Besides the shroud perforation of the working FAs an additional by-pass cross section
due to incorrect seating of 117.5 cm2 was assumed.
Results

The most important results are graphically presented in Figs. 4.1 to 4.14 as the parameter
AEKI.2. Additional figures can be found in Appendix B/VI.
The mass flow rate at the inlet of the working assemblies is continuously increasing, theopposite is true for the follower ones. However, looking at the flow rates entering the heatedpart of the assemblies, i.e. above the elevation of the shroud perforation, all the flow rates
display similar behaviour, which means that in the case of the working FAs the difference
between the two flow rates exits via the shroud perforations. The flow rates of the different
assemblies are slightly different due to the different power. The by-pass flow rate iscontinuously increasing and reaches at 10000 s almost 75% of the total flow at the FA inlet.
As a consequence, the coolant temperatures start to rise. The saturation temperature is reachedat the outlet between 5400 and 6400 s, depending on FA power. The void formation under thetank cover starts at 6500 s and by 8000 s it is completely voided. It takes about a further 1000s that the level in the tank drops to roughly the lower 1/3 elevation of the tank. This is in good
agreement with the phenomenon observed during the incident, where about 2 h 20 min after
change-over to cooling mode B the level of the pool rose by ~7 cm in cca. 20 min.
A number of parametric calculations were also performed in order to assess the impact of themost important parameters: an overview is given in App. B/VI.
3.4.4

Fuel behaviour calculation

3.4.4.1

Code description

Fuel behaviour during the incident was simulated with the FRAP-T6 code. The code
simulates the thermal and mechanical behaviour of one piece of fuel rod considering the heattransmission of the given subchannel. The AEKI version of the FRAP-T6 code includes
special WWER models (phase transition, oxidation, burst, mechanical properties as functionof oxidation) that are based on experiments with E110 type cladding. According to thethermohydraulic calculations all coolant flowed out at the bottom of the fuel assemblies
without cooling the active section.

34

In the FRAP-T6 calculations the coolant circulation flow rate was set to constant in thesubchannel and other heat transfer methods (e.g. radiation) were not taken into account. The
enthalpy of the coolant at the entry was 167 kJ/kg. The model included one fuel rod in a onedimensional nodalisation.
The fuel assemblies with similar power histories and of the same type (follower or working)were grouped into 6 groups and only one representative assembly was calculated for eachgroup. The average linear power, the burnup, the pressure in the fuel rod and the axial
distribution were taken from the database. The calculations were made in the 10 evenly
spaced axial part of the fuel rod.
3.4.4.2

Results

After the formation of steam volume in the cleaning tank the temperature of upper part of
cladding exceeded 800 °C within 130 minutes. The fuel rods ballooned up and bursted, andthe accumulated fission gases were released into the coolant. The burst of fuel rods occurred
about 2.7 hours after steam volume formation. It is likely that this process was detected by
activity detectors after 21.30 at the Paks nuclear power plant.
After bursting, the temperature of fuel rods increased and the maximum cladding temperature
exceeded 1200 °C. The thickness of zirconium-oxide layer on cladding growed quickly
because of the intensive oxidation. Its local value was 300 μm, after 5-6 hour oxidation.
Time of failure of every assembly was calculated by the FRAP-T6 code and those data wereused in the activity release calculations.
3.4.5
3.4.5.1

Calculation of activity release
Model description

The calculational method estimates the release from the fuel rods in the percentage of isotope
inventory. The method makes a distinction between release from the gap and pellet. First the
gap inventory releases from the damaged fuel assemblies and then one part of the radioactive
isotopes (that are located in the pellet) with different mechanisms, like fragmentation of pellet
or leaching.
Parameters of the best-estimate method were derived from processing of different
experimental data. This method is applied in the safety analyses of design basis accidents suchas LOCA.
3.4.5.2

Calculated results

Apart from the noble gases 10% of the total release is taken to be released in the dry phaseand 90% is released after the rupture is recovered. All the noble gas release is taken to occur
in the dry phase.
The calculations of activity releases were performed for every assembly. The fuel assemblies
with similar power histories and of the same type (follower or working) were grouped into 6
groups. In the next table and figure it could be seen that good agreement was found between
calculated and measured activities.
In case of volatile isotopes good agreement was found between the measured activity release
and results of calculations. The most significant difference appeared for non-volatileelements.

35

Measured data

Calculated data

1.00E+14
1.00E+12
1.00E+10

e13
3
X

La
-1
40

I13
1

1.00E+08
1.00E+06
1.00E+04
1.00E+02
1.00E+00
Ba
-1
40
C
e14
4
C
s-1
34
C
s-1
36

Activity release [Bq]1.00E+16

Figure 3.4.3. Comparison of measured and activity data.
3.4.6

Summary of calculated results

More detailed description of the AEKI analyses is available in Appendices B/V, B/VI, B/VII
and B/VIII.
TABLE 3.4.1. CALCULATED TIMES OF MAIN EVENTS
thermal hydraulic calculation
saturation in the cleaning tankwater level drop in the cleaning tankfuel behaviour calculation
maximum cladding temperature exceeds 800°C
maximum cladding temperature exceeds 1200°C
start of intense Zr oxidation
failure of the first fuel rod due to ballooning
failure of the last fuel rod before quenching
3.5

RELAP52830 s10560 sATHLET
5400 s8000 s
FRAP-T6
15800 s17720 s17000 s18850 s20010 sParticipant BME NTI

On 10-11th of April, 2003 a serious incident caused heavy damage of 30 fuel assemblies in acleaning tank which was installed into the refuelling pit of the 2nd unit of the Paks nuclear
power plant. The incident was investigated by several scientific institutes and with different
analytical and numerical methods and tools. Among other methods, the early stage of theincident when one-phase flow was in progress was investigated with 3D CFD calculations in2003 in the BME NTI. In the same institute newer 3D CFD calculations were performed on
this stage of the serious incident in the framework of the OECD-IAEA Paks-2 Fuel Project in2007. Robust meshing, detailed geometry and the best known real parameters were built into
the model. By modifying the model’s geometry the effects of displacements of one or two
working fuel assemblies were also investigated.
3.5.1

The 3D CFD models of the cleaning tank

The 3D CFD model of the cleaning tank contained 11 working and 19 follower fuelassemblies. The assemblies were modelled as hexagonal prisms with a coaxial cylinder cut
out from it. The cylinders’ cross section was equal with the assemblies’ flow cross section.

36

The 12 perforations on the bottom part of the working assemblies were replaced by 6
perforations of 13 mm in diameter per each assembly. The 12 perforations on the top partwere replaced by 2 perforations. The leg parts of the working assemblies were modelled in a
detailed manner.
Three different versions of the model were developed. In the W11_F19_DIS1 and
W11_F19_DIS2 model versions displacements of one and two working fuel assemblies weretaken into account such a way that apertures have been opened at the corresponding
assemblies’ very bottom end. The cross section of the apertures was 2030 mm2. The
W11_F19_DIS0 model version did not contained any displacements.
The 4.1 million control volumes mesh contained unstructured tetrahedral mesh withhexahedral core combined with extruded pyramid elements.
3.5.2

Boundary conditions, parameters of the model

For all calculations the laminar flow model was used. The buoyancy was taken into accountwith the Boussinesq-approximation. All walls were modelled as free-slip walls.
The fuel assemblies were divided into six groups by its decay heat power profile. The decay
heat power density profile inside the FAs’ models was given as point slope functions. The
data set for the heat power profiles was given on the basis of the OECD-IAEA Paks-2
project’s database. Different volumetric quadratic resistance values were set inside theworking and follower fuel assemblies based on literature data.
The inlet parameters were constant static pressure and 36 °C inlet temperature. The outlet
setting was constant mass flow rate of 5.555 kg/s. All calculations were run as transient
calculations of 8,000 seconds. The time-step was set to be 2.5 s. The sensitivity for the time
stepping was investigated. The aim of the runs was investigating the one-phase thermal
hydraulic processes during the ‘B’ operational mode which started at 16:40, 10 April, 2003.
3.5.3

Calculational results

The calculations for the different model versions gave qualitatively very similar results. These
characteristics are summarized in the following bullets and the most important numericalresults are summarized in Figure 3.5.1.

Figure 3.5.1. Left side: By-pass rate through the working FAs’ bottom perforations and thedisplacement orifices and mass flow rate through the active parts Right side: Maximum
temperature values, whole flow domain volume average temperatures and mass flow weighted
temperature averages for the outlet.

37

(1)

(2)

(3)

The temperature stratification developed from the very beginning. After 1200 seconds
such temperature fields developed that it kept its characteristics for the further part of
the transients: The tank outlet temperatures froze at about 41-44°C temperature values
which means that the decay heat power was not removed from the cleaning tank for anyof the calculations since the differences between inlet and outlet temperatures were lessthen 10,5°C. After 1200-1800 s, the temperatures increased with similar rates in time in
the flow domain at different levels above the level of the working FAs’ bottom
perforations.
At 1800-2400 s, the sum by-pass flow become higher than the net flow through theFAs’ active parts and increased further. At 8000 s, it reached 70% in case of the
W11_F19_DIS0 calculation, 75% for the W11_F19_DIS1 calculation and 81% in caseof the W11_F19_DIS2 calculation.
The calculations show that the maximum temperature did not reach the saturation
temperature even after 8000 s. On the basis of trend line fitting it can be stated that theincreasing of the maximum temperatures had a rate of about 3.4 and 4.6 K/1000 s after
1200 s for the different model versions. With these rates the saturation temperature(∼120°C) would have reached after 22900, 17400 and 14700 s for the W11_F19_DIS0-(4)

DIS1-DIS2 calculations respectively.
On the other hand, it has to be emphasized that the initial temperature of the cleaningtank which was valid at 16:40, 10 April 2003 is unknown. The results’ characteristic andthe temperature values presented in Figure 3.5.1 refers to that independently from theinitial temperature, a similar stratified temperature field could be developed after awhile. It is likely that in case the initial temperature had been higher, the saturation
temperature would have been reached earlier.

From the above explained details, the following most important conclusions may be drawn:••

The calculations showed that even with or without any fuel assembly displacements thetemperature stratification could have been developed and blocked the flow through thefuel assemblies and the saturation temperature could have been reached after enough long
period of time.
On the other hand, by taking into account the fact that the maximum temperature reached
the saturation within 8400 s during the incident, the results point out that beside theworking fuel assemblies’ bottom perforations more by-pass area was present in the bottom
part of the cleaning tank.

More detailed description of the BME NTI analysis is available in Appendix B/IX.

