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62 Scientific and Technical Report 2007 - IRSN

2Accidents in nuclear facilities

IRSN - Scientific and Technical Report 2007 63

2 AccIdeNtS in nuclear facilities . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 64

2.1 FIRSt ReSultS of the Phebus FPT3 test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 66

2.2 Study oF RutheNIum chemIStRy in the containment building under severe accident conditions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 73

newsflashnewsflashnewsflashnewsflashnewsflashnewsflash

2.3 KIcK-oFF FoR chIP, an experimental program in Grenoble . . . . . . . . . . . . . . . . 80

2.4 uF6 behAvIoR IN AN AccIdeNtAl ReleASe coNtext Studies and experiments to quantify accidental UF6 release in front-end fuel cycle facilities . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 81

2.5 Study oF PhySIcAl PheNomeNA and consequences associated with criticality accidents . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 89

2.6 ANAlySIS oF the mechANIcAl behAvIoR oF containment on CPY 900 MWe PWRs under severe accident conditions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 98


2.7 INItIAl ReSultS of the Prisme Door campaign . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 109

2.8 locAl-ScAle theRmohydRAulIcS R&d to support LOCA studies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 110

2.9 A New ReSeARch PRogRAm to study fuel assembly behavior at La Hague storage pools in the event of accidental dewatering . . . . . . . . . . 112

2.10 Key dAteS: Dissertations defended and other major events . . . . . . . . . . . . . . . 113


An important aspect of the Institute's mission consists of

preventing nuclear facility accidents and assessing mea-

sures to limit the potential consequences to human health

and the environment. These objectives are achieved through mission-

oriented research programs to improve knowledge of accident

phenomena, and through the development of computer codes and

assessment methods. Once validated, these tools are used to assess

accident risks and analyze the measures taken by operators to

manage these risks.

IRSN has focused considerable efforts on research in these areas,

presented in detail in the following articles.

Pressurized water reactor core meltdown accidents, such as the

1979 accident in the US at the Three Mile Island (TMI) nuclear

power plant, have been the subject of several research programs in

France and throughout the world. Core meltdown accidents are

considered highly improbable, due in large part to measures taken

since the TMI accident, and they can occur only if several independent

safety systems fail. Nonetheless, the potential consequences,

namely radioactive release to the environment, justify pursuing

research efforts in this area. Three articles have been dedicated to

this topic.

The first presents results from the final test of the international

program Phebus FP (fission products). It highlights the influence of

control rod materials on iodine release inside containment.

The second article describes research on the behavior of ruthe-

nium in the containment building. Large quantities of this highly

radiotoxic element could be emitted during certain accident

scenarios, partly in gas form.

The third article reports on mechanical simulations of a 900 MWe

PWR containment. Serving as the ultimate barrier around a reactor,

the containment building prevents radioactive release to the envi-

ronment. A multiscale approach was used to finely evaluate the

strength limits for this structure, a critical component in nuclear

safety.

A brief article sums up the ChIP program that IRSN is conducting

in partnership with the CNRS. ChIP focuses on the chemical forms

iodine can take during transfer from the degraded reactor core to

the containment building, and on the volatility of these iodine

species.

To control uranium hexafluoride containment in fuel cycle facilities,

models are needed to understand dispersion of heavy UF6 gas, as well

as the chemical reactions between UF6 and steam, which produce

caustic hydrofluoric acid. An in-depth article examines progress made

in this important area.

Another significant risk requiring preventive measures, particularly

in fuel cycle facilities, is criticality risk. This phenomenon can occur

in fissile material if the geometry and material configuration are

AccIdeNtS in nuclear facilities

Michel SchwarzPrevention of Major Accidents


"critical", triggering an uncontrolled nuclear fission chain reaction.

Criticality accidents can result in serious injury, especially for workers

in the immediate vicinity. The state of knowledge on criticality

accidents is the focus of another in-depth article.

A nuclear facility fire can have severe consequences, particularly if

safety functions are jeopardized. A brief article presents initial results

from the international PRISME research program on fire propagation

in confined and ventilated facilities.

Spent fuel assemblies are stored in vast pools at the La hague

reprocessing plant. Pool water dissipates the residual heat released

by these assemblies. A brief article describes a new research program

IRSN is conducting with AREVA, to assess the consequences of a pool

dewatering accident.

Operator demand for higher fuel burnup and longer cycle times is

driving the development of ever more sophisticated computer tools,

capable of analyzing the impact of these changes on reactor safety. This

is the backdrop to a short article on current R&D at the Institute aimed

at developing and validating LOCA assessment tools.

2. 1


The Phebus FP program consisted of five tests successfully

conducted from 1993 to 2004 [Clément et al., 2006]. FPT3 was

the fifth and final test, carried out from November 18 to 22,

2004. Its specific feature consisted of testing the fuel using the

neutron-absorbing boron carbide (B4C) material used in 1300 MW

PWRs, whereas earlier Phebus tests had employed the silver-

indium-cadmium (Ag-In-Cd) alloy used in 900 MW PWRs.

Raw experimental data collected during the test, and in subsequent

non-destructive test campaigns, have been processed and

researchers are currently analyzing overall data consistency. This

article presents insights gained from FPT3 results [Clément et

al., 2005a; March et al., 2006; Simondi-Teisseire et al., 2006],

which will be consolidated by non-destructive tests currently

underway.

Experimental facility

In the Phebus facility, experimental conditions are representative

of a PWR core meltdown accident [Schwarz et al., 1999; Clément

Béatrice Simondi-Teisseire, Bruno Biard, Jérôme Guillot, Christelle Manenc, Philippe March, Frédéric PayotExperimentations and Measurements of Accidental Releases Laboratory

Since the accident in the US at the Three Mile Island Unit 2 (TMI-2) nuclear power plant on March 28, 1979,

resulting in partial core meltdown and limited fission product release, a number of experimental safety

research programs have been conducted by different organizations around the world. In 1988, IPSN

launched PhebUS FP, a major international research program on severe water reactor accidents (involving

core meltdown). Conducted in the like-named experimental reactor operated by the CeA, PhebUS involved

a series of integral experiments, reproducing expected physical core meltdown phenomena as realistically

as possible. experimental results from the PhebUS FP program combined with results from separate effect

tests are key to validating the simulation codes used in light water reactor safety analyses [birchley et al.,

2005; Clément, 2003a; Clément et al., 2006; evrard et al., 2003; Schwarz et al., 1999; Schwarz et al., 2001], in

particular ASTeC [Van Dorsselaere et al., 2004], developed by IRSN in collaboration with GRS, as well as the

ICARe/CAThARe code.

FIRSt ReSultS of the Phebus FPT3 test

Accidents in nuclear facilities 2. 1


et al., 2003b], appropriate for the study of fuel rod and absorber

rod degradation, molten pool formation, and release and transport

of degradation products (fission products emitted from the fuel,

gases and/or aerosols from fuel rod/absorber rod degradation)

in the reactor coolant system and containment. Researchers are

particularly interested in iodine behavior, since environmental

release of iodine in the days following core meltdown could have

considerable radiological impact.

FPT3 investigated physical phenomena occurring in the following

systems:

the reactor core ( 1 ), simulated by an assembly comprising

18 fuel rods previously irradiated in the BR3 reactor with burnup

of 24.5 GWd/tU, two instrumented rods containing fresh fuel,

and a neutron-absorbing boron carbide rod; the fuel rods have

Zircaloy cladding and the absorber rod has steel cladding with

a Zircaloy guide tube;

the reactor coolant system simulated by experimental circuits,

consisting of a 700°C hot leg ( 2 ) and a 150°C cold leg ( 4 ),

connected by an inverted U-shaped tube 4 m high simulating

the steam generator ( 3 ), where a sharp drop in coolant

temperature takes place;

the containment building, simulated by a 10 m3 vessel(1) with

an electropolished surface, a 120-liter tank at its base filled with

ph 5 buffer solution simulating the reactor sump 6 (for technical

reasons, the sump represented only 10% of the vessel cross-

section in Phebus), a gas containment ( 5 ) and, in the upper part,

condensing, cooled, painted surfaces(2) ( 7 ). The cold leg discharges

into the open area of the vessel, simulating a break downstream

from the steam generator.

These three zones are reproduced at a scale of 1:5000 with

respect to a 900 MWe PWR (Figure 1) and are equipped with

various instruments to measure flow rate, temperature, radiation

(high count-rate gamma spectrometry), concentrations of

hydrogen, oxygen, and carbonaceous gases, and to take sequential

samples of experimental circuit fluid, containment atmosphere,

and sump liquid.

Non-destructive measurements were performed in the facility

after the test to quantify the gamma emitters retained in the

experimental circuit and vessel samples, and to characterize fuel

degradation (using X-ray radiography, computed tomography,

and gamma spectrometry to establish the γ emitter distribution

profile for the rod assembly).

(1) Referred to as "vessel" in the rest of this article.

(2) Referred to as "condensers" in the rest of this article.

Figure 1 The Phebus FP facility.

Scale = 1:5000

PWR

PHEBUS FP

Paint

Containment

Reactorcoolantsystem

Break

1

1

2

2

3

3

4

4

5

5

7

7

6 6

2. 1


diffusiophoresis on condensers, and deposition on electropolished

vessel walls);

any potential FP poisoning effect on catalytic hydrogen

recombiners used in nuclear reactor containments.

FPt3 test scenario

Prior to the actual experimental phase, the test rod assembly

was re-irradiated in the Phebus reactor for a period of 8.5 days,

to obtain a representative inventory of short-lived fission

products (e.g. iodine-131 with a half-life of around 8 days).

Re-irradiation was followed by a transition phase (rod assembly

drying , limit condition adjustment, and xenon poisoning

reduction), lasting around 37 hours.

The degradation phase lasted around 5 hours. Experimental

circuit pressure was kept at 0.2 MPa and the steam injection

flow rate in the lower part of the rod assembly at 0.5 g/s. Power

in the rod assembly was increased by successive ramps and

plateaus until the degradation objectives were reached, at which

point the reactor was shut down. The fuel rod assembly was then

cooled for around one hour, followed by isolation of the vessel

from the circuit.

Between reactor shutdown and the end of steam injection, the

hydrogen recombiner coupons were placed in the containment

atmosphere for 30 minutes.

The experimental phase then moved into a long-term phase

lasting 4 days, with three successive stages:

an "aerosol" stage lasting around 37 hours, aimed at analyzing

aerosol deposition mechanisms in the vessel;

a 13-minute washing phase, in which aerosols deposited by

gravitational settling on the hemispheric vessel bottom were

transferred to the sump;

a 2-day chemistry stage, dedicated to iodine chemistry in the

sump and containment atmosphere, with particular focus on

iodine speciation. Water temperature during this stage was

brought to 90-100°C to favor a representative 0.73 g/s

evaporat ion/condensat ion cyc le between sump and

condensers.

main results for rod assembly degradation

During the experiment, reactor power was increased gradually

(Figure 2). The first plateaus were maintained at low levels to

FPt3 objectives

FPT3 objectives can be divided into three groups according to

the relevant system: test device, experimental coolant system,

and vessel.