38

3.6

Participant VEIKI

Calculations performed with the MELCOR 1.8.5. version on IBM PC using two different
models of MELCOR the PWR and BWR model.
3.6.1

MELCOR PWR and BWR models of the cleaning tankPWR Model
The parts included into the MELCOR modelling scope for the cleaning tank are given in
Figure 3.6.1.
The system was modelled with a total of 49 control volumes. The surge tank was connected tothe cleaning tank at the top of the upper plenum. Eleven fixed FA have been considered with
12 holes (each with d=9mm) at the bottom and 12 holes at the top. The 19 follower FA were
without holes. The geometry of fixed FA and followers were considered to be the same using
the fixed FA as base.
Heat loses from the inner vessel to the vacuum gap between the inner and outer vessel has
been modelled. Subsequently the outer vessel communicated the heat to the pool of shaft
No.1. Inner vessel wall assumed to be core boundary HS was calculated with radiation heattransfer taken into account through the vacuum gap between the two vessels. The flow-paths
connecting the control volumes are also summarised in Figure 3.6.1.
Degraded core calculations were performed by the COR package. The decay power has been
specified as a value valid 13 days after the shutdown. MELCOR calculated it as 258.2 kW.
The core (Figure 3.6.2) has been modelled with 6 radial rings (ring1 is empty representing thecentral hole) and 21 axial levels in the core. The fuel assemblies were divided into 5 groups
and placed into rings 2-6. (Ring1 was empty).
The COR 6 radial rings and 21 axial levels have been grouped into several control volumes asof Fig 3.6.1. The power distribution is given in Figure 3.6.3. Initial conditions were thefollowing:ooooPool temperature
30 °C
Tank coolant temperature 57 °C
Fuel assembly temperatures
Coolant+1-2 °C
Pump flow rate
22 t/hTABLE 3.6.1. DISTRIBUTION OF ASSEMBLIES IN RADIAL RINGS OF THE MODEL
Ring1

Ring2

Ring3

Ring4

Ring5

Ring6

CentralHole

Fixed FA 1-6

Fixed FA 7-11

Followers 1318

Followers 1924

Followers 2530

Follower 1239

Figure 3.6.1. MELCOR 1.8.5 tank PWR Figure 3.6.2. MELCOR 1.8.5 core BWR
nodalization.
nodalization.
BWR Model
The MELCOR BWR model – different from PWR model mainly by the presence of by-passCVs - can model the channel boxes of the WWER-440/213. It allows to simulate heat transfer
to space outside of the fuel assemblies modelled as by-pass region of a BWR core. In this way
it was possible to avoid the modelling of the oxidation of WWER-440 Zr channel boxes by adouble surface area.
The space outside of the fuel assemblies (by-pass) has been divided along the vertical in thesame way as the CVs along the core axial levels.
Opposite to the PWR model where the fuel-containing outer-most core ring (No6) wastransferring heat to internal vessel wall simulated as core wall in the BWR model spaceoutside of the fuel assemblies (by-pass) transferred heat the to wall of the inner vessel
modelled as core wall. (However core wall still receives radiative heat from the outermost
core ring.) These by-pass CVs also communicated with the space inside the fuel assemblies
by heat conduction via the assembly walls (channel boxes).
3.6.2

Safe envelope of operation

Simple hand calculations were done to identify the safe margins of inlet flow-rates atassembly bottom without and with steam generation. The results show that about 12% of
nominal flow-rate would have been enough to remove the decay heat without boiling and an
even smaller flow if the vent line could remove the steam generated to avoid level depression
and uncovery.

40

Figure 3.6.3. Radial and axial distribution of FA decay power in the Paks-2 cleaning tank.
Calculating the time needed to fill the upper plenum by steam and plotting it vs. steady stateflow for heat removal with steam generation one can obtain the time range for accident
development (Figure 3.6.4). Figure 3.6.4 shows a very steep change around 12% of nominal
flow. This figure also contains the flow-rates experienced with MELCOR before the strong
two-phase flow, which shows that some channels are generating steam and some are not bothin MELCOR and in simple hand calculations. The figure shows that uncovery by release of
dissolved gases is not a real danger.

Figure 3.6.4. Time needed to fill the upper plenum (vessel cap) by steam vs. relative flow-rateensuring steady-state heat removal with steam generation from groups of FA of Paks-2
cleaning tank. (Relative to nominal flow to group assuming 22 t/h for the whole vessel.).
3.6.3

Transient calculations

If vent line was suddenly open at moderate steam generation rate, most of the steam got
discharged from the tank and the FA were re-flooded. However if vent line was open all thetime then strong water level depression was experienced. The final calculations were donewith vent line closed.

41

Ring No. 3 was assumed to be miss-positioned (area=0,0120m2) and inlet flow area to ring 3(lowermost CV = FA unheated end tail containing the bottom holes) has been reduced by thesame amount.
The calculation was started from t= 0 s which corresponded to switching to low flow-ratepump – operation ‘AMDA C’. The timing of the main events is shown in Table 3.6.2. The
observed water temperature exiting the tank was about 37 C, which was reproduced by thecalculations well.
The results were sensitive to size a vacuum (0.01 bar in BWR model) between the two tanks
and surface emissivity (0.8 in BWR model) of the vessel walls. These influenced first of all
the radiative heat loss to water pool in shaft No1 at later stages. However the heat-up times
did not change too much because radiation starts to be important only at elevatedtemperatures. More detailed description of the VEIKI analysis is available in Appendix B/X.
TABLE 3.6.2. COMPARISON OF CALCULATED AND OBSERVED EVENTS USINGMELCOR PWR AND BWR MODELS
Vent line closed
PWR: Run9nA
BWR: Run H

Accident
phase

Mark

Time, sPWR model

Time, s
BWR model

Time, sObserved

Initial heat-up
Initiating event
Core starts to boil

0.0
Boil-off
Ring4
elev.
415

0.0

B

5600

7190

NA

8400

Core uncovery
Core exit temp. above 550 C

Core heat-up
Ring3-4
Ring4

U
5

9600
15921

9800
15706

Start of Zr-H2O reaction

Core damage
Ring4

Z

17804

17664

17977
22123
21030

17999
21711
23600
23141
19901

Gap release

Molten Zr (T>1825 C)

Ring4
Ring3
Ring5
Ring6

Not reached

18600-24600(1)(1) 18600-24600s corresponds to time interval of April 10, 16:40-23:30h

3.6.4

Conclusions

Based on several calculations with different MELCOR models (PWR and BWR) it can beconcluded that MELCOR reproduced the incident phenomenon well, and the experience canbe utilized in plant calculations. There is no big difference between the MELCOR PWR andBWR models results in term of accident progression and activity release. However the
MELCOR BWR physical model describes the heat transfer and oxidation phenomena of fuelassembly channel boxes special for this case better.
A coarse nodalisation in the core could not achieve conditions when steam generation started
for the present case with low heat generation.
Calculations were done with the vent line closed. Some studies were also done with vent lineon the tank open. Preliminary results suggest, that some core damage might have happened

42even in case of vent line open. However lack of precise data prevents us from drawing strong
conclusions on this case.
3.7

Participant KI/IRSN

The presented results of the Paks-2 cleaning tank incident simulation withICARE/CATHARE code are obtained in the frame of technical collaboration between RRC‘Kurchatov Institute’ (Russia) and IRSN (France), and should be regarded as a common RRC
KI and IRSN contribution to OECD-IAEA Paks fuel project.
This work was carried out in accordance with recommendations from OECD-IAEA Paks fuelproject meetings and input transient and modelling parameters are set following current
database of Paks incident. An analysis of the calculation results is done accounting therequirements for the participants of the project.
3.7.1

Code description

The ICARE/CATHARE code, developed in IRSN (France), is devoted to calculate in detailed
mechanistic way core degradation during severe accidents in LWRs. Both parts of the code,
ICARE2 and CATHARE, were developed separately as for analyses of wide range of hightemperature core phenomena as for realistic simulations of thermal hydraulics in different
reactor components. Recent wide validation of coupled ICARE/CATHARE code against
numerous integral experiments and SFD reactor scenarios showed its applicability and highprediction power, in particular, to WWER reactors and WWER type experimental facilities.
The first part of the current work included simulation of thermal hydraulic phase of theincident using CATHARE2 V1.3L_1 code. The second part concerned phenomena during
SFD phase, which appeared after dryout of the tank, heat-up and the high temperature
evolution of the fuel assemblies. The SFD phase was simulated using the updated version of
the ICARE/CATHARE code with ICARE2 V3.2 code version.
3.7.2

Thermal hydraulic phase of the transient

It was assumed that 3 of 30 fuel assemblies have additional flow by-pass at their entry. Thisby-pass is formed due to incorrect positioning of these assemblies on the support plate of thetank. Cross-flow area of the by-pass 120 cm2 was chosen in accordance with decisions of the
1st Meeting on OECD-IAEA Paks fuel project. Moreover ordinary by-pass due to perforationin the lower part of 11 fixed assemblies was taken into account. Geometry and hydraulicresistance values of the fuel assemblies were reproduced in the input data.
Water inlet mass flow 5.5 kg/s with temperature about 30°C was supplied with a pump to thetank from the cooling pond to cool the fuel assemblies.
The code predicted that major part of the water, supplied into the tank in the cooling mode,
was lost through the by-pass and the perforation. Only 10-20% of supplied water took part inthe cooling process. It led to insufficient cooling of the fuel assemblies and their gradual heatup. The temperatures reached the values leading to water boiling ~2 h after the moment offlow reduction at the beginning of cooling mode. Assembly dry-out began at ~2 h 20 min andcontinued ~20 min. Water level stabilization was predicted at elevation ~0.7 m, timing ofsteam void formation and final level position correlated with known plant data.
An uncertainty analysis was performed to estimate the variation range of physical parametersat the beginning of SFD phase. Four series of variant calculations were carried out, including43

by-pass flow area, gas relief pipe flow capacity and decay heat power. It was shown thatduration of the period from inlet flow reduction until the onset of the assemblies’ dry-out
varied in wide range (from 6% to 15%) due to these parameter uncertainties. Pressure loss
coefficients of the assemblies located in the tank were found to be less important parameters
for the incident simulation.
All known events and parameters of initial (thermal hydraulic) phase of the cleaning tankincident were successfully reproduced in CATHARE2 calculation. The results of thiscalculation formed initial and boundary conditions for subsequent ICARE2 calculations.
3.7.3

SFD phase of the transient

According to Paks database all 30 assembles inside cleaning tank are divided on 6 groups,
basing on burn up history and type of the assembly:
1.
2.
3.
4.
5.
6.

Fix assemblies 1-6 (burn up 10.9 MW day/kgU, power 7.159 KW);
Fix assemblies 7-11 (burn up 27.0 MW day/kgU, power 9.006 KW);
Follower
12 (burn up 9.2 MW day/kgU, power 8.154 KW);
Followers
13-18(burn up 21.3 MW day/kgU, power 6.694 KW);
Followers
19-24(burn up 13.9 MW day/kgU, power 8.837 KW);
Followers
25-30(burn up 13.7 MW day/kgU, power 8.719 KW).

Total decay power at time of incident was 241 KW.
Axial meshing in ICARE2 simulation included 20 nodes along heated region of the fuel rods.
A nodalization scheme with 3 parallel channels was used. Additionally, recently developed
best-fitted correlations for Zr+1%Nb alloy oxidation have been applied.
The first characteristic event of SFD phase appeared at 14360 second, when maximum
cladding temperature exceeds 800 ºC. The assembly of the 3rd group (single assembly № 12)
first reaches this temperature that totally corresponds to the peculiarities of the local decay
power, which according to Paks database is the highest for this assembly at the sameelevation. Later increase of radiation heat transfer smoothes this dependence and hottest zone
is shifted to central region of the cleaning tank. The assemblies of the 1st group (assemblies№ 1-6) first reach temperature of 1200 oC at time 19700 s.
According to simulation results maximum fuel rod temperatures does not exceed 1500 ºC,
which is realized approximately after 24000 seconds of the transient. Our analyses show thatgrowth of maximum temperatures is limited by oxygen starvation. The temperature between
fuel and cladding generally is rather small and differs by 5-10 ºC.
The evolution of internal fuel rod pressure, which is obtained during the simulation, indicatesthat first failure of fuel rod occurred at ≈ 16700 second in the central assembly. The latest
failure occurs about 500 seconds later on the periphery. It should be noted that these results
should be regarded only as a preliminary ones, as they were obtained with application of
mechanical behaviour model CREEP, which was developed for Zircaloy claddings.
The analyses of evolution of oxidation state of the assemblies show that total extent of Zr
oxidation is not very high increasing to approximately 15% at the end of the transient. At thesame time local oxidation of the fuel rod cladding appears to be substantial (about 80%) at
narrow region in central part of the cleaning tank.
Specific Zr oxidation behaviour of the regarded transient is realized in simulation of hydrogen
release. Following the computation the entire transient can be subdivided in three periods.
Firstly, when time is less than 20000 s, simulated temperatures are relatively low with no44

limitations on oxygen supply. Afterwards the hydrogen release rate is limited by oxygen
starvation and blanketing and remained constant. At the end of the transient the primary Zr in
major part of hot zone is converted into α-Zr(O) and total hydrogen release rate decreased.
Total hydrogen release reaches in this simulation the value about 11.5 kg .
Axial distribution of simulated external diameter of fuel rods at the instant before quenchingand zirconia layer thickness at the same instant correlates one with another. Therefore it canbe supposed that 25% of cladding radial increase during ballooning led to subsequent increasein Zr oxidation rate and is one of the reasons of heavy oxidation at these elevations. The
comparison of simulation results for different assembly groups shows certain dependence onradial position of the group. Namely, the central group (number 1 in the database list) issimulated with highest temperatures and oxide scale thickness, while these values decrease
from the center to the periphery. The latter effect is mainly the consequence of substantial roleof radiation transfer inside the cleaning tank.
A current analysis is based on recently developed by NSI RRC KI best-fitted correlations for
Zr+1%Nb oxidation, which were obtained from results of several experimental groups. The
difference between this kinetic and previously used VNIINM one (so-called Sokolov
correlations) falls mainly in the temperature regions T > 1500 oC and T < 1300 oC, where
best-fitted correlations predicts higher rate of oxygen mass gain. Here the results, obtainedwith both sets, are compared for Paks incident scenario. The comparison shows thatmaximum temperatures in variant simulation are lower and are reached at later instants thanin base case.
Generally, simulations of current studies confirmed behaviour of accident scenario withrespect to the uncertainties of the models and incident settings, which was outlined as a resultof previous investigations.
Further examinations are foreseen (outside of the OECD-IAEA Paks fuel project) with the
improved modelling of mechanical properties of Zr+1%Nb alloys. An update of CREEP
module of the ICARE2 code with mechanical properties of Zr1%Nb claddings is foreseen in
the nearest future with application to Paks incident at the second half of 2007 year. This work
will be performed in the frame of technical collaboration between RRC ‘Kurchatov Institute’and IRSN. More detailed description of the KI/IRSN analysis is available in Appendix B/XI.
3.8

Participant IVS

Analysis was performed using the integral code ASTEC (Accident Source Term Evaluation
Code), version V1.3 rev0.
3.8.1

Code description

The ASTEC V1 series has been developed jointly by IRSN and GRS since 1998 with the aim
to get a fast running code for the simulation of the total sequences of severe accidents in LWR
from the initiating event up to the possible fission product release to the environment. The
code version V1.3 rev0 was released in December 2006. The integral code ASTEC consists ofseveral modules. Only several of them were used in the analysis of Paks cleaning tankaccident:••

CESAR for RCS two-phase thermal hydraulics during the front-end phase and thedegradation phase,
DIVA for core degradation including late phase phenomena (molten pool, corium slump
to lower head, corium in lower head) and vessel failure,

45

•
•

ELSA for release of FP from fuel rods and debris and of materials from control rods,
using a semi-empirical approach,
SOPHAEROS for FP vapour and aerosol transport in RCS.

In typical reactor applications the ASTEC modules are working in coupled mode exchanging
the relevant data through common database. The front-end thermal hydraulic phase of theaccident is calculated by CESAR module. When the core heat-up takes place, the CESAR
thermal hydraulics in the downcomer, lower plenum and core is replaced by DIVA moduleand only rest of system is henceforth analyzed by CESAR module. Pre-defined basic
configuration of the core and lower part of reactor vessel, which represents arrangement of
typical PWR, is assumed in DIVA. Typically, the core region is split into arbitrary number ofradial rings and axial layers. The same axial power profile has to be considered in all rings.
Contrary to DIVA the CESAR nodalisation is very flexible and thus applicable to ‘arbitrary’thermal hydraulic systems. All above mentioned points imply certain limitations in modellingof Paks cleaning tank arrangement. The two most important are as follows:
1) Uniform axial power profile has to be considered in all fuel assemblies (at least duringcore heat-up and degradation phase),2) After start of DIVA module (i.e. after start of core heat-up and replacement of CESAR
core nodalisation by DIVA nodalisation), it was not possible to model direct by-pass fromthe lower holes of fuel assemblies to outlet nozzle from cleaning tank.
To cope with this, two successive calculations were performed:••

CESAR stand alone up to start of fuel heat-up (analysis focused on precise modelling of
the front-end thermal hydraulics of the cleaning tank),
All modules in coupled mode since the beginning of the fuel heat-up (450 °C); theoriginal fine CESAR nodalisation was replaced by simplified one and proper inlet/outlet
boundary conditions (analysis focused on fuel degradation).

3.8.2

Nodalisation scheme

The 30 fuel assemblies were split into 5 groups (6 assemblies in each group) based on theirlocation from the centre of cleaning tank:
1st (central) ring: assemblies No. 1 to 6,
2nd ring: assemblies No. 7 to 12,
3rd ring: assemblies No. 13 to 18,
4th ring: assemblies No. 19 to 25,
5th (outer) ring: assemblies No. 26 to 30.
The burnup of fuel assemblies within one group was relatively close each to other. The
exception was assembly No 12 in 2nd group. Average power within each group was
proportional to average burnup. This defines radial power distribution in 5 parallel corechannels (CESAR) and in 5 core rings, respectively. The total ‘core’ power was 241 kW.
TABLE 3.8.1. RADIAL POWER DISTRIBUTION (FROM CENTRE TO OUTER PART)ring No.
average power of one FA [W]

46

1

2

3

4

5

7159

9006

6694

8837

8719

Uniform axial distribution was used for both, CESAR and DIVA modules. This distribution
represents an average calculated from all fuel assemblies.
Only simplified approach was used in modelling of fission product release and transport.
Generic data were used for initial FP inventory instead of specific and the same inventory wasconsidered in all fuel assemblies. Individual isotopes were not modelled, only elements.
TABLE 3.8.2. AXIAL POWER DISTRIBUTION (FROM BOTTOM TO TOP)
axial mesh

1

2

3

4

5

6

7

8

9

10

relative power 6.372 9.126 1.001 1.151 1.200 1.199 1.159 1.081 9.575 7.028
10-2
10-2
10-1
10-1
10-1
10-1
10-1
10-1
10-2
10-2
3.8.2.1 CESAR nodalisation
Main features of CESAR nodalisation are as follows:-

Core (heated volume) represents only part inside fuel shrouds;
By-pass represents part outside shrouds,
5 parallel core channels modelled,
14 axial nodes used in total (2 lower, 2 upper and 10 heated nodes containing fuel),
Lower head with BC (representing inlet from cooling pump) and water inlet to fuelassemblies,
Upper head – outputs from fuel assemblies and by-pass; degassing line on the top wasmodelled too,
10 horizontal junctions (holes in the shrouds, improperly positioned FA was notconsidered),Heat transfer through shrouds modelled,
Outlet pipe and pressure BC.

3.8.2.2 DIVA nodalisation--

Basic components: Lower plenum, core (5 radial rings, 19 axial meshes – 15 of themheated) and by-pass,
Urbanic-Heidrick model was used for cladding oxidation,
DIVA is not able to model forced circulation cooling of lower part of tank (i.e. direct bypass through lower holes to outlet from tank; ‘once through’ model have to be usedinstead,Because of once-through model, the inlet BC was chosen to keep constant water level ~ 1
m from the bottom of cleaning tank (i.e. applied inlet flow rate was chosen just tocompensate water boil off from flooded part of FAs, real inlet flow rate was higher due to
direct flow by-pass through the lower holes of FAs to outlet nozzle from cleaning tank.
Heat losses from the tank walls were artificially increased to compensate this effect.

3.8.3

Main results

Overall system behaviour was reasonably well predicted by ASTEC code. In total 4.68 kg ofhydrogen was produced during the analysis. Maximum cladding oxidation in central part ofupper, uncovered core region reached 100%. Burst of cladding occurred in all rings except ofoutermost one. Maximum cladding temperature during the accident was lower than 1400 °C.
More detailed description of the IVS analysis is available in Appendix B/XII.

47

TABLE 3.8.3. CHRONOLOGY OF MAIN EVENTS IN ANALYSIS
EventTime [s]Transition from cleaning to cooling regime (16:40 h. real time)0

Local saturation in the fuel assembly~ 9 500

Increase of PRZ water level 7 cm

10 700 – 12 100

Max. cladding temperature 450 °C (start of DIVA module)

14 014Max. cladding temperature 800 °C

17 872

Failure of the 1st fuel rod due to ballooning

19 311

Start of intense Zr oxidation (production > 2 g/s)

22 000

Failure of the last fuel rod before quenching*

22 019Max. cladding temperature 1200 °C

24 930End of analysis

34 410* fuel assemblies in outer ring remained intact.

3.9
3.9.1

Participant VUJE
Description of the input model

The RELAP5/Mod3.2.2 code was used for thermal hydraulic analyses of Paks incident. The
code has been developed for the analyses of light water reactor coolant systems duringtransients and postulated accidents. More detailed code description is listed in Appendix B.
Calculation model of cleaning tank with 30 pieces of fuel assemblies was developed.
Nodalisation of the model is shown in Fig. 3.9.1 [19].
The model is composed from interconnected simple volumes. Coolant inflow into tank isrealized by connection no.30. Type of this connection is ‘TMDPJUN’. Volume under fuelassemblies is represented by one ‘BRANCH’ type volume (no.31).
Fuel assemblies are lumped into 6 groups. Each group of fuel assemblies is modelled by‘PIPE’ type component (volumes 1 - 6). These volumes are split in axial direction onto 13
sections. Fuel part of assemblies is divided onto 10 axial sections. These fuel sections are
represented by heat structures 1001 - 1006. In each group of fuel assemblies, decay heat andaxial profile of decay heat is defined according to data specified in the database of the project.
Volume of cleaning tank above fuel assemblies is modelled by single volume no.34.
Component no.35, connected into upper part of volume no. 34, represents tank relief valve.
This valve opens when tank pressure exceeds 0.25 MPa.
Coolant flows from tank through volume no. 38 (‘PIPE’ type) into ‘TMDPVOL’ component
no.40, where the constant pressure is set (boundary condition). Tank heat losses are modelledby heat structures no. 1020 and 1021.
More detailed description of the input model is presented in the Appendix B.