Test device

The main objective was three-fold: to obtain substantial

degradation in the fuel rods and neutron-absorbing rod, substantial

volatile FP release in a low-pressure hydrogen-rich atmosphere,

and an overall displaced fuel weight of around 1 kg, of the 10 kg

initially present in the rod assembly.

Experimental coolant system

The main objective was to apprehend emission of fission products

from the fuel, gases and aerosols produced by fuel rod and

absorber rod degradation, as well as transport and deposition

of fission products in the reactor coolant system under low

pressure (0.2 MPa). Another objective was to collect data on

fission product chemistry, in particular interactions with the

walls of high-temperature lines and with carbon and boron

compounds produced by B4C oxidation. Methane formation is

particularly important because it can promote organic iodine

formation. To investigate this phenomenon, the FP release phase

must be sufficiently long and take place under highly reducing

conditions (achieved by transforming nearly all injected steam

into hydrogen through oxidation of B4C and Zircaloy

cladding).

Vessel

The main objective was to study FP physicochemistry in the

hours and days following their emission from the rod assembly,

along with the effects of boron and carbon compounds.

Researchers focused on iodine radiochemistry in the sump water

and vessel atmosphere, using several dedicated instruments.

Painted surfaces installed in both the sump and the containment

area (condensing cooled painted surfaces) provided a source of

organic compounds capable of interacting with iodine.

In addition to these objectives, researchers also aimed to

characterize:

the size of aerosols released in the vessel and the aerosol

deposition processes (gravitational settling to vessel bottom,



failure in the rod assembly midplane, with a maximum rod assembly

temperature of 1650°C at 500 mm. As in FPT2, no significant

material displacement was detected during the main oxidation

phase, in contrast to oxidation during FPT1, which was more

violent;

displacement of materials, particularly the fuel, beginning in

the last plateau and continuing until the start of the subsequent

power ramp phase, likely resulting in formation of a small molten

pool.

Molten material progression in FPT3 was apparently faster and

more penetrating than in earlier tests, given that cooled materials

were found under the rod support plate during post-test

examinations. This could be explained by a lower melting

temperature, due to the presence of boron and steel compounds

in the U-Zr-O mixture (containing melted Zircaloy cladding and

make sure that temperatures reached the expected values.

On-line measurements of the rod assembly, experimental circuits,

and vessel indicated that the following events occurred:

cladding failure, occurring around the rod assembly midplane

(–500 mm) at a temperature close to 800°C, as in earlier tests;

absorber rod failure at around 500 mm. Maximum guide tube

temperature was roughly 1450°C, at least 100 degrees higher

than in earlier tests (which used an Ag-In-Cd absorber rod). CO

was first detected in the containment atmosphere a few seconds

after absorber rod failure, which is consistent with temperatures

measured in the experimental rod assembly. CO2 entered the

vessel later, at the end of the first oxidation phase, once the

hydrogen concentration in the experimental circuit had fallen. The

concentration of CH4 was too close to the detection limit to draw

any conclusions regarding its formation.

the main oxidation phase, occurring shortly after absorber rod

0 18 0004 000 6 000 8 000 10 000 12 000 14 000 16 000

2 650

2 400

2 150

1 900

1 650

1 400

1 150

900

650

400

150

Time (s) - T0: 11h49m00s2 000

Start of test: 11:49:00 (0 s)Duration: 04:49:30 (17,370 s)Reactor shutdown: 16:38:30 (17,370 s)

Calibration Plateau P4

Cooling

HeatingPre-oxidation

Oxidation

P1 start: 360 sP1 end: 3960 s



P4a

P4b

P4c

P4 start: 11,100 sP4 end: 15,420 s

Start of heating phase: 15,420 sReactor shutdown: 17,370 s

Hydrogen Insulation - 100 mm - 349° Insulation - 100 mm - 169°

Cladding failure

B4C failure

1st riseunder lower grid

2nd riseunder lower grid

Temperature (°C)

Hydrogen first detected: 8440 sStart of oxidation: 9480 sStart of low-steam phase: 10,000 s

Co

re p

ow

er (a

.u.)

/ SD

HY

700

hyd

rog

en (a

.u.)/

SD

HY

700

hyd

rog

en (a

.u.)

Fuel - 300 mm Fuel - 500 mmCore power

Figure 2 General timeline of FPT3 degradation phase. Start time is set at the beginning of the reactor power ramps.

2. 1


dissolved uranium oxide fuel). Total weight of relocated UO2 was

around 1 kg, in agreement with experimental objectives, although

it was spread over a greater height than in earlier tests.

FPT3 hydrogen production kinetics were similar to FPT2 kinetics,

and more gradual than FPT0 and FPT1: the lower steam injection

rate – around 0.5 g/s in FPT2 and FPT3, as opposed to 2 g/s in

FPT0 and FPT1 – resulted in slower progression of the Zircaloy

cladding oxidation front.

hydrogen production lasted longer, however, leading to a

17-minute period in which the volume concentration of hydrogen

in the experimental circuit exceeded 70%. During degradation,

Zircaloy and boron carbide oxidation produced 60 moles of

hydrogen (52-53 moles from Zircaloy and 7-8 moles from boron

carbide). In other words, 73% of Zr and 77% of B4C were oxidized

during the degradation phase.

As in earlier tests, a less-significant oxidation stage occurred in

the final rod assembly degradation phase, when materials flowed

into the lower part of the assembly. Maximum hydrogen

concentration in the experimental circuit reached 20% during

this late oxidation stage, which lasted around 13 minutes.

As expected, moderate but significant rod assembly degradation

was obtained during FPT3, as shown by the X-ray images taken

after the test (Figure 3).

main results from experimental circuits and vessel

FP emission, transport, and deposition in the

experimental circuit

For the elements measured to date, overall release from the

FPT3 rod assembly was similar to overall release in earlier tests

[Clément et al., 2006; Dubourg et al., 2005]. The elements

measured can be classified according to overall release:

elements released in substantial amounts (around 80% of the

initial rod assembly inventory [i.i.]) such as noble gases (e.g.

Xe);

I, Te, and Cs volatile fission products with a released fraction

in the 45-75% range;

elements released in low or very low amounts, such as Ba or

Zr; around 3% of the initial Ba rod inventory was released, and

fractions were much lower for some elements.

As in FPT2, the low steam injection rate (0.5 g/s) resulted in

significant deposition of volatile fission products (Cs, I, Te, and

Mo) in the upper part of the rod assembly. This stands in contrast

to FPT0 and FPT1, in which a higher steam injection rate (2 g/s)

led to deposits downstream (in the upper plenum above the rod

assembly, and in the tube simulating the steam generator). FPT3

also resulted in significant cesium and iodine deposits in the

steam generator tube, representing 9.4% of the initial cesium

inventory and 7.1% of the initial iodine inventory, a two-fold

increase compared to FPT2.

Aerosol release was most significant at the end of the first

oxidation phase and during the fuel displacement phase.

Volatile FP transport through the experimental circuit toward

the vessel began in the first oxidation phase and ended with

reactor shutdown. Xe, I, and Cs entered the vessel at a relatively

stable rate throughout the release phase. Te entered the vessel

with a significant delay compared to Cs and I. Measurements

also indicate that Te was deposited in the hot leg in greater

quantities than the other volatile fission products. Mo was only

measurable in the tank after the first oxidation phase, suggesting

that release was limited in the hydrogen-rich phase and only

became significant in the high-steam phase that followed. This

is consistent with the fact that oxidized Mo is more volatile than

the metal.

Noble gases, which did not react with coolant system surfaces,

reached the containment atmosphere without being retained in Figure 3 X-ray images of Phebus FP rod assembly before and

after FPT0, FPT1, FPT2, and FPT3.



therefore differed significantly from aerosol FP behavior. For

example:

during the degradation phase, iodine was deposited around

three times faster in the vessel than aerosols of the other

elements, due to gaseous iodine absorption by the cooled painted

surfaces, on which 55% of the iodine weight released in the

vessel was deposited;

during wash-down, no significant additions of iodine to the

sump were measured, indicating that very little iodine was

deposited on the hemispheric vessel bottom. This is consistent

with the fact that iodine was mainly in gas form during the

aerosol sedimentation phase.

In the degradation phase, as iodine transport into the vessel

began to slow, gaseous iodine measured in the vessel reached a

maximum of 13% of the initial rod assembly iodine weight,

which is much higher than the maximum values measured in

earlier tests [Clément et al., 2006 ; Girault et al., 2006; Jacquemain

et al., 1999]. Then, as the aerosol stage began, gaseous iodine

fell very rapidly to 0.8% of the initial iodine inventory. This sharp

decline suggests that iodine trapping on the cooled painted

surfaces was very efficient. A similar decrease was observed in

FPT1 and FPT2, but with a lower initial gaseous iodine

concentration in the vessel. During the aerosol stage of FPT3,

the gaseous iodine fraction continued to decrease until it reached

0.1-0.15% of the initial rod inventory weight.

During the chemistry stage (after aerosols deposited on the

hemispheric vessel bottom were washed into the sump), the

gaseous iodine fraction decreased from 0.1-0.15% to a plateau

around 0.03% of the rod assembly inventory (compared to 0.01%

for FPT2), probably indicating a physicochemical iodine

equilibrium inside the vessel.

Throughout the experimental phase, the gaseous iodine fraction

was predominantly (over 75%) inorganic, as it was in FPT2, in

contrast to FPT0 and FPT1, where organic iodine was the major

component. FPT3 experimental data do not indicate significant

iodine desorption from the vessel walls or condensers during

the experimental timescale.

Iodine collected in the sump water (ph 5 ± 0.2, 90°C in aerosol

stage, 100°C in chemistry stage) was mostly in soluble form

throughout the test.

The iodine water solubility observed in FTP3 is consistent with the

use of the B4C rod instead of the Ag-In-Cd rod, which reduces the

trapping efficiency of iodine in the insoluble form AgI [Funke, 1996].

Iodine recovered in the sump, estimated at 5.4% of the initial rod

inventory weight, resulted mainly from condenser draining.

the experimental circuit. By contrast, volatile fission products

(such as I, Cs, and Te) were released in comparable fractions,

around 45-75% of the initial fuel weight, but they reached the

vessel in different amounts: 34% i.i. for iodine, 5% i.i. for tellurium,

and 31% i.i. for cesium. The fractions measured for FPT2 were

much higher: 57% i.i. for iodine, 28% i.i. for tellurium, and 41%

i.i. for cesium.

For FPT3, the lower Te and Cs fractions carried into the vessel

can be attributed to the large deposits of these elements

measured in the hot leg, mainly in the vertical line and steam

generator for Te, and upstream from the steam generator for Cs.

A weight balance must be established prior to any conclusions

about iodine retention in the experimental circuits.

Aerosol behavior in the vessel

The rate of aerosol deposition on condensers in the vessel was

similar for FPT3 and FPT2, consistent with similar condensation

rates (in the degradation phase and early "aerosol" stage) and

involving diffusiophoresis (aerosols entrained by steam

condensation on cooled painted surfaces). The rate of aerosol

deposition by gravitational settling was lower by a factor of two

compared to earlier tests. This can be attributed to smaller

aerosol size and/or lower aerosol concentration in the vessel

(which reduces aerosol agglomeration and hence size), and/or

lower aerosol density. For Cs and Te, however, gravitational

settling remained the major aerosol deposition mechanism in

the vessel, with around 50-60% of the vessel inventory deposited

on the hemispheric bottom. Measurements indicated non-

negligible deposits on the vessel walls, containing around 15%

of the I, Cs, and Te weight initially carried into the vessel at the

end of the "aerosol" stage, a higher proportion than in FPT2 and

FPT1.

Iodine behavior in the vessel

Around 34% of the iodine weight initially contained in the fuel

was carried into the vessel. This is lower than the corresponding

fraction in earlier tests [Clément et al., 2006; Girault et al., 2006;

Jacquemain et al., 1999]. In FPT3, iodine released in the vessel

was mainly in gas form (mean gaseous iodine in the atmosphere

of the vessel during degradation was around 80%). The absorber

rod (B4C in FPT3, rather than the Ag-In-Cd used in previous tests)

appeared to have a significant impact on the physicochemical

form of iodine released in the vessel. In FPT3 iodine behavior

2. 1


and the lower-than-expected liquefaction temperature of the

U-Zr-O-B-steel mixture (known as "corium"). Further studies

[Clément et al., 2005b] are underway, both to provide additional

experimental data and to apprehend the physicochemical

phenomena behind these results. Researchers are also in the

process of transposing the results to reactor conditions, in

order to evaluate their impact on the assessment of any

radioactive release in an accident situation.

conclusions and outlook

In general, FPT3 reached the test objectives satisfactorily

[Albiol et al., 2004]. Certain preliminary results deduced from

available FPT3 data were unexpected, however, particularly

the large gaseous iodine fraction carried into the vessel

during degradation (although it quickly dropped within a few

hours to produce iodine values similar to those in earlier tests)

References

T. Albiol, S. Morin (2004), FPT3 test statement, CPEX/PH/2004-02050, Document Phébus IP/04/560.

J. Birchley, T. haste, h. Bruchertseifer, R. Cripps, S. Güntay and B. Jäckel (2005), Phebus-FP: Results and significance for plant safety in Switzerland, Nuclear Engineering and Design, vol. 235, pp. 1607-1633.

P.D.W. Bottomley, P. Carbol, J.P. Glatz, D. Knoche, D. Papaioannou, D. Solatie, S. Van Winckel, A.-C. Gregoire, G. Grégoire and D. Jacquemain (2005), Fission product and actinide release from the Debris bed test Phebus FPT4: synthesis of the Post test Analyses and of the Revaporisation testing of the plenum samples performed at ITU, International Congress on Advanced Power Plants (ICAPP-05), May 15-19, Seoul, Korea.

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B. Clément, N. hanniet-Girault, G. Repetto, D. Jacquemain, A.V. Jones, M.P. Kissane and P. Von der hardt (2003b), LWR severe accident simulation: synthesis of the results and interpretation of the first Phebus FP experiment FPT-0, Nuclear Engineering and Design, vol. 226, pp. 5-82.

B. Clément, O. De Luze and G. Repetto (2005a), Preliminary results and interpretation of Phebus FPT-3 test, MELCOR Cooperative Assessment meeting, September 20-21, Albuquerque (NM) USA.

B. Clément, N. Girault, G. Repetto and B. Simondi-Teisseire (2006), Les enseignements tirés du programme PhEBUS PF, RST IRSN 2006, 84-96.

B. Clément, R. Zeyen (2005b), The Phebus Fission Product and Source-Term international programs, International Conference Nuclear Energy for New Europe 2005, Bled, Slovenia, September 5-8.

J.C. Crestia, G. Repetto and S. Ederli (2000), Phebus FPT-4 First post test calculations on the debris bed using the ICARE V3 code, Proc. 4th technical seminar on the Phebus FP programme, Marseille, France, March.

R. Dubourg, h. Faure-Geors, G. Nicaise and M. Barrachin (2005), Fission product release in the first two Phebus tests FPT-0 and FPT-1, Nuclear Engineering and Design, vol. 235, pp. 2183-2208.

J.M. Evrard, C. Marchand, E. Raimond and M. Durin (2003), Use of Phebus FP Experimental Results for Source Term Assessment and Level 2 PSA, Proc. 5th Technical seminar on the Phebus FP programme, Aix-en-Provence, France, June 24-26.

F. Funke, G.-U. Greger, A. Bleier, S. hellmann and W. Morell (1996), The reaction between iodine and silver under severe PWR accident conditions, Chemistry of Iodine in Reactor Safety, Workshop proceedings Würenlingen, Switzerland 10-12 June, NEA/CSNI/R(96)6.

N. Girault, S. Dickinson, F. Funke, A. Auvinen, L. herranz and E. Krausmann (2006), Iodine behaviour under LWR accidental conditions: lessons learnt from analyses of the first two Phebus FP tests, Nuclear Engineering and Design, vol. 236, pp. 1293-1308.

D. Jacquemain, N. hanniet, C. Poletiko, S. Dickinson, C. Wren, D. Powers, E. Krausmann, F. Funke, R. Cripps and B. herrero (1999), An Overview of the Iodine Behaviour in the Two First Phebus Tests FPT-0 and FPT-1, OECD Workshop on Iodine Aspects of Severe Accident Management, Vantaa, Finland, May 18-20.

Ph. March et al. (2006), First results of the Phebus FPT-3 test, Proc. of the 14th International Conference on Nuclear Engineering, July 17-20, 2006, Miami, Florida, USA.

M. Schwarz, G. hache and P. Von der hardt (1999), Phebus FP: a severe accident research programme for current and advanced light water reactors, Nuclear Engineering and Design, vol. 187, pp. 47-69.

M. Schwarz, B. Clément and A.V. Jones (2001), Applicability of Phebus FP results to severe accident safety evaluations and management measures, Nuclear Engineering and Design, vol. 209, pp. 173-181.

B. Simondi-Teisseire, B. Biard, J. Guillot, C. Manenc, P. March, F. Payot, C. Gaillard, B. Morassano, M. Pepino (2006), FPT3 Phébus Test: first results on iodine behaviour, Cooperative Severe Accident Research Program (CSARP), September 25-28, Albuquerque (NM) USA.

J.-P. Van Dorsselaere et h.-J. Allelein (2004), ASTEC and SARNET, Integrating Severe Accident Research In Europe, Proc. EUROSAFE Forum, Berlin, Germany, 2004.


context

Safety in nuclear power plants is based on defense-in-depth

and containment of radioactive materials. A key safety measure

to prevent environmental release consists of containing the

radioactive products in the reactor core within three successive

barriers: the fuel cladding, the reactor coolant system (RCS),

and the containment building.

A severe accident (SA) has an extremely low probability of

occurrence(1), since it implies coolant loss concurrent with

partial or total failure of safety systems. Nonetheless, such an

event would result in core meltdown and loss of the first two

containment barriers, allowing the release of fission products

into containment.

Ruthenium (Ru) metal, found in nuclear fuel as a fission product,

is considered to have low volatility; studies show that the Ru

fraction emitted from a UO2 pellet heated in a mixed oxidizing

atmosphere (h2O/h2) to around 2300°C (SA reactor core

temperature) ranges from 1% to 10% [Ducros et al., 2005;

Libmann, 1996]. however, under oxidizing conditions, the metal

species oxidizes, releasing greater quantities of far more volatile

ruthenium oxides, which may then reach the containment. This

explains why ruthenium is of particular concern when studying

accidents involving air ingress in the reactor vessel (the most

oxidizing accident conditions).

Two predominant scenarios resulting in fuel-air contact have

been identified (Figure 1). The first scenario involves accidental

draining of the reactor pool during refueling, with simultaneous

core dewatering [Powers et al., 1994]. The second corresponds

to the core meltdown phase that follows failure of the vessel

bottom due to corium (molten core materials), where gas (air)

circulates between the vessel pit, the reactor vessel, and the

RCS break [Seropian, 2003], [Freydier et al., 2006].

In the second configuration, the ruthenium released from the

fuel is transported through the RCS, characterized by a strong

thermal gradient, before reaching the containment. Recent

experimental studies indicate that a fraction of the ruthenium

is not trapped in the coolant system; the rate of trapping, or

retention, varies according to the species transported and the

thermal gradient involved, which in turn depends on where the

break is located. Ruthenium can take gaseous forms, such as

ruthenium trioxide (RuO3(g)) and ruthenium tetroxide (RuO4(g)),

along with condensed forms such as RuO2. It may also be

contained in mixed aerosols (e.g. Cs2RuO4).

Other accident types involving fuel-air contact – mainly spent

fuel handling and transport accidents or accidental draining of

a spent fuel storage pool – can result in Ru metal oxidation

Christian Mun, Laurent Cantrel Corium and Radioelements Transfer Research Laboratory

Study oF RutheNIum chemIStRy in the containment building under severe accident conditions

2. 2

(1) Level 1 probabilistic safety assessments (PSAs) conducted by IRSN on 900 MWe PWRs (PSA1 900) estimate the probability of occurrence for core meltdown acci-dents to be 10-5/year.reactor. Although the probability is very unlikely, core meltdown accidents would have significant radiological consequences, justifying in-depth studies on postulated scenarios and their progression.


2. 2

(in the presence of oxygen), depending on the temperature

reached.

Finally, ruthenium release was also detected during high-level

waste vitrification in a spent fuel reprocessing plant. RuO4(g)

formation during the process is postulated (hEPA filters(2) do

not retain RuO4(g)).

Purpose

Ruthenium is a fission product with high radiotoxicity, mainly

due to the isotopes 106Ru (T1/2=369 d) and 103Ru (T1/2=39.3 d).