48

36

40

35

39

4

3

2

1005

1004

1003

1002

1021

38

1

7

42

1020

5

1001

6

1006

34

41
37

31

29
30

Figure 3.9.1. Nodalisation of cleaning tank.
3.9.2

Initial and boundary conditions

Initial and boundary conditions for analysed cases were set in order to reach as good
agreement with the real conditions during incident as possible.
In this chapter the reference case is showed. Initial and boundary conditions for this case wereset as follows:
Parameter

Value

By-pass flow area

120 cm2

Decay power of assemblies No. 1 - 6

42.952 kW

Decay power of assemblies No. 7 - 11

45.028 kW

Decay power of assembly No. 12

8.154 kW

Decay power of assemblies No. 13 - 18

40.164 kW

Decay power of assemblies No. 19 - 24

53.021 kW

Decay power of assemblies No. 25 - 30

52.312 kW

Total decay power

241.7 kW

Initial coolant temperature in container

56 °C

Initial container pressure

0.23 MPa

External pool temperature

30 °C

Temperature of inlet coolant

30 °C

Inlet mass flow

21 t/h49

Axial profile of decay heat distribution for each assembly group was defined according todata provided in the database of the project. Graphical form of axial profile of decay heat
distribution is presented in Appendix B/XIII.
3.9.3

Description of results

Process analysis begins at the moment of start of the intermediate cooling with coolant mass
flow through tank of 21 t/h. Due to the lower flow rate of the intermediate cooling pump, theby-pass flow through the perforations in the assembly shroud and at the bottom of incorrectlyseated assemblies became much more significant, than it was during the cleaning operation
with high flow rate. Up to 90% of water added into cleaning tank flows through these
perforations and through the by-pass. Cooling of fuel assemblies is not sufficient and heating
up of the water leads to saturation state at the top of the cleaning tank at the 4500 s.

TABLE 3.9.1 CHRONOLOGICAL SEQUENCE OF IMPORTANT EVENTS
EventStart of cooling mode (mass flow 21 t/h)Saturation temperature in the cleaning tank
Start of water level drop in the cleaning tankTime [s]
0
4500
7260

Air letdown valve opening

9440

Maximum cladding temperature exceeds 800°C
Start of intensive Zr oxidation
Maximum cladding temperature exceeds 1200°C

13900
16400
20600

Maximum cladding temperature (1470 °C in assembly no.12)

29700

End of calculation

36000

Water level in tank begins to decrease after 2 hours since start of intermediate cooling regime.
Water level is stabilised at level 1.8 m from lower plate. This decreased water level leads tothe temperature increase of the fuel assemblies. Due to its highest and peaked axial decay heat
profile, maximum cladding temperature 1470 °C is found in assembly 12 at time 29700 s.
The air letdown valve opens at 9440 s, when tank pressure exceeds 0.25 MPa. Flow rate
through this valve is approximately 5 g/s.
Chronological sequence of important events is presented in TABLE 3.9.1. The results ingraphical form are presented in Appendix B/XIII.
3.10 Participant USNRC
This summary report describes the analysis of the Paks event using the MELCOR severe
accident analysis code. The Paks event was a fuel degradation accident in a cleaning tank inthe spent fuel pool of the reactor at the Hungarian Paks Nuclear Power Station. In the event,30 assemblies in the cleaning tank overheated due to inadequate heat removal.

50

3.10.1

Code description

A MELCOR model was developed to represent the inner cleaning tank, the outer tank, the30 assemblies within the tank, and the surrounding pool. The specified boundary conditions
for the model included the tank inlet flow rate, the assembly decay heat, the hydrostaticpressure at the tank exit and degassing line, and the operation of the hydraulic lid locking
system of cleaning tank. It was assumed that a 0.01 m2 leakage hole opened at 9.5 hours when
the tank lid was unlocked. Some of the inlet flow leaked around misaligned assemblies at thebase plate of the inner tank (i.e. FL-101) and by-passed the fuel assemblies. The leakage area
was specified to be 120 cm2 as agreed at the second workshop. The leakage was due toincorrect seating of the assemblies into the lower support plate. The perforations at the top
and bottom of the shroud of the standard assemblies were also modelled as leakage paths.
At the top of the inner tank, there was a 10 mm OD degassing line. In normal operation, thedegassing line is full of water. It was assumed that water within the degassing line precluded
vapour flow until the inner tank steam bubble was established and the tank was starting topressurize. From the Paks database, the degassing line resistance corresponded to a length of
15 m and its elevation change was 13.1 m. The MELCOR model was subdivided into 5 rings
to represent the 30 assemblies.
The MELCOR CORSOR-Booth fission product release model was used and the volatile
specie release coefficients were modified to match measurements from recent French
experiments. The vapour pressure, compound form, and initial inventory default inputs weremodified to represent Cs2MoO4 and CsI as the dominant compounds for the cesium andiodine releases. The initial fission product inventory in the gap between the fuel and cladding
was estimated to be <0.1%, based on the calculated released fission product gas.
3.10.2 Thermal hydraulic response
The accident simulation was started when the system was configured to a single submersible
pump (i.e. the conditions at 4:40 pm on April 10, 2003) and was terminated 0.5 hours after thetank lid was unlocked, which flooded the tank (i.e. 10 hours). The inlet flow to the cleaningtank was initially 5.69 kg/s and 30oC. From the previous operation, the water temperature in
the tank was 57oC (330 K).
Following the shift to the low-flow submersible pump and the storage pool as the water
source, relatively cold water (30oC) was introduced into the initially warm tank (57oC). 65%(i.e. the initial maximum) of the cold water flowed upward into the assemblies and wasapproximately evenly distributed across the 30 assemblies. The calculation assumed that someof the water leaked through the gap between baseplate holes and the assemblies (initially~25% of the tank inlet flow) and some leaked through the perforations on the standard
assembly inlet nozzles (initially ~10% of the tank inlet flow). The cold water cooled thebottom of the assemblies and increased the hydrostatic pressure drop through those pathways
relative to the leakage pathways. Hence, the calculated leakage flow across the baseplate andthrough the standard assembly perforations steadily increased while the assembly flow
decreased. As the accident progressed, the calculated assembly flow dropped to 12% of theinlet flow by 1.7 hours while the leakage flow increased to 88% (52% across the baseplateand 36% through the perforations in the standard assemblies).
As a result of the flow reduction into the assemblies, the calculated fluid temperature at theupper portion of the assemblies steadily increased towards saturation conditions. At 1.97
hours into the simulation, the water at the exit of Assembly 12 started to boil. By 2.44 hours,

51all the assemblies were boiling. As steam exited the assemblies, it created a steam bubble,
which filled the top of the tank and depressed the water level in the tank and in theassemblies. One of the few measurements available during the accident was the primarycircuit pressurizer level, which was hydraulically connected to the storage pool with thecleaning tank. At 2.6 hours into the accident (7:20 pm on April 10, 2003), the pressurizer tankshowed a 70 mm water level increase (i.e. an undiagnosed indication of water displacementfrom the cleaning tank). The timing of the calculated drop in the cleaning tank water level
closely corresponded to the measured increase in the pressurizer tank level.
The calculated level response shows an initial level depression to 1.2 m followed by arecovery to 1.4 m (i.e. measured from the bottom plate). Although the rate of steam
production exceeded the degassing flow, condensation of steam against the relatively cool
inner tank wall reversed the level trend. After 3.2 hours, non-condensable hydrogen from thezirconium-steam oxidation reaction started to replace steam in the upper tank. The water levelremained at 1.4 m until 9.5 hours when the hydraulic latch on the lid was opened.
At a water level of 1.4 m, the top 60% of the fuel was uncovered. The steam flow due toboiling below the 1.4 m water level was inadequate to cool the fuel. The top portions of all theassemblies heated from 2.5 hours to 9.5 hours. Due to thermal radiative heat transfer, there
was a temperature gradient from the assemblies closest to the inner tank wall (#19 -#30)
versus the assemblies at the center of the tank (#1 - #6). The peak temperatures of theassemblies in the middle ring (#7 through #18) lied between the inner and outer rings of fuelassemblies.
The peak cladding temperature was calculated to be 1743 K, which occurred in Ring 1(Assemblies 1 through 6) just prior to the tank reflood. The corresponding peak temperatures
calculated for Rings 2 through 5 were 1540 K, 1672 K, 1667 K, and 1403 K, respectively.
The temperature response of the fuel was a complex function of the magnitude of the decay
heat and oxidation powers, the axial power profile, the magnitude of steam generation below
the water level in each assembly to drive oxidation, the radiative exchange to surrounding
assemblies or the tank wall, and the convective heat removal rate. The radiative heat losses tothe tank wall had a significant impact of the peak temperatures. For example, the heat losses
to the tank wall exceeded the decay power of the assemblies in Ring 5 after 6.1 h. However,the fuel temperatures in Ring 5 continued to increase due to radiative heating from theassemblies in the inner rings. Because the assemblies in Ring 1 were the furthest from thetank wall, they reached the highest temperature. In the middle ring of assemblies, Rings 3 and4 had higher decay heat power in the region above the water level and therefore reached ahigher temperature than the assemblies in Ring 2. Although Ring 3 (Assembly 12) had thehighest decay power above the water level and initially heated the fastest, radiative exchange
to the surrounding assemblies and low steam production below the water level (i.e. forcontinued oxidation) limited its peak temperature to below Ring 1.
3.10.3 Fission product release
The first failure of the fuel cladding and release of the gap fission products was calculated tooccur at 4.0 h in Assembly 12. The other rods subsequently failed from 4.2 h to 4.9 h. The
thermally driven fission product releases continued following the rod failures in each group of
assemblies. After about 6 h, the release increased more rapidly. Approximately half of thevolatile xenon, cesium, and iodine releases occurred in the last 1.5 hours (i.e. between 8 h to
9.5 h).