This makes ruthenium a significant radiocontaminant in the

short and medium terms (as recognized by a French decree

published in 2003). If Ru particles were accidentally dissemi-

nated in the environment, their high specific activity could

lead to considerable external irradiation [Pöllänen, 1997].

Moreover, the risk of internal exposure cannot be overlooked,

since volatile ruthenium species (typically RuO4) could be

inhaled.

Significant amounts of ruthenium are formed during nuclear

reactor operation (essentially by direct fission of 235U and 239Pu). These amounts increase in proportion to fuel burn-up.

In addition, ruthenium content is higher in MOX fuel than in

conventional UO2 fuel. Gradual adoption of MOX fuel and the

trend toward higher burn-up rates could increase the amount

of ruthenium formed over the fuel's lifetime.

(2) High-efficiency particulate air filter.

Scientific approach

Ruthenium behavior, poorly modeled at present, is the focus

of an R&D program included in the European SARNET [2007]

excellence network. While IRSN deals with ruthenium behavior

in containment, other network partners (VTT, AEKI, CEA, etc .)

are conducting experimental studies on ruthenium release

and transport in the reactor coolant system. The ultimate

goal of all these programs is to supplement the experimental

knowledge base needed to develop and validate models for

the ASTEC integral code [Van Dorsselaere et al., 2005].

Regarding the IRSN program, a literature review revealed a

lack of quantified data on RuO4 gas phase stability and on

behavior of the oxides RuO2(c) and RuO4(g) in radiolytic

conditions. It also highlighted uncertainties, or even contradic-

tions, concerning interactions between the gaseous tetroxide

and containment surfaces (316L stainless steel or epoxy

paints). This absence of relevant data in the literature [Mun,

2007] led researchers to conduct an in-depth experimental

and theoretical study of ruthenium chemistry in severe acci-

dent containment conditions, focusing particularly on RuO2

and RuO4. Study parameters included a temperature range of

40-140°C, moist or dry atmosphere, and more or less oxidizing

conditions. With regard to RuO4(g) reactivity with steel sur-

faces, a number of papers have been published and several

hypotheses advanced concerning reduction of RuO4 deposited

on steel. Questions remain, however, as to the exact nature

of such deposits. There is also a total lack of information on

interaction between ruthenium tetroxide and epoxy paint.

And yet this is a key point, given the very large number of

painted surfaces in the containment.

Vessel breachAir

Degraded rods

Break

Air ingress following failureof vessel bottom

Air ingress following an accident causing dewatering of reactor vessel

Figure 1 Severe accident scenarios with air ingress in the reactor vessel.

Accidents in nuclear facilities


2. 2

Ruthenium deposit surface characterization

X-ray photoelectron spectroscopy (XPS) was used to analyze

the interactions between RuO4 deposits and the two PWR(3)

containment substrates. The findings led to the following

conclusions:

there is strictly no difference in the types of ruthenium

species found in deposits on epoxy paint and stainless

steel;

there are no chemical bonds linking the deposited ruthenium

to the paint polymer or the iron oxides; in other words, no

chemical reaction takes place at the deposit surface.

The XPS spectra also established that the species deposited

on the two containment substrates were similar to those

detected in the commercially available reference sample of

hydrated ruthenium dioxide. Analysis of the Ru3d and O1s

orbitals indicated that hydroxylated forms of Ru(IV), e.g .

RuO(Oh)2, made up most of the Ru deposits (at least at the

outermost surface, i.e. ≈10 nm). They are the only species

whose presence can be explained in both the commercial

hydrated Ru dioxide reference powder and the experimental

samples [Mun, Ehrhardt et al., 2007]. These results, coupled

with the RuO4(g) stability results, led researchers to conclude

that ruthenium tetroxide decomposition is a direct gas phase

process, followed by condensation of the reaction products

on painted surfaces, rather than an adsorption process. Using

XPS analysis to elaborate on the rare indications in the litera-

ture, researchers were able to propose an overview of RuO4(g)

decomposition reactions (Table 1).

Ru deposit oxidation study

The stability of containment ruthenium deposits in accident

conditions, i.e. in a partially oxidizing environment, was inves-

tigated using two approaches:

tests without irradiation, using an ozone generator to

determine Ru deposit oxidation kinetics constants under the

effect of ozone;

radiolytic tests conducted in an irradiation facility (EPICUR,

an ICPE(4) delivering a dose rate of around 4 kGy/h, Figure 2),

aimed at realistically reproducing physicochemical contain-

ment conditions during an accident, particularly the inventory

of radiolysis products (Oh•, O3, etc .).

(3) Pressurized water reactor.

(4) ICPE : Facilities classified for environmental protection.

In addition, the various reactions between ruthenium dioxide

and air radiolysis products suggest that deposits of RuO2, for

example, could undergo significant oxidation, leading to

gaseous tetroxide formation, with partial pressures potentially

reaching 10-7 to 10-5 bar [Mun et al., 2006]. Oxidation of

ruthenium species in the containment sump water is also

conceivable, to the extent that radiolysis, as induced by

suspended or dissolved fission products, can transform the

water into an oxidizing medium. The sump would then con-

stitute a potential source of volatile ruthenium (typically

RuO4).

IRSN aims to gather enough information to be able to assess

releases of RuO4(g) to the environment. Potential release

pathways include natural containment leakage, pathways

related to the French procedure of controlled and filtered

containment venting (known as "U5"), and the ground in the

case of basemat melt-through.

Study outcomes

Study of RuO4(g) "stability"

Although gaseous ruthenium tetroxide is often described as

"unstable", its stability must be evaluated at a severe accident

timescale, i.e. during the first 24 hours and thereafter. For

this purpose, a reliable and reproducible method of generating

pure ruthenium tetroxide crystals was developed, since this

compound is not available commercially.

Experimental results show that RuO4(g) decomposition is

slower than would be expected, based on the few indications

given in the literature. Thus, in conditions representative of

containment during a severe accident, i.e. at 90°C and in the

presence of steam, half-life time for the gaseous tetroxide is

around five hours.

Furthermore, decomposition does in fact appear to follow a

first-order rate law relative to the RuO4 concentration, as

predicted by certain authors [Ortins de Bettencourt et al.,

1969; Debray et al., 1888], although their results were obtained

under very different conditions. Regarding substrate interac-

tions, the results contradict other findings in the literature,

showing that RuO4(g) has no special affinity with ferrous

substrates or epoxy paints. In fact, the type of substrate has

no influence on gaseous tetroxide decomposition kinetics

[Mun, Cantrel et al., 2007].

Finally, the tetroxide decomposition reaction is accelerated by

the presence of steam and deposits of ruthenium oxides (RuO2

or similar compounds), which act as catalysts.


2. 2

n(Ru dep): quantity of ruthenium forming the deposit (mol)

V: volume (l)

[O3]: ozone concentration (mol.l-1)

The oxidation reaction is a first-order reaction with respect

to [O3] and [h2O].

Based on the rate laws established during the study without

irradiation, but nonetheless in the presence of radiolysis

products (from O3) (RuO4(g) decomposition and oxidation of

the deposits forming RuO4(g)), the Ru fractions revolatilized

under irradiation were calculated, then compared with the

experimental irradiation results obtained in EPICUR. The

calculated fractions are under-estimated by one order of

magnitude. Therefore, oxidation is enhanced under γ-radiolysis,

The first qualitative study, involving the ozone generator,

revealed revolatilization of ruthenium oxide deposits in the

40°C-90°C temperature range, in both dry and moist air.

Revolatilization was induced by the oxidizing effect of O3 on

active Ru deposit sites, producing RuO4(g).

The same oxidation reaction was also detected in radiolysis

tests, under the same temperature and humidity conditions.

Thus, it was experimentally shown that temperature and

humidity represent two key factors, whether or not ionizing

radiation is present. More specifically, increasing these two

parameters clearly favors the oxidation reaction. The strong

effect of an increased humidity rate is assumed to be attribut-

able to the hydroxyl radical (Oh•), an extremely powerful

oxidant with one electron. Based on the ozonation tests, an

ox idat ion ra te l aw fo r the ru then ium depos i t s was

proposed:

[ ] ( ) ][O V

n O)X(H k k

dt

RuOd3

)dep(Ru

2OHO

(g)4

23

+=

Where:

kO3 and kh2O: oxidation kinetics constants related to the action

of O3 and h2O(g) (l.mol-1.s-1)

X(h2O): molar fraction of steam

Reactions Steps Speed

RuO4(g) → RuO3(g) + 0,5 O2

RuO4(g) + H2O(g) → H2RuO5Initiation

Slow

Fast

RuO3(g) → 0.5 Ru2O5 + 0.25 O2

H2RuO5 → 0.5 Ru2O5,2H2O + 0.75 O2First reduction

+VI → +V Slow

+VIII → +V Medium

RuO3(g) + RuO2 → Ru2O5 Catalytic effect of RuO2 Fast

Ru2O5 → 2 RuO2 + 0.5 O2

Ru2O5,2H2O + 2 H2O → 2 (RuO2,2H2O) + 0.5 O2Final reduction

+V → +IV ?

+V → +IV ?

RuO2 + H2O(g)amb. → RuO(OH)2

RuO2,2H2O + H2O(g)amb. → RuO(OH)2 + 2H2O

Surface hydroxylationby steam in the environment

?

Table 1 Analysis of RuO4(g) decomposition, with and without steam(5).

Figure 2 View of ePICUR facility (IRSN/DPAM/SeReA, Cadarache center).

(5) This study refers to two distinct "types" of water vapor: the first, H2O(g), present in the system during stability tests with humidity; and the second, H2O(g) amb., produced after exposure of the samples (ruthenium depots on steel or painted substrates) to ambient air (following stability tests).

Glove box (used only during irradiation tests on iodine samples)

Irradiator(60Co sources)

Irradiation cell



2. 2

ruthenium releases;

a two-fold increase in the corium mass pouring from the

reactor vessel into the reactor pit results in a clear reduction of

release by a factor of 6. While at first surprising, this outcome

is explained by the indirect effects of containment temperature

and pressure. With the doubled corium mass, the mean contain-

ment temperature increases (by around 30 K), with a correlative

increase in RuO4(g) decomposition. Additionally in this case,

as compared to the ozonation tests; this is explained by the

predominant role of the O. and/or Oh. radicals. During the

γ-radiation tests, additional quantities of these radicals were

produced directly by air and steam radiolysis.

Although this research focused primarily on gas phase ruthe-

nium chemistry, aqueous solutions of ruthenium (in the form

of perruthenate: RuO4 -) were irradiated in a few exploratory

tests, to determine whether volatile ruthenium tetroxide could

be formed from an aqueous phase subjected to radiolysis.

Initial results revealed formation of RuO4(g). Depending on

the experimental conditions, the revolatilized Ru fractions

can reach values of around 12%. however, at this stage, the

influence of key parameters such as ph, temperature, and the

integrated γ dose has yet to be determined and quantified.

Investigations will be pursued by IRSN in 2008.

First evaluation of Ru releases

Based on the experimental results detailed above, a kinetic model

of RuO4(g) decomposition and volatile tetroxide formation via

Ru oxide deposit oxidation was proposed and implemented in

ASTEC [Van Dorsselaere et al., 2005], the benchmark European

code for severe accident simulation. A "reactor case" simulation

for a 900 MWe PWR was then carried out using realistic boundary

conditions (thermohydraulic parameters, dose rates, etc.). The

simulated accident scenario was an h2 sequence(6), involving

reactor vessel failure followed by air ingress, with the ruthenium

fraction released in the containment estimated at 10%.

Several simulations were run to study the sensitivity of certain

parameters (dose rate, amount of corium). The results of one of

the "reactor" simulations(7) are shown to illustrate this approach.

Figures 3 to 5 respectively show the mass of ruthenium in the

form of RuO4(g) inside containment, the massof ruthenium

deposited on the inner walls, and the mass of ruthenium released

to the environment as RuO4(g).

Gaseous ruthenium releases in this simulation were on the order

of a few grams (Figure 5); similar values were obtained in the

other simulations, despite differing initial conditions. The sen-

sitivity study highlighted the following trends:

dose rate plays an important role in volatile Ru formation,

and decreasing it by a factor of 2 results in a two-fold drop in

Figure 3 Mass of Ru in containment as RuO4(g).

00

1

1.105 2.105 3.105 4.105 5.105

2

3

4

5

6Ru (kg)

1 Ru

t (s)

H2 sequence

Gaseous ruthenium tetroxide (RuO4)

1

1

1

1

1

1

1

1

1

1

1 1 1 1 1 1

1

1

1 Ru

00

1.105 2.105 3.105 4.105 5.105

20

15

10

5

Ru (kg)

t (s)

H2 sequence

Ru deposited on containment walls

1

1

1

1

1

1 1 1 1 1 1

1 Ru

010-3

1.105 2.105 3.105 4.105 5.105

102

101

100

10-1

10-2

Ru (g)

t (s)

Ruthenium released to environment

Environment

1

1

1

1 1

1

1

11 1 1

(6) The H2 sequence results in the combined loss of the normal steam generator (SG) feedwater system and the emergency SG feedwater system.

(7) Simulation run with the following initial conditions: a dose rate of 10 kGy.h-1 prior to the U5 procedure (controlled and filtered containment venting), and a corium weight of 82 tons. The U5 procedure is implemented at 2.5 days. Figure 4 Mass of Ru deposited on walls.

00

1

1.105 2.105 3.105 4.105 5.105

2

3

4

5

6Ru (kg)

1 Ru

t (s)

H2 sequence


1

1

1

1

1

1

1

1

1

1

1 1 1 1 1 1

1

1

1 Ru

00

1.105 2.105 3.105 4.105 5.105

20

15

10

5

Ru (kg)

t (s)

H2 sequence


1

1

1

1

1

1 1 1 1 1 1

1 Ru

010-3

1.105 2.105 3.105 4.105 5.105

102

101

100

10-1

10-2

Ru (g)

t (s)


Environment

1

1

1

1 1

1

1

11 1 1


2. 2


With regard to gaseous phase ruthenium chemistry, this

research identified the key reaction parameters and

opened the way for model development. Applied to the

"reactor case", the models indicate non-negligible ruthe-

nium release, the radiological impact of which remains to

be quantified. These models are currently being integrated

in the MER(8) for PSA2(9) studies.

Research on ruthenium behavior will be pursued at IRSN

and in the SARNET [2007] program in order to reduce

experimental uncertainty associated with ruthenium gas-

eous phase chemistry, and to broaden knowledge of

aqueous phase ruthenium chemistry in the presence of γ

radiation. These objectives coincide with the priorities set

by the SARNET committee of experts.

(8) Release assessment model.

(9) Level 2 probabilistic safety assessment.

steam partial pressure is less than 0.2 bar on average, which also

favors a decrease in the gaseous tetroxide production rate.

Figure 5 Mass of Ru released to environment as RuO4(g).

00

1

1.105 2.105 3.105 4.105 5.105

2

3

4

5

6Ru (kg)

1 Ru

t (s)

H2 sequence


1

1

1

1

1

1

1

1

1

1

1 1 1 1 1 1

1

1

1 Ru

00

1.105 2.105 3.105 4.105 5.105

20

15

10

5

Ru (kg)

t (s)

H2 sequence


1

1

1

1

1

1 1 1 1 1 1

1 Ru

010-3

1.105 2.105 3.105 4.105 5.105

102

101

100

10-1

10-2

Ru (g)

t (s)


Environment

1

1

1

1 1

1

1

11 1 1



2. 2

References

h. Debray, A. Joly, Compte rendu des séances de l’académie des sciences, 1888 (106) : p. 328-333.

Décret relatif à la protection des travailleurs contre les dangers des rayonnements ionisants, n° 2003-296, paru au J.O. n° 78 du 02/04/03. 2003.

G. Ducros, Y. Pontillon, P.P. Malgouyres, P. Taylor, Y. Dutheillet. Ruthenium release at high temperature from irradiated PWR fuels in various oxidising conditions; main findings from the VERCORS program. in Nuclear Energy for New Europe 2005. Bled (Slovenia).

P. Freydier, J.L. Rousset, Evaluation of Air Ingress in the Reactor Vessel with the SATURNE Code. SARNET-ST-P19. EdF n° HI-83/05/006/A. 2006.

J. Libmann, Éléments de sûreté nucléaire. Les éditions de physique IPSN. 1996.

C. Mun, Étude du comportement du produit de fission ruthénium dans l’enceinte de confinement d’un réacteur nucléaire, en cas d’accident grave. Thèse univ. Paris XI. 2007.

C. Mun, L. Cantrel, C. Madic, A Review of Literature on Ruthenium Behaviour in Nuclear Power Plant Severe Accidents. Nuclear Technology, 2006. 156(3): p. 332-346.

C. Mun, L. Cantrel, C. Madic, Study of RuO4 Decomposition in Dry and Moist Air. Radiochimica Acta, 2007. 95 (11): p. 643-656

C. Mun, J.J. Ehrhardt, J. Lambert, C. Madic, XPS Investigations of Ruthenium Deposited onto Representative Inner Surfaces of Nuclear Reactor Containment Buildings. Applied Surface Science, 2007. 253(18): p. 7613-7621.

A. Ortins de Bettencourt, A. Jouan, Volatilité du ruthénium au cours des opérations de vitrification des produits de fission (2e partie). Rapport CEA-R-3663 (2). 1969: CEN de Fontenay-aux-Roses.

R. Pöllänen, Highly radioactive ruthenium particles released from Chernobyl accident: Particles characteristics and radiological hazard. Radiat. Protec. Dos., 1997 (71): p. 23-32.

D. Powers, L.N. Kmethyk, R.C. Schmidt, A review of the technical Issues of air ingression during severe reactor accidents. NUREG/CR-6218. 1994.

C. Seropian, Analysis of the potential for in-vessel air ingress during a severe accident in a PWR 900 MWe. Note technique IRSN/DPAM/SEMIC/LEPF. 03/01/2003.

J.P. Van Dorsselaere, J.C. Micaelli, h.J. Allelein. ASTEC and SARNET. Integrating severe accident research in Europe. in ICAPP’05. 2005 (15-19 mai). Séoul (Corée).http://sarnet.grs.de/default.aspx 2007.



(1) USNRC, AECL, PSI, Suez/Tractebel, the European Commission.

(2) Process and Materials Science and Engineering Laboratory, CNRS UMR 5266, INPG, UJF, Saint-Martin d’Hères.

(3) ASTEC: Accident Source Term Evaluation Code.

Marie-Noëlle Ohnet, Didier Jacquemain

Separate Effect Test Programs Laboratory

Benoît Durville, Christophe Marquie

Experimental Equipment and Instrumentation Engineering Laboratory

Phebus FP experiments on PWR core

meltdown demonstrated that reactor

accident simulation codes do not take

into account the fact that a significant

port ion of iodine emissions in the

containment are released in a gaseous

state. This may entail a greater risk of

iodine release to the environment in the

event of a reactor accident.

In partnership with CEA, EDF, and other

international organizations(1), IRSN has

launched a research program aimed at

generating experimental data on ther-

modynamic and kinetic constants for

chemical reactions between the main

elements in the reactor coolant system

that could influence the volatile iodine

fraction fuel is transported

during an accident.

The experimental facility developed

for ChIP (a separate-effect test program

on reactor coo lant system iod ine

chemistry in accident conditions) was

developed and coupled to a h igh-

temperature mass spectrometer provided

by CNRS/SIMaP-Grenoble(2).

More than 20 subcontractors contributed

to the development of this complex

system, featuring over 600 miniaturized

components – a success crowning two

years of fruitful cooperation between

engineer ing pro ject managers and

research scientists.

The first thermokinetic studies in the

reactor began in April 2008 and will

continue into 2011. Experimental results

will be used to validate the ASTEC safety

code(3) developed by the Institute, while

aiming to reduce uncertainty in iodine

source term assessments.

KIcK-oFF FoR chIP, an experimental program in Grenoble2.3

Figure 1 Thermokinetic reactor. Figure 2 Instrumented furnace column. (reactor heating)


2. 4

Research topics

UF6 behavior in an accidental leak situation is complex because

it involves several interacting physicochemical phenomena

(Figure 1). For example, liquid UF6 (the worst-case scenario for

accident situations and the most complex in terms of the

phenomena involved) is an unstable compound in ambient

pressure and temperature conditions. Accidental containment

loss on a container of liquid UF6 stored in hot, pressurized

conditions leads to rapid decompression of UF6. This results in

equivalent proportions of UF6 in solid and vapor form, which

behave very differently. Depending on the ambient thermody-

namic conditions, the solid may sublimate and the vapor recrys-

tallize. UF6 vapor reacts violently with moisture in the air,

producing hF vapor and UO2F2 aerosols. The aerosols may then

be deposited in the location where leakage occurred, and may

Abdalkarim Abbas, Cyril HuetNuclear and Radiological Emergency Management Unit

A common interest program on UF6 behavior in accidental release contexts was conducted from 1998 to

2006 to achieve a better understanding of the consequences of environmental dispersion of UF6 and its

hydrolysis products. This program, conducted by IRSN, was co-funded by three operators: AReVA NC,

eurodif, and FbFC Romans.

Uranium hexafluoride (UF6) is the most volatile uranium-bearing compound. It is used in the front end of

the nuclear fuel cycle, during the conversion, enrichment, and fabrication stages.