52

The calculated releases were 1.1% of the noble gases, 1.1% of the cesium, 0.8% of the iodine,
0.02% of the barium, and 0.03% of the cerium and lanthanum. The data from themeasurements of 8 representative nuclides were integrated for the following time periods after
the accident: 1 day, 3 days, 7 days, 14 days, and 2 years. In general, the comparisons of the
MELCOR results to the 14 day integrated tank measurements were relatively good. The
temperature profile from the inner assemblies to the outer assemblies meant the calculated
fission product releases were higher in Ring 1 versus Ring 5. The overall calculated andmeasured results show that approximately 1% of the volatile inventory (i.e. the noble gases,
cesium, and iodine) was released during the high temperature portion of the accident, which isin good agreement with the 14 day data. The measured time-evolving release of somenuclides raises some uncertainty in the comparison of MELCOR to some nuclides (i.e. what is
attributed to the delayed transport from the tank from the high temperature accident versus
continued low-temperature dissolution release) . The comparison of the calculated and 14 day
barium release data suggests higher releases than calculated. The cerium and lanthanum
release models had good agreement with the 14 day integrated data. More detailed descriptionof the USNRC analysis is available in Appendix B/XIV.
TABLE 3.10.1. CHRONOLOGY OF MAIN EVENTS
Calculated

A.

Event

TimingMeasured or
Estimated Timing

Assembly 12 reaches saturation near theassembly outlet

1.72 hours

n/a

Water level starts to drop in the cleaningtank

2.36 hours

Water level stops decreasing

2.67 hours

Peak cladding temperature exceeds 800oC

3.69 hours

Hydrogen production >2.5 g/s

4.8 hours

Zr oxidation shows signs of steam limiting>4.8 hours

First fuel rod failures (Assembly 12)

4.00 hours

Second fuel rod failures (Assemblies 13-18)

4.17 hours

Third rod failures (Assemblies 1-6)

4.31 hours

Fourth rod failures (Assemblies 7-11)

4.39 hours

Last fuel rod failures (Assemblies 19-30)

4.94 hours

Gap + thermal noble gas release is >0.1%

4.94 hours

Hydrogen production >3 g/s

5.10 hours

Peak cladding temperature exceeds 1200oC

5.23 hours

Start of quench

9.50 hours

2.33 - 2.66 hours

n/a

5.22 hours (9:53
pm)(First rad. alarm)
Note An/a
9.50 hours

The gap releases in the calculation were assumed to be small (<0.1%) based on fuel behaviour calculations in the
Paks database. The calculated releases did not exceed 0.1% of he initial inventory until 4.94 hours, which is inbetter agreement with the timing of the first alarm measurement than the first gap release.

53

4

COMPARISON OF CALCULATED RESULTS

The results of the calculations were provided by 11 organisations from eight countries for thisfinal report. Some organisations carried out all three types of calculations (thermal hydraulics,
fuel behaviour and activity release), while other organisations focused their work on one or
two selected calculation type.
The name of analysts, the codes and the type of performed calculations are summarised inTable 4.1.
TABLE 4.1. SUMMARY OF PERFORMED CALCULATIONS
Calculations
ThermalFuelhydraulics behaviour

Country
Organization

AnalystCode

Belgium
TRACTEBEL

Jarne Verpoorten

MELCOR 1.8.5+

Pasi Junninen

APROS

+

Kari Pietarinen

FRAPCON-3
FRAPTRAN

Klaus Trambauer

ATHLET-CD+

László Perneczky

RELAP 5 mod 3.3+

István Trosztel

ATHLET+

Attila Molnár

FRAP-T6

Emese Szabó

EU LOCA model

Gábor Légrádi

CFX+

Gábor L. Horváth

MELCOR 1.8.5++

Yuri Zvonarev

ICARE2 V3.2 /CATHARE2
V1.3L_1++

Peter Matejovic

ASTEC

+

+

Martin Fogel

RELAP 5+

Kenneth Charles
WagnerMELCOR+

Finland
VTT
Germany
GRS

Hungary
AEKI

Hungary
BME NTI
Hungary
VEIKI
RussianFederation/
FranceKI/IRSN
Slovakia
IVS
Slovakia
VUJE
United Statesof America
US NRC

54+

Activity
release
+

+
+

+

+
+

+

+

+

The calculated time of the main events as provided by the participants are summarised inTables 4.2 and 4.3. Participant BME NTI is not listed in the tables, for their calculation
covered the initial single-phase period of the event and the analysis did not reach saturation in
the cleaning tank that was the first event requested to report. Plots of selected parameters areshown in chapter 4.4.
TABLE 4.2. CALCULATED TIME OF MAIN EVENTS AND SOME INTEGRALPARAMETERS BY PARTICIPANTS SUEZ - TRACTEBEL ENGINEERING, VTT, GRS
AND AEKI

saturation in the cleaning tank (s)water level drop in the cleaning tank(s)maximum
cladding
temperature
exceeds 800°C (s)maximum
cladding
temperature
exceeds 1200°C (s)start of intense Zr oxidation (s)failure of the first fuel rod due toballooning (s)failure of the last fuel rod before
quenching (s)maximum cladding temperature (ºC)
total Zr oxidation (%)
final hydrogen production (kg)

TRACTEBEL
8520

VTT

GRS

AEKI.1

AEKI.2

7260

7060

2830

5400

8520

8400

8640

10560

8000

16250

13800

14690

15800

20110

18780

21240

17720

16610

14820

15840

17000

16440

16036

18850

18300

19842

20010

1628
46.7*
15.1

1381
4.6
4.63

1152
4.1
3

1400
12
13

AEKI.1 and AEKI.2 correspond to thermal hydraulic calculations by RELAP and ATHLET codes respectively.
*
Only the zirconium of rod cladding was considered in VTT calculations.

TABLE 4.3. CALCULATED TIME OF MAIN EVENTS AND SOME INTEGRALPARAMETERS BY PARTICIPANTS VEIKI, KI/IRSN, IVS, VUJE, USNRC
VEIKI
saturation in the cleaning tank (s)water level drop in the cleaning tank(s)maximum
cladding
temperature
exceeds 800°C (s)maximum
cladding
temperature
exceeds 1200°C (s)start of intense Zr oxidation (s)failure of the first fuel rod due toballooning (s)failure of the last fuel rod before
quenching (s)maximum cladding temperature (ºC)
total Zr oxidation (%)
final hydrogen production (kg)

IVS

VUJE

USNRC

5600
9600

KI/
IRSN
7200
9600

9500
12100

4500
7260

6192
9612

17500

14360

17872

13900

13284

18688

19700

24930

20600

18828

17804

18000

22000

16400

17280

17977

16700

19311

-

14400

21030

17200

22019

-

17784

1895
-

1471
12.6
11.5

1350
10.9
4.68

1440
-

1470
11.0
7.79

55

4.1

Thermal hydraulics

In the first part of the incident – until a fairly stable steam volume in the upper part of thecleaning tank is formed – the course of the incident is determined by thermal hydrauliceffects. The main thermal hydraulic phenomena in this period are the following:
•
•
•
•
•

Distribution of the tank inlet flow rate among the fuel assemblies, shroud perforations
and by-pass due to incorrectly seated assemblies,
Change in the above flow distributions with heat-up of the coolant inside theassemblies and in the down-flow part of the cleaning tank,
Steam formation in the upper sections of the assemblies and its accumulation in thetank upper head,
Flow stagnation in the fuel assemblies and decrease of the tank level,
Stabilization of the tank level.

A large variety of codes has been used by the 11 participants. MELCOR and RELAP5 werethe most widely used ones: with 3 and 2 participants, respectively. One user evaluated thethermal hydraulic part with the ATHLET and APROS code respectively, and the CFX CFD
code was also applied for the first part of the incident. All other users applied tools withsevere accident capabilities: ATHLET-CD, ICARE/CATHARE, ASTEC.
The modelling approach applied by the participants differed in the number of groups the 30
fuel assemblies were distributed to. The most widely used grouping consisted of 5 or 6
groups, based on type (working or follower) and on similar power of the assemblies, but 4, 3
and 1 groups were also applied by GRS, Suez - Tractebel Engineering and AEKI in its
RELAP5 analysis, respectively. The draw-back – especially of the two last calculations – isthat the difference in the inlet flow rate of working and follower assemblies cannot be
captured, but the overall behaviour in the first phase can be fairly well predicted even by these
simple models, since the flow rates in the heated part of the different assemblies are very
similar.
Most of the participants divided the heated part of the fuel channels to about 20 axial nodes,
in an earlier parametric study performed by the ATHLET code it was demonstrated that this issufficient for convergence. In some cases only 10 axial nodes were applied that might impact
the accuracy of the results.
In order to assess the thermal hydraulic phenomena listed above Figs. 4.1 to 4.14 are ofinterest from the comparison plots.
The coolant outlet flow rate from the tank is shown in Fig. 4.1. As it can be expected in thefirst period, until steam formation, it corresponds to the inlet flow rate that was used as aboundary condition in the analyses. The flow rate increase due to substantial steam formation
around 8000-10000 s is shown by all calculations.
The tank outlet temperature (Fig. 4.2) varies in the early phase between 35 and 40 oC – thevalue depends on the amount of by-pass flow calculated by the different participants. The
only exception is one calculation, which applied an inlet temperature of 60°C. Again, theeffect of steam production is visible in each of the results, although the resulting temperaturepeak differs strongly.
The initial system pressure (Fig. 4.3) shows large variation, which is due to the different
approach chosen in the calculations to impose this boundary condition. All analyses show56pressure increase, when steam is produced in the tank that must be related to the increased
pressure drop between the highest point of the tank outlet tube (where most calculations useda pressure boundary condition) and the tank upper plenum due to the increased flow rate in
two-phase conditions.
Figures 4.5 and 4.6 display the outlet temperatures of assemblies 1-6 and 12, respectively, i.e.
those of working and follower assemblies. The figures show considerable differences among
the participants during the heat-up phase to saturation, which can be explained mainly by thefact that the flow rates along the heated parts are rather different. (To some extent differencesin assembly group powers and fuel heat capacities may be the cause as well.)Figure 4.11 shows that there is a factor 3 difference among the calculated inlet flow rates tothe follower assembly, but all the calculations predict correctly the decreasing trend of theflow rate during the subcooled period, which is the result of the continuously changing
elevation head balance between the regions inside and outside the assemblies. Similardifferences can be seen at the inlet of the working assemblies, as shown by Figs. 4.8 and 4.9,
although neglecting the two extreme results, the flow rates are closer. Most of the flow rate
entering the working assemblies exits via the shroud perforations at the bottom, as shown by
Figs. 4.13 and 4.14. Here again, the tendency of all the calculations is correct, indicating an
increased loss of the coolant via the perforations with time. An important amount of the inlet
flow to the cleaning tank passes through the assumed by-pass due to incorrect seating, as
evidenced by Fig. 4.12. The trend here is similar to that of the perforation, i.e. the flow rateby-passing the assemblies is increasing with time. It is interesting to note that there is a sort ofcompensation between these two by-pass flows: when in a given calculation the perforation
flow rate is low, the baseplate by-pass is high.
For the differences in the calculated flow rates the following reasons can be suspected,
different by-pass areas and loss-coefficients used by the participants, but also the models offriction losses in the laminar-turbulent transition region may be responsible. In spite of therelatively large differences in different flow rates it can be said that – in view of the very
delicate balance between elevation head losses and laminar flow friction losses – the range ofthe results is acceptable.
Tables 4.2 and 4.3 summarize the times to saturation in the different calculations: thedispersion is rather important, ranging from 2830 to 9500 s, even if most of the cases put it
around 7000 s. It should be mentioned that most flow rates, but especially those through theheated part (Fig. 4.9) show strong oscillation when saturation conditions are reached at thefuel assembly outlet.
The calculated levels of the cleaning tank are compared in Fig. 4.4 and in Tables 4.2 and 4.3.
These indicate that the analyses predict the level decrease due to steam accumulation in theupper head with an uncertainty range of 5000 s, although most of the calculations predict it
between 8000 and 10000 s.
4.2

Fuel behaviour

Fuel behaviour calculations have been carried out with the MELCOR, ICARE2, ATHLETCD, ASTEC, FRAPTRAN and FRAP-T6 codes. The initial state of the fuel rods at the
beginning of the incident was defined using the results of TRANSURANUS and FRAPCON3 analyses.