UF6 reacts violently with water, particularly steam, producing solid uranyl fluoride (UO2F2) and gaseous

hydrofluoric acid (hF). The environmental consequences of a UF6 accident depend primarily on the chem-

ical toxicity of UF6 and hF, in addition to the radiotoxicity of uranium.

Accidents involving UF6 release represent a significant risk, which nuclear operators take into consid-

eration when establishing safety practices and emergency plans for their facilities. but UF6 behavior in

an accidental release situation is poorly understood. The various assumptions currently used to quantify

UF6 release are associated with considerable uncertainty, which is particularly unsatisfactory for safety

assessment purposes. This observation provided sufficient motivation to justify the UF6 common interest

program.

The first step in this research program was to review current knowledge on UF6 properties. Researchers

focused on data relevant to accident situations involving UF6 release. Conclusions from this review were

used to define the various research topics investigated subsequently by the common interest program.

uF6 behAvIoR IN AN AccIdeNtAl ReleASe coNtext Studies and experiments to quantify accidental UF6 release in front-end fuel cycle facilities


2. 4

not sublimate. The study aimed to acquire the knowledge and

tools necessary to evaluate solid UF6 sublimation, for conditions

representative of the accident situations taken into consider-

ation in emergency plans.

Approach

The literature review yielded neither the data nor models

needed to assess solid UF6 sublimation kinetics and quantities

in an accident situation. Consequently, researchers decided to

develop a sublimation model and validate it experimentally.

Sublimation model

The model was developed by adapting Acacia [Ducruet et al.]

to UF6. Acacia is a model developed by IRSN to study evapora-

tion/condensation-induced changes in the size of a free-falling

water droplet in a facility with variable relative humidity,

representative of nuclear reactor containments. The resulting

model is based on calculating heat and material exchanges

between solid UF6 and the atmosphere, occurring via a solid/

vapor interface. Transfer coefficients were determined from the

system's thermodynamic and air flow conditions (solid – inter-

face – atmosphere). Phase transition kinetics were based on a

series of equilibrium states (quasi-static model). Developing

this model also helped identify the various parameters influenc-

ing the sublimation process, in particular the temperature inside

be retained by ventilation system filters in the facility where

the accident occurred. The filters may be damaged, however,

due to the acidity of the hF gas.

To make the most realistic impact assessment, the amounts of

uranium and hF released to the environment must be deter-

mined, taking into account all the phenomena mentioned above.

The common interest program comprised tests and studies

aimed at expanding knowledge of the various phenomena

involved in the accident process, in order to improve release

assessment. The research focused on phase transitions, par-

ticularly sublimation, and on dispersion of UF6 (a heavy, very

reactive gas) inside an enclosure, UO2F2 aerosol deposition,

and the strength of filters exposed to hF.

Results of the uF6 common interest program

Evaluating sublimation

Scope

Phase transitions are central to assessing the consequences of

an accident involving UF6. A key issue is the fate of solid UF6,

which forms after accidental leakage of liquid UF6.

According to current assumptions, solid UF6 does not influence

accident consequences once deposited and, in particular, does

UF6 Liquid

UF6 gas

UF6 gas + H20

UF6 gas + UO2F2 solid + HF gas

Heat

Moist air

UF6 solid + UO2F2 solid

UO2F2 solid retained on HEPA filters, filter efficiency impaired after exposure

to HF

UF6 phase distribution UF6 liquid decompression

Sublimation / recrystallization

UF6 gas hydrolysis

UF6 gas dispersion

UF6 release

UF6 gas, HF gas, and UO2F2 solid released to environment via ventilation system

or direct leakage

UO2F2 solid deposition

Figure 1 Process of accidental UF6 release in a ventilated enclosure.



2. 4

conditions, gravitational effects can lead to stratification, in

turn producing large concentration gradients. Gravitational

effects can also strongly influence hydrolysis, which depends

on: the quality of the mixture between gaseous UF6 and moist

air; sublimation, a function of the concentration of near-solid

UF6; and UO2F2 aerosol deposition in the enclosure, which

depends on the height of aerosol formation.

This study aimed to improve characterization of UF6 vapor

dispersion in a ventilated enclosure by integrating the gravi-

tational effects of this high-density gas.

Approach

Full-scale tests are usually required for experimental charac-

terization of heavy gas dispersion. The nature of gaseous UF6,

however, rules out large-scale tests, given the necessary mea-

sures that must be taken to prevent the risk of environmental

release. Researchers thus opted to model gaseous UF6 behavior

using a multidimensional computer tool. The first step was to

select one of the commercially available numerical tools and

assess its ability to simulate gravitational effects, which were

characterized experimentally using a chemically inert simulant.

The physicochemical properties of UF6 were then integrated

in the numerical model so that researchers could conduct a

multidimensional simulation campaign for conditions repre-

sentative of accidents involving UF6 release.

Selection and experimental validation

of a multidimensional tool

Although several commercially available multidimensional tools

were evaluated, only CFX-5 [Ansys, 2003] met the requirements.

The validation tests were conducted in two enclosures with

different volumes (36 m3 and 1500 m3), at the IRSN Saclay

site. Sulfur hexafluoride (SF6) was used, given its high density

(d = 5 g/cm3). The tests examined 19 different configurations:

15 for the 36-m3 enclosure and 4 for the 1500-m3 enclosure.

A large majority of these configurations showed strong strati-

fication of the gas, with high floor-level concentrations as soon

as injection began. Time-dependent changes in concentration

levels were mainly linked to gas injection characteristics. high

injection rates were experimentally observed to favor dispersion

of the gas. Although injection remained the predominant dis-

persal mechanism, mechanical ventilation had a more pro-

nounced effect in the 1500-m3 enclosure. hence, SF6

stratification was less pronounced for spaces with greater

volume.

The two experimental enclosures were modeled using CFX-5,

following a sensitivity study on the selected mesh. All tests

were then simulated, and agreement with experimental results

the enclosure, UF6 partial pressure, the size and shape of solid

UF6, and air flow conditions around the solid.

Experimental validation

Validation tests were conducted with simulants, such as dry

ice, using the Bise test bench [Gelain, 2004], located at IRSN's

Saclay site. This test bench creates a perfectly controlled air

flow around a sample. CO2 concentration in the air circulating

around the sample was measured to evaluate the sublimation

rate for dry ice. Results showed satisfactory agreement between

measured sublimation rates and laws used in the model, thereby

validating the model. The tests were also used to reproduce

the influence of various model parameters, such as sample size

and air flow rate around the sample. Iodine sublimation tests

rounded out the study, confirming the influence of partial

pressure on sublimation.

Computation results

Once validated, the sublimation model was used to quantify

the degree of sublimation in situations representative of acci-

dents considered in emergency plans. In all cases, computation

shows that the fraction sublimated in a few hours is consider-

able, and in certain cases is even complete, thereby confirming

that sublimation cannot be ignored in assessing the conse-

quences of an accident involving UF6.

A sensitivity study on the various influential parameters quanti-

fied their impact on the sublimation rate. Some of these param-

eters could change considerably during accident progression,

such as the gaseous UF6 concentration, which depends on

dispersion of the gas in the building where the accident has

taken place. The common interest program examined this

specific phenomenon, discussed below in the Study of gaseous

UF6 dispersion in a ventilated enclosure below.

Assessment tool

The sublimation model was implemented in SUBLI_UF6. This

tool was developed to take into account all phenomena affecting

the influential parameters in the model, such as gaseous UF6

dispersion in the enclosure and hydrolysis.

Study of gaseous UF6 dispersion in a ventilated

enclosure

Scope

According to the current assumption, gaseous UF6 distribution

is homogeneous throughout the enclosure. Gaseous UF6 has a

high density (d = 12 g/cm3) and could be emitted at very high

concentrations in the event of accidental release. Under these


2. 4

available water vapor. These results are related to the strong

reactivity of UF6 and the air circulation ensured by the ventila-

tion system in the enclosure.

The simulations confirm that assuming homogeneous gas

distribution in the enclosure is acceptable. This assumption has

thus been integrated as an operational rule in SUBLI_UF6.

Effect on sublimation

The concentration (or partial pressure) of UF6 as it nears the

solid phase is a parameter that influences sublimation. The

higher its value, the lower the sublimation rate (which falls to

zero when saturating vapor pressure is reached). Simulation

results show that UF6 concentrations near the ground are very

high during the injection phase, and then decrease rapidly.

Figure 3 illustrates that maximum concentration (7%) is reached

at the end of injection (120 s), followed by a rapid drop (under

4% at 180 s).

The test configurations resulted in concentrations that could

reach nearly 30% at the end of injection, before decreasing

rapidly. The operational criterion for integrating the effect of

UF6 stratification on sublimation assumes that no sublimation

occurs in the enclosure during the injection phase and part of

the drop-off phase, with injection lasting up to 20 minutes in

the accident situations considered for emergency plans. This

criterion remains to be integrated in SUBLI_UF6.

Effect on UO2F2 aerosol deposition

The height at which UO2F2 aerosols are formed is a parameter

required to evaluate deposits in the enclosure. By default, this height

is set to the enclosure height, a conservative assumption for release

was generally satisfactory, as shown by the comparative example

in Figure 2.

More specifically, the simulated injection ranges were in good

agreement with test results; SF6 stratification and the concen-

tration levels reached at various measurement points were

clearly reproduced in the simulations, even though the code

had a slight tendency to over-estimate them. In addition, the

time-dependent changes in SF6 concentration were not always

accurately simulated; the decrease in concentration level was

generally faster in tests than in simulations. Based on these

results, simulation of gravitational effects by CFX-5 was con-

sidered satisfactory enough to pursue this approach [Bouilloux

et al., 2006].

Applying the model to UF6

Researchers adapted the numerical model to UF6 by integrating

thermodynamic and physicochemical properties of the com-

pound, in particular the exothermic hydrolysis reaction. A

simulation campaign was defined to characterize the effects

of gaseous UF6 stratification on UF6 hydrolysis as well as on

solid UF6 sublimation and UO2F2 formation height, a data item

required to assess deposition. Simulation parameters included

orientation of the injection stream (vertical, horizontal), total

amount of UF6 emitted in the enclosure, and air renewal rate

in the enclosure.

Results and assessment tool applications

Effect on hydrolysis

Simulation results show that regardless of configuration, the

simulated hydrolyzed fraction always exceeds 80% of the

maximum hydrolysable fraction, determined by considering all

00 500 1 000 1 500 2 000

t(s)

00 500 1 000 1 500 2 000

t(s)

B

M

H

5 000

10 000

15 000

20 000[SF

6]ppm

5 000

10 000

15 000

20 000[SF

6]ppm

Figure 2 Changes in SF6 concentrations calculated and measured experimentally at different heights in the 36-m3enclosure. h, M, and b respectively represent measurement heights of 2.50 m, 1.50 m, and 0.55 m, for an enclosure height of 3 m.



2. 4

UO2F2 aerosol formation height must be known to use the

deposition charts presented in the following section on UO2F2

behavior.

Characterization of UO2F2 behavior

Scope

hydrolysis of gaseous UF6 produces l’UO2F2 aerosols, which

may be deposited in the building where the accident has

occurred. Until now, the assessment of accident consequences

involving UF6 has not taken this deposition into account, thereby

ensuring conservative assessments. Data on UO2F2 character-

istics, including particle size, are, however, available in the

literature, and deposition computer codes exist as well .

Researchers thus set out to examine the conservative quality

of this assumption.

Approach

The approach adopted was based on three objectives: to quan-

tify deposition for conditions representative of accidents con-

sidered in emergency plans, using data from the literature and

computer codes at IRSN; to identify the predominant deposition

assessment. The simulations indicate that beyond a certain height,

the UF6 concentration becomes very low; elsewhere it remains

relatively homogeneous within the horizontal planes and shows a

strong correlation to the UF6 injection characteristics (Figure 4).

For vertical injection, the height at which UF6 concentration starts

to decrease considerably is close to the injection height (below

enclosure height). The Baines semi-empirical law can be used to

evaluate injection height, which could in turn be used to estimate

UO2F2 deposition.

For cases of non-vertical injection, the simulations show that

the height of near-zero UF6 concentration is lower than for

vertical injection, and close to the exhaust outlet height.

Consequently, depending on the case, several UO2F2 formation

heights can be used:

enclosure height, for configurations where the atmosphere

is relatively homogeneous, or for a conservative simulation of

release;

injection height, for vertical injections, estimated using the

Baines law;

exhaust outlet height, for other cases.

Figure 3 Changes in concentration near the ground when UF6 is injected for 120 s.

0.043

0.043

0.043

0.042

0.042

0.042

0.042

0.041

0.041

0.041

UF6 mass fraction

180 s

(Volume 1)

0.008

0.007

0.007

0.006

0.005

0.005

0.004

0.003

0.002

0.002

UF6 mass fraction

600 s

(Volume 1)

0.052

0.047

0.042

0.038

0.033

0.028

0.024

0.019

0.014

0.00960 s

UF6 mass fraction(Volume 1)

0.071

0.068

0.065

0.062

0.058

0.055

0.052

0.048

0.045

0.042120 s

UF6 mass fraction(Volume 1)


2. 4

mechanisms; and if necessary, to propose a simple tool to

evaluate the deposit inside an enclosure.

Evaluation of UO2F2 deposit

UO2F2 is a very hygroscopic aerosol whose characteristics

depend to a great extent on how the aerosol was formed (UF6

hydrolysis). These characteristics could potentially change as

the aerosol circulates within the facility or the environment,

for example, following a reaction between the aerosol and

humidity in the air. Nonetheless, after reviewing the literature,

researchers were able to define characteristics they considered

representative of the aerosol. More specifically, particle size is

well represented by a log-normal distribution. The mass median

aerodynamic diameter (MMAD) falls within a range of 1-10 µm,

which can be reduced to 2-8 µm if the most frequent values

in the literature are considered. Finally, the geometric standard

deviation for the log-normal distribution varies from 1.5 to 2

and the density of UO2F2 is around 4.

Based on these data, a study was conducted using AEROSOLS_B2

[Gauvain et al.], for conditions representative of accidents

considered in emergency plans. The main deposition phenomena

(gravitational settling, Brownian diffusion, thermophoresis, and

diffusiophoresis) and particle agglomeration phenomena (gravi-

tational, Brownian, and turbulent coagulation) were taken into

account.

The simulations indicate that settling is the predominant

deposition mechanism. Wall deposits by thermophoresis, dif-

fusiophoresis or Brownian diffusion are negligible, regardless

of the accident scenario considered.

The aerosol characteristics (MMAD and particle size distribu-

tion) are major parameters in determining what portion of the

aerosol will be deposited. For the characteristics considered,

this study shows that deposition can be very significant in

certain situations, and overlooking it amounts to a very penal-

izing or even unrealistic assumption. This pointed to the need

for a simple assessment tool.

Assessment tool

Using AEROSOLS_B2, researchers developed a tool for evaluating

UO2F2 deposition in an enclosure, only taking into consideration

deposition by settling. Results are presented in the form of

charts (curves and tables). Since this simple tool only considers

deposition by settling, it requires only a limited number of

parameters, i .e. aerosol characteristics (MMAD, standard

deviation, density) and the characteristics of the enclosure

where the accident occurred (height and air renewal rate).

Figure 4 UF6 concentration in the enclosure at the end of injection.

Figure 5 hePA filtering material exposed to hF.

UF6 mass fraction

0.833

0.714

0.595

0.476

0.357

0.238

0.119

0.000

UF6 mass fraction

1.000

0.857

0.714

0.571

0.429

0.286

0.143

0.000

UF6 mass fraction

1.000

0.857

0.714

0.571

0.429

0.286

0.143

0.000



2. 4

Assessment tool

A comprehensive tool for determining the condition of a filter

exposed to hF was not developed because researchers were

unable to determine the filter's limit strength. But they were

able to develop a graphic representation defining three distinct

ranges of filter strength:

filtering efficiency maintained (filter intact) (t < 15 min;

C < 400 ppm);

total loss of filtering efficiency (brittle filter) (t > 60 min;

C > 1000 ppm);

intermediate range where filter strength was not characterized

(15 min < t < 60 min; 400 ppm < C < 1000 ppm).

Summary of resultsThe UF6 common interest program compiled an inventory of

physicochemical accident phenomena, reviewed the literature

for each phenomenon, and in some cases, gained insight into

how these phenomena influence impact assessment. These

results were used to develop a set of "relatively" simple

tools.

The test results obtained in the UF6 common interest program

provided the basis for building the SUBLI_UF6 computer tool,

the charts to evaluate UO2F2 aerosol deposition in the building

where the accident occurred, and the graphic tool (based on

hF concentration and exposure time) that characterizes the

hEPA filter operational range.

All these tools were used to assess the consequences of several

accident scenarios, and the results were compared to values

obtained using the "old" assumptions. The common interest

program tools appear to provide more realistic assessments,

particularly in terms of kinetics, considering that release time

was previously a fixed value in most cases. Moreover, since

the new assessments are based on the same approach (a series

of physicochemical evaluations) regardless of the type of UF6

release (e.g. liquid or gaseous), there is better physical con-

sistency between results for the various types of accident,

which was not necessarily the case up until now, since previous

computing methods varied according to the operator and

postulated scenario. While these results generally confirm the

degree of risk associated with UF6 activities at the front end

of the fuel cycle, each accident situation can nonetheless be

assessed more realistically, potentially highlighting significant

differences, which may in turn lead individual facilities to

redefine their most severe scenarios.

Charts were produced for two aerosols (MMAD of 2 and 8 µm,

geometric standard deviation of 1.75, density of 4), for heights

and renewal rates characteristic of industrial facilities that use

UF6. The tool offers a total of 264 configurations.

Characterizing HEPA filter strength

Scope

high-efficiency particulate air (hEPA) filters consist of filtering

cells in a galvanized or stainless steel frame or structure that

holds the filtering material. hF could severely damage this

filtering material, composed of borosilicate glass fibers. The

study aimed to characterize the behavior of hEPA filters

exposed to hF, focusing on any changes in filter efficiency.

Approach

For the accident situations considered in the common interest

program, hEPA filters in facilities could be exposed to high

hF concentrations (potentially reaching 5x104 to 4.105 ppm),

including concentrations greater than 104 ppm for durations

varying from 45 minutes to 9 hours. The only existing data

[Del Fabro et al., 2002], identified by the pre-test literature

rev i ew, i nvo lves re l a t i ve ly l ow exposu re t imes and

concentrations (filters remain efficient after 15 minutes of

exposure to hF concentrations between 300 and 400 ppm).

These data are not representative of the postulated accident

conditions, which are far more severe. Researchers thus

collected experimental data that could be used to characterize

hEPA filter efficiency at hF concentrations and exposure times

representative of the accidents considered in emergency plans,

with the addit ional aim of assessing l imit strength, i f

possible.

Test results

The tests were performed by Comurhex, a company based in

Pierrelatte, France. Their experimental facility can provide

tested hF concentrations no lower than 1000 ppm. The tests

[Bouilloux et al., 2004] consisted of exposing the filtering

material to various hF concentrations and measuring the

resulting changes in filtering characteristics. The results showed

that exposing the filtering material to the facility's lowest hF

concentration led to total loss of hEPA filter efficiency; the

filtering material was destroyed in less than 60 minutes.

Figure 5 shows filtering material exposed to hF.


2. 4

on deposition. Setting up an experimental program to char-

acterize the properties of the UO2F2 aerosol is one possible

follow-up project to the common interest program.

Researchers were unable to determine the limit strength of

HEPA filters exposed to HF, given that the lowest available

HF concentration (1000 ppm) at the experimental facility was

too high. Future efforts may focus on characterizing filter

behavior across the entire concentration range. This, how-

ever, would require an experimental system capable of

obtaining lower HF concentrations (below 1000 ppm).

outlook

The tests characterizing heavy gas dispersion in a ventilated

enclosure revealed phenomena that have yet to be explained.

At the end of the injection phase, test results showed that a

mechanism that reduces gas levels accentuates the well-

modeled decrease in gas concentration linked to air renewal

in the enclosure. This translated as a faster dropping phase

in the tests as compared to the simulations. Additional stud-

ies are necessary to explain these unexpected results.

The UO2F2 study revealed gaps in the literature on the aero-

sol characteristics and demonstrated its very strong influence

References

ANSYS Company, CFX-5 Solver Theory Manual CFX Ltd., Oxfordshire, 2003.

L. Bouilloux, C. Prevost, L. Ricciardi, R. Sestier-Carlin, Modélisation de la dispersion d’un gaz lourd dans un local ventilé. Rapport scientifique et technique de l’IRSN 2006, p. 117–123.

L. Bouilloux, , E. de Vito, O. Norvez, Bilan de l’étude du comportement des filtres ThE en cas de rejet accidentel d’UF6. Note technique DSU/SERAC/LECEV/04-26, septembre 2004.

D. Ducruet, J. Vendel, Description détaillée d’Acacia : algorithme appliqué à Caraidas pour la capture de l’iode et des aérosols. Rapport d’étude SERAC/LPMC/98-17.