57After the formation of steam volume and low water level in the cleaning tank the fuel rod
temperatures started rapid increase. The initial heat-up rate in the dry phase was similar in
most of the calculations (about 0,1 ºC/s). In 2-3 hours after the beginning of dry phase themaximum fuel temperatures in most of the calculations reached 1000 ºC. The further increaseof temperature was much slower (Figs. 4.19, 4.20 and 4.21). There were two typicaltemperature histories for the high temperature phase (above 1000 ºC) of the event:•Some calculations produced a temperature plateau and the maximum temperature
seemed to converge to a given value. Obviously in these analyses some kind of
equilibrium was reached between the internal heat sources (decay heat and heatzirconium oxidation) and heat removal from the cleaning tank (losses to thesurrounding water as shown in Fig. 4.16, coolant circulation by the pump, gas andsteam release through the air letdown valve).

•

Some other calculations indicated peak in the temperature histories and showeddecreasing values in the last phase of the incident. The peak values can be associatedwith the period of very intense zirconium oxidation.

The calculated cladding temperatures were close to the fuel temperatures, for the thermal
power in the rods two weeks after reactor shutdown was very low: the average linear power
was 0.26 W/cm. The maximum cladding temperatures reached 800 ºC between 14000-18000
s and 1200 ºC between 19000-25000 s. The typical value of maximum cladding temperature
was between 1200-1400 ºC, but some calculations showed even higher temperatures close to
the melting point of zirconium (Figs. 4.22, 4.23 and 4.24).
The maximum cladding temperatures were received not for the same fuel assemblies bydifferent participants. Basically two cases can be identified for the location of maximum
temperature: the fuel assemblies with maximum power, or fuel assemblies in the centre of thecleaning tank. It can be explained by the following considerations:•

The power of fuel assemblies was not the same but varied between 7 and 9 kW.
Furthermore there were significant differences between the axial power distributions,since the normal working assemblies had cosine shaped profile and the follower
assemblies had local peak in the upper section of the fuel rods. The nodalisation
schemes developed by the participants included different groups of assemblies withdifferent powers and power profiles.

•

Some numerical models represented the fuel assemblies in form of parallel channels
without any connections between them. Some other calculations specified thermal
connections between the channels or used a ring structure for the simulation of theposition of fuel rods.

The calculated axial profile of temperature distribution in the fuel rods was also influenced by
the above factors. In most of the calculations the maximum temperature was reached in theupper-middle part of the rods. The bottom part of the rods had cladding temperature close to
the temperature of cooling water. The shape of the temperature profile usually was similar tothe power profile applied in the modelling of the given group of fuels rods. Comparing thetemperature profiles at the moment when 800 ºC cladding temperature was reached large
differences (200-300 ºC) can be found e.g. for the central fuel assemblies that can be
explained by the applied radial profiles (Figs. 4.17 and 4.18).

58

The fuel rod internal pressure showed increasing trend until the burst of fuel rods (Figs. 4.24).
Some calculations captured well the short period when the intense plastic deformation
resulted in pressure increase before the cladding integrity was lost. The failure of the first fuelrod was indicated by the participants between 14000 and 19000 s. It is in very good
agreement with the observation of first noble gas activity release that was detected at 18600 s(21.50 real time) during the Paks-2 incident. The calculations confirmed that the first activity
release from the cleaning tank was the result of plastic deformation and burst of fuel rods.
The fuel rod internal pressure at the moment of burst was typically between 12 and 20 bars.
The simulation of cladding plastic deformation was carried out with Zircaloy-4 models bysome participants and with WWER specific (E110 cladding) correlations by otherparticipants. However the most significant effect on the calculated failure pressure had theinitial pressure of the fuel rods (Figs. 4.24).
The ballooning and burst of fuel rods continued after the failure of the first rod. Most of thecalculations indicated the failure of all fuel rods in the 30 assembly. The last rods failedusually within one hour after the burst of first rod (Tables 4.2 and 4.3). According to theresults of some calculations several fuel rods remained intact until the opening of the cleaningtank. The calculated fuel rod external diameters (Fig. 4.29) indicated that the neighbour rods
could touch each other and it could significantly reduce the flow area in the assembly. The
deformation in some cases was overestimated, since the single rod models were not able totake into account the effect of neighbour rods on the maximum deformation.
The intense zirconium oxidation started at about 15000 s in most of the calculations (Tables
4.2 and 4.3). The simulation of zirconium oxidation in several calculations covered not only
the fuel rod cladding, but the shroud of the assemblies as well. The maximum cladding
oxidation at the end of the incident reached 100% in one calculation and varied between 1080% in the other analyses (Fig. 4.25). The calculated shroud oxidation was between 20-90%(Fig. 4.26). The profile of oxide layer thickness showed maximum in the upper-middle part of
the rods, the typical value was between 100-600 µm (Figs. 4.30 and 4.31). The typical valueof total oxidation of zirconium components at the end of calculation was between 4-12% (Fig.
4.27). Only one calculation showed considerably higher (47%) oxidation value. The
corresponding mass of produced hydrogen was between 3 and 15 kg (Tables 4.2 and 4.3). The
average value of hydrogen production was less than one g/s (Fig. 4.28). The large scatter inthe calculated degree of zirconium oxidation can be explained by several factors:•

The temperature histories (Figs. 4.21, 4.22 and 4.23) and temperature profiles (Figs.
4.17 and 4.18) had obviously very important effect on the oxidation process. However
it must be mentioned that the maximum temperatures and maximum degree ofoxidation were produced in different calculations.

•

The oxidation process was influenced by the availability of steam. Since the release of
produced hydrogen was limited through air letdown valve (Fig. 4.15), theaccumulation of hydrogen in the cleaning tank could reduce the access of steam tozirconium surface and suppress the oxidation process.

•

The calculation of oxidation kinetics was carried out with different correlations.
Models with conservative correlations (e.g. Baker-Just) produced much moreoxidation than some other models with realistic WWER specific models (e.g. AEKI
correlation).

•

In some calculations only the oxidation of cladding was taken into account and theoxidation of Zr shroud was not considered.

59

The quenching of hot fuel rods was simulated only by one calculation. Quenching did not leadto temperature or hydrogen production peaks in the calculations (Figs. 4.21, 4.22 and 4.23),
because the zirconium surfaces were heavily oxidised in the dry period of the incident andbecause the temperature of cladding and shrouds were too low for melt formation andsubsequent strong escalation.
The opening of the cleaning tank at the end of the incident was simulated only from thermal
point of view. The applied codes did not have detailed models for the simulation of thefragmentation of fuel assemblies due to thermal and mechanical stresses. The hydrogen
uptake of zirconium probably played very important role in the embrittlement of fuelassemblies, but the hydrogen uptake and its effect on the mechanical strength of cladding can
not be simulated by the currently available fuel behaviour codes.
4.3

Activity release

Activity release simulations have been carried out with the MELCOR and ATHLET-CD
codes and one series of hand calculations was performed using the EU model for calculating
the releases and consequences of a large LOCA model.
Most of the calculations produced data not for isotopes, but for elements or groups ofelements. The calculated release rates in those cases were multiplied by the isotope inventoryof 30 assemblies to receive the requested activity release values.
The beginning of activity release from the fuel rods corresponded to the first fuel failures due
to ballooning and burst. The good capture of timing of fuel failure was already discussed inchapter 4.2. The further activity release was related to the increase of number of failed fueland to the increase of fuel temperatures.
The main observation for some important elements can be summarised as follows:•

Noble gas release calculations were compared to activity measurements in thechimney of the reactor hall. The transport time between the cleaning tank and thechimney can be estimated as less than one hour considering the ventilation system of
the reactor hall. Some calculations were close to the measured values, but there were
calculations showing much higher release from the fuel, too. The estimated relative
release from measured data was 1.19% for 133Xe (Fig. 4.32).

•

Iodine release was generally slightly underestimated by the calculations. However thevalues were close to the measured value that was 1.41% for 131I isotope. The measured
value was determined from coolant activity measurements considering the first two
weeks after the incident. Two calculations produced data with about one order aboveand below the measured relative release (Fig. 4.33).

•

Cesium release was slightly overestimated by all calculations (Fig. 4.34). The
measured values were 0.74% and 0. 53% for 134Cs and 137Cs isotopes, respectively.

•

Tellurium (132Te, Fig. 4.35) and cerium (141Ce, Fig. 4.36) releases were well
reproduced by some of the participants. The release of actinides (242Cm) was
overestimated by all calculations (Fig. 4.37).

The scope of activity release calculations was limited to the release from fuel and did notinclude the transport of radioactive isotopes in the spent fuel storage pool or in the reactor
hall.

60

Plots of selected parameters

Cleaning tank coolant outlet flowrate [kg/s]

10

TRACTEBELVTTGRSAEKI.1
AEKI.2
BME-NTI
VEIKI
KI/IRSN
IVS
VUJE
USNRC

9
8
7
6
5
4
3
2
1
0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.1. Cleaning tank coolant outlet flow rate.
Cleaning tank coolant outlet temperature [°C]80
TRACTEBELVTTGRSAEKI.1
AEKI.2
BME-NTI
VEIKI
KI/IRSN
IVS
VUJE
USNRC

70

60

50

40

30

20
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.2. Cleaning tank coolant outlet temperature.
Pressure in the upper volume of the cleaning tank [bar]

4.4

3.0

TRACTEBELVTTGRSAEKI.1
AEKI.2VEIKI
KI/IRSN
IVS
VUJE
USNRC

2.9
2.8
2.7
2.6
2.5
2.4
2.3
2.2
2.1
2.0
1.9
1.8
1.7
1.6
1.5
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.3. Pressure in the upper volume of the cleaning tank.

61

1345013400

Water levelchangeWater levelchange

5

13350
13300

4

13250
3
13200
TRACTEBEL
VTT
GRS
AEKI
VEIKI
KI/IRSN
IVS
VUJE
USNRC

2

1

13150
13100

Pressurizer water level [m]Water level in the cleaning tank [m]

6

13050

13000
5000 10000 15000 20000 25000 30000 35000 40000 45000 50000

00Time [s]Figure 4.4. Water level in the cleaning tank.
Outlet temperature of assemblies No. 1-6 [°C]

1400
VTT
GRS
AEKI.1
AEKI.2
BME-NTI
VEIKI
KI/IRSN
IVS
VUJE
USNRC

1200

1000

800

600

400

200

0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.5. Outlet temperature of assemblies No. 1-6.
Outlet temperature of assembly No. 12 [°C]

1600
VTT
GRS
AEKI.2
BME-NTI
VEIKI
KI/IRSN
IVS
VUJE
USNRC

1400
1200
1000
800
600
400
200
0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.6. Outlet temperature of assembly No. 12.

62

Coolant temperature above the assemblies
in the cleaning tank [°C]1200
TRACTEBELVTTGRSAEKI.1
AEKI.2
BME-NTI
VEIKI
KI/IRSN
IVS
VUJE
USNRC

1000

800

600

400

200

0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Inlet flowrate per assembly of assemblies No. 1-6 [kg/s]

Figure 4.7. Coolant temperature above the assemblies in the cleaning tank.
1.0
VTT
GRS
AEKI.1
AEKI.2VEIKI
KI/IRSN
IVS
VUJE
USNRC

0.8
0.6
0.4
0.2
0.0
-0.2
-0.4
-0.6
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Inlet flowrate per assembly of assemblies No. 1-6 [kg/s]

Figure 4.8. Inlet flow rate per assembly of assemblies No. 1-6.
0.40
VTT
GRS
AEKI.1
AEKI.2VEIKI
KI/IRSN
IVS
VUJE
USNRC

0.35
0.30
0.25
0.20
0.15
0.10
0.05
0.00
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000

Time [s]Figure 4.9. Inlet flow rate per assembly of assemblies No. 1-6 (first 10000 s).

63Inlet flowrate per assembly of assembly No. 12 [kg/s]

0.6
VTT
GRS
AEKI.1
AEKI.2VEIKI
KI/IRSN
VUJE
USNRC

0.4

0.2

0.0

-0.2

-0.4
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Inlet flowrate per assembly of assembly No. 12 [kg/s]

Figure 4.10. Inlet flow rate per assembly of assembly No. 12.
0.07
VTT
GRS
AEKI.1
AEKI.2VEIKI
KI/IRSN
VUJE
USNRC

0.06

0.05

0.04

0.03

0.02

0.01
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000

Time [s]By-pass flowrate due to incorrect seating [kg/s]

Figure 4.11. Inlet flow rate per assembly of assembly No. 12 (first 10000 s).
8
TRACTEBELVTTGRSAEKI.1
AEKI.2VEIKI
KI/IRSN
VUJE
USNRC

7
6
5
4
3
2
1
0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.12. By-pass flow rate due to incorrect seating.

64

By-pass flowrate through perforations per
assembly of assemblies No. 1-6 [kg/s]

0.6
VTT
GRS
AEKI.1
AEKI.2
BME-NTI
VEIKI
KI/IRSN
IVS
VUJE
USNRC

0.5

0.4

0.3

0.2

0.1

0.0

-0.1
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.13. By-pass flow rate through perforations per assembly of assemblies No. 1-6.

By-pass flowrate through perforations per
assembly of assemblies No. 1-6 [kg/s]

0.24
VTT
GRS
AEKI.1
AEKI.2
BME-NTI
VEIKI
KI/IRSN
IVS
VUJE
USNRC

0.22
0.20
0.18
0.16
0.14
0.12
0.10
0.08
0.06
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000

Time [s]Figure 4.14. By-pass flow rate through perforations per assembly of assemblies No. 1-6 (first
10000 s).
Flowrate through the air letdown valve [kg/s]

0.03
TRACTEBEL
VTT
GRS
AEKI
KI/IRSN
VUJE
USNRC

0.02

0.01

0.00

-0.01
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.15. Flowrate through the air letdown valve.

65350000
TRACTEBEL
VTT
GRS
VEIKI
KI/IRSNVUJE
USNRC
Decay heat

Heat loss to the environment [W]

300000250000200000150000100000

50000

0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.16. Heat loss to the environment.
900
TRACTEBEL
VTT
GRS
AEKI
VEIKI
KI/IRSN
IVS
VUJE
USNRC

Axial cladding temperature profile ofassemblies No. 1-6 [°C]

800
700
600
500
400
300
200
100
0
0

1

2

3

Node position [m]Figure 4.17. Axial cladding temperature profile of assemblies No. 1-6.
900
TRACTEBEL
VTT
GRS
AEKI
VEIKI
KI/IRSN
IVS
VUJE
USNRC

Axial cladding temperature profile ofassembly No. 12 [°C]

800
700
600
500
400
300
200
100
0
0

1

2

3

Node position [m]Figure 4.18. Axial cladding temperature profile of assembly No. 12.

66

Max. fuel temperature of assemblies No. 1-6 rod [°C]2000

TRACTEBEL
VTT
GRS
AEKI
KI/IRSN
IVS
USNRC

1800
1600
1400
1200
1000
800
600
400
200
0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Max. fuel temperature of assembly No. 12 rod [°C]

Figure 4.19. Max. fuel temperature of assemblies No. 1-6 rod.
2000

TRACTEBEL
VTT
GRS
AEKI
KI/IRSN
IVS
USNRC

1800
1600
1400
1200
1000
800
600
400
200
0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.20. Max. fuel temperature of assembly No. 12 rod.
2200
TRACTEBEL
VTT
GRS
AEKI
VEIKI
KI/IRSNIVS
USNRC

Max. cladding temperature of assemblies
No. 1-6 rod [°C]

2000
1800
1600
1400
1200
1000
800
600
400
200
0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.21. Max. cladding temperature of assemblies No. 1-6 rod.

67

2000
TRACTEBEL
VTT
GRS
AEKI
VEIKI
KI/IRSNIVS
USNRC

Max. cladding temperature of assembly
No. 12 rod [°C]

1800
1600
1400
1200
1000
800
600
400
200
0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.22. Max. cladding temperature of assembly No. 12 rod.

Max. cladding temperature of assemblies
No. 25-30 rod [°C]2000
TRACTEBEL
VTT
GRS
AEKI
VEIKI
KI/IRSN
IVS
USNRC

1800
1600
1400
1200
1000
800
600
400
200
0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.23. Max. cladding temperature of assemblies No. 25-30 rod.
30

Fuel rod internal pressure of assemblies
No. 1-6 [bar]

28

VTT
GRS
AEKI
KI/IRSN
IVS

26
24
22
20
18
16
14
12
10
8
6
4
2
0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.24. Fuel rod internal pressure of assembly No. 1-6.

68

Max. local cladding oxidation of a fuel rod [%]

110
VTT
GRS
AEKI
KI/IRSN
IVS
USNRC

100
90
80
70
60
50
40
30
20
10
0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.25. Max. local cladding oxidation of a fuel rod.
100
GRS
KI/IRSNIVS
USNRC

Max. local oxidation of a shroud [%]

90
80
70
60
50
40
30
20
10
0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Total Zr oxidation of all fuel rods and shrouds [%]Figure 4.26. Max. local oxidation of a shroud.
50
TRACTEBEL
VTT
GRS
AEKI
KI/IRSN
IVS
USNRC

40

30

20

10

0
0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.27. Total Zr oxidation of all fuel rods and shrouds.

69Hydrogen production [kg/s]

10

-2

10

-3

10

-4

10

-5

TRACTEBEL
VTT
GRS
AEKI
KI/IRSN
IVS
USNRC

0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.28. Hydrogen production rate.

External diameter of fuel rods inassemblies No. 1-6 at time = 34200 s [mm]13
TRACTEBEL
VTT
GRS
AEKI
KI/IRSN
IVS
USNRC

12

11

10

9
0

1

2

3

Node position [m]Figure 4.29. External diameter of fuel rods in assemblies No. 1-6 at time = 34200 s.

Oxide layer thickness of fuel rods inassemblies No. 1-6 at time = 34200 s [mm]1000
TRACTEBEL
VTT
GRS
AEKI
KI/IRSN
IVS
USNRC

900
800
700
600
500
400
300
200
100
0
0

1

2

3

Node position [m]Figure 4.30. Oxide layer thickness of fuel rods in assemblies No. 1-6 at time = 34200 s.

70Oxide layer thickness of fuel rods inassemblies No. 25-30 at time = 34200 s [mm]1000
TRACTEBEL
VTT
GRS
AEKI
KI/IRSN
IVS
USNRC

900
800
700
600
500
400
300
200
100
0
0

1

2

3

Node position [m]Figure 4.31. Oxide layer thickness of fuel rods in assemblies No. 25-30 at time = 34200 s.
17

Integral activiy release of isotope

133

Xe [Bq]10

TRACTEBEL
GRS
AEKI
VEIKI
USNRC
Measured

16

10

15

10

14

10

13

10

12

10

0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.32. Integral activity release of isotope 133Xe.
16

10

131

I [Bq]

TRACTEBEL
GRS
AEKI
VEIKI
USNRC
MeasuredIntegral activiy release of isotope

15

10

14

10

13

10

12

10

0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.33. Integral activity release of isotope 131I.

71

16

Cs [Bq]10

TRACTEBEL
GRS
AEKI
VEIKI
USNRC
Measured15

Integral activiy release of isotope

137

10

14

10

13

10

12

10

11

10

0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.34. Integral activity release of isotope 137Cs.
15

10

132

Te [Bq]

GRS
AEKI
VEIKI
USNRC
MeasuredIntegral activiy release of isotope

14

10

13

10

12

10

11

10

0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.35. Integral activity release of isotope 132Te.
14

141

Ce [Bq]

10

VEIKI
USNRC
Measured13

Integral activiy release of isotope

10

12

10

11

10

10

10

0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.36. Integral activity release of isotope 141Ce.

72

Integral activiy release of isotope

242

Cm [Bq]

10

12

10

11

10

10

VEIKI
USNRC
Measured

9

10

8

10

7

10

0

5000

10000

15000

20000

25000

30000

35000

40000

Time [s]Figure 4.37. Integral activity release of isotope 242Cm.
5

OUTLOOKThe OECD-IAEA Paks Fuel Project was a valuable international programme that was actively
supported by 9 Member States. Phase 1 included a thorough development of a comprehensive
database to provide input data for severe accident analysis calculations. In support of the
project, numerous simulations of the thermal hydraulic response, the fuel behaviour, and thefission product source term were performed by the participants. The three workshops
provided an excellent forum to exchange ideas, to discuss the progression of events and key
phenomena in the incident, and to assess the capabilities of various accident analysis codes. It
is recommended that future programs used to investigate nuclear incidents include theseeffective features.
AEKI has performed an experimental simulation of the Paks-2 cleaning tank incident in theCore Degradation Experiment (CODEX) facility. The experiments in the CODEX facility
simulated the whole scenario of the incident using electrically heated fuel rods. The test
conditions were selected in such a way that several unknown parameters of the incident werecovered. The final state of the fuel rods showed many similarities with the conditions
observed after a severe core damage incident at a nuclear power plant. Hence, it is probablethat the thermal conditions and chemical reactions were also similar in the tests and in theincident. Since the initial and boundary conditions were well controlled in the experiments, it
is recommended that follow-on CODEX calculations be performed by the participants andpresented at a future working group meeting.
The damaged Paks fuel has been removed from the cleaning tank and is stored in special
containers in the spent fuel storage pool of the Paks nuclear power plant. The nuclear power
plant intends to send all damaged fuel to Russia for reprocessing. Phase 2 was intended tocover a hot cell examination of the damaged fuel. At present, there is no funding or
programme to perform Phase 2. The following benefits were envisioned from the Phase 2
program.
•

The damaged Paks fuel represents actual irradiated nuclear fuel from a power reactor that
endured severe damage. Most of our current knowledge on the behaviour of fuel under
severe accident conditions is based on small-scale experiments which use electricalheating. The examination of the Paks fuel could confirm or enhance the understanding of73

ballooning, oxidation, hydrogen uptake, and change of mechanical properties under severe
accident conditions.
•

The similarities between the Paks-2 event and a large break loss-of-coolant accident(LOCA) include high temperature oxidation in steam and quench by cold water. However,the differences are significant. These include the distance between assemblies, low decay
heat (2 weeks after shutdown), very long oxidation time, stagnant steam volume, andhydrogen-rich atmosphere. In addition, the Paks fuel has a very sharp transition zone. The
bottom of the fuel was under water and the top was heavily oxidized. Between thoseregions, there is a transition zone where varying levels of fuel damage exist. There is
uncertainty when the fuel will fail under heavily oxidized conditions. An examination of
the transition zone in the Paks fuel would provide data on the fuel cladding strength andintegrity as function of oxidation under beyond design basis conditions.

•

Many severe accident analysis codes do not model the mechanical integrity of oxidized
fuel during a reflooding event. It is believed that the Paks fuel was extensively damaged
during the reflood of the cleaning tank. A characterization of the cladding strength at theassembly failure locations as described in the previous bullet could lead to more realistic
thermo-mechanical failure models for the severe accident computer codes.

The removal of the fuel from the Paks cleaning tank was videotaped. Any further
characterization of the event should include a review of the videotapes to further quantify thephysical damage state. The review should include any observations on the axial, radial, andazimuthal damage patterns (e.g. for insights into heat loss and heat-up behaviour), a best
estimate level depression, qualitative observations on the material strength and behaviour
when cutting the debris for removal, a description of the debris on the baseplate and lower
vessel head (e.g. whole pellets, size of oxidized cladding flakes), and a description of picturesof rod ballooning or other damage insights where views were possible. It is recommended touse the videotapes for the further evaluation of the state of the damaged Paks-2 fuel.
The database of the OECD-IAEA Paks Fuel Project includes valuable measured data on theactivity release from damaged fuel under wet storage conditions. The use of such data wasbeyond the scope of the project, since the numerical analyses focused on the simulation of theincident. It is recommended that this part of the database be used in other (e.g. waterchemistry, leaching of fuel, long term waste disposal) projects in the future.
6

CONCLUSIONSThe original objectives of the OECD-IAEA Paks Fuel Project were successfully met. The
basic information that was necessary for the simulation of Paks-2 incident had been collected.
Numerical analyses based on a common database were carried out with sophisticated models
using codes that are used in the evaluation of the safety of nuclear power plants. The
simulations covered thermal hydraulics, fuel behaviour and activity release aspects of thePaks-2 incident.
The calculations captured well some events, such as:• the time and speed of water level change in the cleaning tank and in the surrounding
pools,• the time of first fuel failures that very probably took place as a result of ballooning andburst of fuel rods at high temperature,• the rate of released activity from the fuel.

74

The numerical analysis improved the understanding of Paks-2 event and helped to makeprecise some unknown parameters of the incident, such as:• the by-pass flow at low flow rate amounted to 75-90% of the inlet flow rate that led to
the formation of steam volume,• the maximum temperature in the tank was between 1200-1400 ºC,• the degree of zirconium oxidation reached 4-12%,• the mass of produced hydrogen was between 3 and 13 kg.
The OECD-IAEA Paks Fuel Project improved the current knowledge on fuel behaviour under
accident conditions and led to recommendations for some further actions for research in this
area.

75

REFERENCES[1]

OKB HYDROPRESS, Analysis of operation of fuel assemblies with reduced flow rate
due to deposit in the Paks NPP, Final Technical Report, U-213-TI-1762, OKB
HYDROPRESS, Podolsk, (2003)[2] SCHUNK, J., BEIER, M., KOVÁCS, F., S. MIKÓ, P. TILKY, H.O. BERTHOLD, I.
JANZIK, G. MARQUARDT, Fuel Assembly Chemical Cleaning at Paks Nuclear
Power Plant, IAEA Meeting on Fuel Failures and Mitigation, Bratislava (2002).
[3] ASZÓDI, A., The Paks-2 Incident, The Planned Recovery Work and the Restart of Unit
2, 47th Symposium of Hungarian Secondary School Physics Teachers, Hungary (2004).
[4] HUNGARIAN ATOMIC ENERGY AUTHORITY, Report to the Chairman of the
Hungarian Atomic Energy Commission on the Authority’s investigation of the incident
at Paks Nuclear Power Plant on 10 April 2003, (2003).
[5] PAKS NUCLEAR POWER PLANT EVENTS INVESTIGATION GROUP, “Inquiry
Report of Fuel Damage in the Cleaning Facility of FANP on 10-11 April 2003”, Reportfor HAEA NSD, NBI 03.11, B20301, (2003)[6] ASZÓDI, A., BOROS, I., CSIGE, A., LÉGRÁDI, G., Thermal hydraulic analysis of thePaks Incident, 2nd Symposium of Nuclear Technology, Budapest (2003).
[7] GYŐRI, CS., HÓZER, Z., TÓTH, I., TROSZTEL, I., Analyses of the 10th April 2003
Paks Incident, Hungary (2003).
[8] PEREZ-FERÓ, E., WINDBERG, P., HÓZER, Z., HORVÁTH, M., NAGY, I.,PINTÉR-CSORDÁS, SZABÓ, E., KULACSY, K., Experimental Database of E110
Cladding Oxidised in Hydrogen Rich Steam, Transactions of TOPFUEL, Salamanca,
Spain (2006) 605-609.
[9] HÓZER, Z., Fuel Failures During the Paks-2 Incident, Enlarged Halden ProgrammeGroup Meeting, Proceedings of the Fuels and Materials Sessions, HPR 366 Vol. 2,
paper F4.1, Norway (2007).
[10] HÓZER, Z., WINDBERG, P., NAGY, I., VIMI, A., VÉR, N., MATUS, L., KUNSTÁR,M., Experimental Simulation of the Paks-2 Cleaning Tank Incident, Transactions,
SMiRT 19, Toronto (2007).
[11] KÓPHÁZI, J., FEHÉR, S., CZIFRUS, S., BERKI, T., CSOM, G., Detailed 3D analysisof the Reactor Physics of the Damaged Fuel inside the Cleaning Tank, 3rd Symposiumof Nuclear Technology, Budapest (2004).
[12] FEHÉR, S., KÓPHÁZI, J., CZIFRUS, S., BERKI, T., “Detailed 3D subcriticality
analysis of the damaged fuel”; Internal Research Report of TU Budapest, BME-NTI281/2004, Budapest (2004).
[13] INTERNATIONAL ATOMIC ENERGY AGENCY, “Evaluation of examination
results of the Hungarian Atomic Energy Agency on the incident in Paks NPP during
fuel cleaning on 10. April 2003”, TC Project: HUN/9/022, IAEA-TCR-01769,
IAEA,Vienna (2003).
[14] CSERHÁTI, A., “The technology of the recovery work in Paks Unit 2”, 3rd Symposiumof Nuclear Technology, Budapest (2004).

77

ABBREVIATIONS
AEKI

Hungarian Academy of Sciences KFKI Atomic Energy Research Institute

AMDA

Automatic Mobile Decontaminating Instrument

CFD

Computerized fluid dynamics

CSNI

Committee on the Safety of Nuclear Installations

FA

Fuel assembly

FE

Finite element

FGR

Fission gas release

FP

Fission product

GRS

Gesellschaft für Anlagen- und Reaktorsicherheit mbH

INESInternational Nuclear Event Scale

IRSN

Nuclear Safety and Radioprotection Institute

ISP

International Standard Problem

KI

Russian Research Centre “Kurchatov Institute”

LOCA

Loss-of-coolant accident

PCMI

Pellet-cladding mechanical interaction

SFD

Severe fuel damage

79

CONTRIBUTORS TO DRAFTING AND REVIEW
Aszódi, A.

Budapest University of Technology andEconomics, Institute of Nuclear Techniques(BME-NTI), Hungary.

Boros, I.

Budapest University of Technology andEconomics, Institute of Nuclear Techniques(BME_NTI), Hungary.

Guillard, V.

Institute de radioprotection et du Sureté
Nucléaire (IRSN), France.

Győri, Cs.

Atomic Energy Research Institute (AEKI),
Hungary.

Hegyi, G

Atomic Energy Research Institute (AEKI),
Hungary.

Horváth, G. L.

VEIKI Institute for Electric Power Research,
Hungary.

Hózer, Z.

Atomic Energy Research Institute (AEKI),
Hungary.

Nagy, I.

Atomic Energy Research Institute (AEKI),
Hungary.

Junninen, P.

Technical Research Centre (VTT), Finland.

Kobzar,V.

Kurtchatov Institute, Russian ResearchCentre, Russian Federation.

Légrádi, G.

Budapest University of Technology andEconomics, Institute of Nuclear Techniques(BME-NTI), Hungary.

Molnár, A.

Atomic Energy Research Institute (AEKI),
Hungary.

Pietarinen, K.

Technical Research Centre (VTT), Finland.

Perneczky, L.

Atomic Energy Research Institute (AEKI),
Hungary.

Makihara, Y.

nternational Atomic Energy Agency

Matejovic, P.

Slovak Academy of Sciences (IVS),
Slovakia.

Szabó, E.

Atomic Energy Research Institute (AEKI),
Hungary.

Trambauer, K.

Gesellschaft für Anlagen- undReaktorsicherheit mbH, Germany.

Trosztel, I.

Atomic Energy Research Institute (AEKI),
Hungary.

Varpoorten, J.

Tractebel Engineering GDF-Suez, Belgium.

Vitanza, C.

OECD/NEA

81

Vogel, M.

Nuclear Power Plant Research InstituteTrnava (VUJE), Slovakia.

Volchek, A.

Kurtchatov Institute, Russian ResearchCentre, Russian Federation.

Wagner, K.C.

Sandia National Laboratories, USA.

Zvonarev,Y.

Kurtchatov Institute, RussianCentre, Russian Federation.

82

Research

No. 21, July 2006

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spine:5.455 mm
81 pages

OECD-IAEA Paks Fuel Project

OECD–IAEA Paks Fuel Project
Final Report

INTERNATIONAL ATOMIC ENERGY AGENCY
VIENNA

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2010-06-29 10:12:33