L. Del Fabro, J.-C. Laborde, C. huet, Réalisation d’une étude bibliographique et du dimensionnement d’essais dédiés au comportement de filtres ThE en cas de rejet accidentel d’UF6. Note technique DPEA/SECRI/02-082, septembre 2002.

J. Gauvain, G. Lhiaubet, ESCADRE mod 1.2, AEROSOLS_B2 Release 3.3, Aerosol behavior in containment, Reference document. Note technique SEMAR 98/57.

T. Gelain, Essais de sublimation de carboglace dans BISE. Rapport DSU/SERAC/LEMAC/04-10, mai 2004.


Introduction

What is a criticality accident?

The risk of criticality (an uncontrolled fission chain reaction) exists

in fuel cycle facilities where non-negligible amounts of fissile

material(1) are handled (typically more than a few hundred grams).

Nuclear power plants are therefore at risk, as are facilities for uranium

enrichment, nuclear fuel fabrication, post-irradiation fuel processing,

and nuclear waste management, in addition to fissile material

transport casks and certain research laboratories. The fissile materials

involved (uranium, plutonium, mixed uranium and plutonium, other

fissile actinides [heavy nuclei]), their physicochemical form (liquid,

solid, gas), and their conditions of use are all highly variable.

When the handled material contains enough fissile material to

initiate and sustain a fission chain reaction, the material is said to

have reached "criticality" or a "critical state". A nuclear reactor in

its normal operating state functions at criticality. Beneath this

threshold, systems are referred to as "subcritical". Facilities that

use fissile materials are designed so that subcriticality is maintained

in all situations, with a safety margin relative to the critical state.

Beyond criticality, in a "supercritical" state, the fission chain reaction

(which produces energy, neutrons, and gamma radiation), develops

at a pace that increases as the reaction goes further and further

beyond critical conditions(2). This is called a criticality accident.

Design and/or operating measures to avert criticality risk in

facilities are subject to analyses and studies conducted by

operators, as detailed in their safety reports. These measures,

which, in France, must comply with a Basic Safety Rule specific

to criticality risks (RFS 1.3.c), are also examined by IRSN

criticality specialists. Nevertheless, despite the measures

taken, criticality accidents cannot be totally excluded.

As part of its mission to provide technical support to public

authorities, IRSN conducts research focused on criticality

accidents to obtain the most relevant information given the

state of the art. The challenge is to preventively assess the

consequences of a criticality accident, with a view to ensuring

that measures taken in the event of an accident will effectively

mitigate the consequences.

The purpose of IRSN's research in this area is to develop and

sharpen the skills required to assess the consequences of a

criticality accident, in order to provide support in an emergency

situation, particularly for assessing radiological consequences

and response capability, and also to analyze the pertinence

of the number of fissions (or the "fission source term"), a

Luis Aguiar, Véronique RouyerCriticality Assessment, Study and Research Department

Matthieu Duluc, Xavier Knemp, Igor Le Bars Assessment Section for Criticality Risks and Accidents

Study oF PhySIcAl PheNomeNA and consequences associated with criticality accidents

2. 5

(1) Material containing nuclei said to be fissile, i.e. having a non-negligible probability of undergoing fission by interaction with neutrons, regardless of their energy level.

(2) Critical conditions define all the characteristics required (mass, geometry, physi- Critical conditions define all the characteristics required (mass, geometry, physi-Critical conditions define all the characteristics required (mass, geometry, physi-cochemical form, etc.) for a neutron-multiplying medium to reach the critical state.


2. 5

parameters, such as how far the situation has gone beyond

the critical state, and the supercriticality kinetics. For example,

if the time required to double the number of neutrons is

around half a second, the power produced can be multiplied

by one million in 10 seconds. By comparison, if doubling time

is on the order of one hundredth or one thousandth of a

second, power can be multiplied by one billion in respectively

three tenths or three hundredths of a second, reaching values

of 10 to 100 MW in a few fractions of a second (100 MW

represents around 3 ×1018 fissions per second, given that one

joule equals approximately 3.1 ×1010 fissions). An accident

results in an initial power spike, generally followed by other

excursions. It can be prolonged over time by power oscillations

of varying frequency and amplitude. Accident duration is also

highly variable, ranging from a few seconds to a few hours.

Some accidents terminate spontaneously, due to physical

dispersion of the materials for instance, while others require

human intervention. The heat and energy released by these

power excursions are usually limited.

In contrast, fission-induced emission of high-intensity gamma

radiation and neutrons can have severe consequences on

human health for people in the immediate vicinity, causing

potentially fatal irradiation to workers who are the closest

to the accident, given that gamma and neutron doses can

reach 25 Gy and 20 Gy, respectively, at one meter from the

source during the initial power spikes (1017 fissions). Although

preventive measures reduce the risk of a criticality accident,

it may be appropriate to install detection systems to sound

an alarm from the first power spike, even though this type of

accident is characterized by its sudden onset, without any

precursor signs. For accidents involving more than one power

spike, these detection systems play a particularly important

role for workers not fatally exposed during the first spike, by

facilitating rapid evacuation of personnel, thereby limiting

their exposure to neutron and gamma radiation.

Finally, no criticality accident has ever resulted in significant

radioactive release to the environment, but the latest accident

(Tokai-mura, Japan, 1999) did highlight the importance of

analyzing the impact on the local population.

IRSN's research in the field of criticality accidents focuses on

accident consequences in terms of radiation (gamma and

neutrons) as well as energy release.

Safety objectives

IRSN focuses on developing the skills and resources necessary to:

parameter operators use to assess the off-site radiological

consequences of a potential criticality accident, which helps

determine the type of mitigation action to be taken (e.g .

which zones to evacuate). Acquiring and developing computer

tools to assess accident consequences is essential to achieving

this objective.

Since 1945, about sixty criticality accidents have occurred in

the world, with about forty taking place in research reactors

or in laboratories conducting research on critical assemblies.

Several lessons have been learned from this experience

feedback with regards to the type of risk involved and the

phenomenology characteristic of this type of accident.

What type of risk is involved?

Criticality accident risk is specific to the nuclear industry, one

of several such risks (e.g . dissemination of radioactive

materials, irradiation) which exist alongside conventional

industrial risks such as fire. The risk of a criticality accident

must be taken into account in all situations, both under normal

operating condit ions (process operations, maintenance

operations, transfer operations, etc .) and subsequent to an

incident situation (procedural errors, fires, earthquakes, floods,

etc .).

Two types of event occur when entering the supercritical

state: events that may occur during fissile material transport,

storage, or processing , where a chain reaction is always

accidental, or events that may occur in a nuclear reactor. The

two contexts are very different, since reactors are specifically

designed for the purpose of initiating and controlling a chain

reaction. Only the first type of event, directly related to fuel

cycle facilities and laboratories, can be called "criticality

accidents" in the strictest sense, and they are the focus of

this article.

Criticality accidents result from an uncontrolled fission chain

reaction, which translates as a rapid multiplication in the

number of fissions, generally interrupted by various physical

feedback effects that allow the system to return to a subcritical

state.

The first direct consequence is the energy generated, each

fission releasing around 200 MeV of energy (1 MeV equals

approximately 1.6 ×10-13 joule). The energy released depends

on the sequence of events, which in turn depends on several



2. 5

conditions. Solid materials are mostly metallic , but dry powder

materials (sintered or unsintered) may also be included in

this category, as well as fissile materials contained in matrices,

solidified slag, etc. Accident phenomenology for these materials

is relatively straightforward: the heat balance consists simply

of calculating the heated temperature of the material, taking

into account nuclear power levels, and heat loss induced

through conduction and heat radiation. Feedback effects are

a lways par t ia l ly exp la ined by expans ion and nuc lear

temperature effects(5) (Doppler and neutron spectrum). The

accident ends either by meltdown (or at least plast ic

deformation) and dispersion of fissile material, or by the

intervention of workers (to remove a reflector, add an absorber,

etc .). In general only a single spike is observed, within a

relatively short timeframe (less than a few seconds). however,

the 1997 Sarov accident [IAEA, 2001] progressed in a much

di fferent manner, involv ing power osc i l lat ion without

significant meltdown or deformation of the fissile material.

Consequently, the evolution of such systems must be assessed

with caution.

Liquids

Accident phenomenology for liquids is more complex, involving

radiolysis gas bubble formation (due to water molecule

decomposition by fission fragments and gamma/neutron

radiation) or vapor bubble formation (due to boiling of the

solution) and other fluid movements within the system, in

addition to the usual feedback effects (i.e. expansion and

neutronic temperature effects). The system thus goes from a

single-phase state to a two-phase state. After the first power

spike, gas bubbles are formed by radiolysis and migrate toward

the surface. Their quenching effect disappears as they migrate,

and the power excursion can then start over. This process

accounts for the oscillation phenomenon generally observed

in solution criticality accidents (Figure 1). The possibility of

partial ejection of the solution must also be considered in

open vessel configurations. Gradual evaporation of the solution

over longer timeframes can lead to either increased or

decreased reactivity, depending on the situation. Post-accident

behavior of the system therefore depends on whether it is

assess the possible consequences of criticality accidents in

nuclear facilities to ensure that accidents do not lead to an

unacceptable dose to the population, and that evacuation

zones and assembly stations are in an appropriate location;

analyze any accident situation, providing support for emergency

response in order to provide public authorities with the

information needed to define containment or exclusion areas,

as well as any necessary response zones, based on assessment

of the radiological consequences; and to help the operator and

safety authorities terminate the accident situation, if shutdown

is not spontaneous, and restore safety.

Achieving these objectives requires excellent knowledge of

accident phenomenology, especially the number of fissions.

Sequence of events in a criticality accident

One of the direct consequences of supercriticality is energy

release, mainly in the form of heat, accompanied by intense

neutron and gamma radiation, as well as fission gas release.

heating of the material initiates feedback mechanisms, which

in turn reduce reactivity(3) until the system becomes subcritical,

even temporarily. These interactions typically result in the

first power spike. From that point on, the sequence of events

varies considerably, based on the following parameters:

the physicochemical form of the supercritical fissile material;

the reactivity of the system, representative of the level of

supercriticality;

the initial spontaneous neutron source (which is different

depending on the features of the media involved, i.e. non-

irradiated enriched uranium, irradiated uranium containing

plutonium, or plutonium alone);

the neutronic feedback effects;

the immediate environment and the equipment configuration

where the accident has occurred (heat exchanges between the

supercritical system and the surrounding materials, containment

of the supercritical system, etc.).

Depending on the type, extent, and kinetics of the different

feedback effects (parameters related to the supercritical fissile

material), criticality accidents are generally classified according

to four categories: non-moderated solids, liquids, powders, and

heterogeneous media.

Unmoderated solids

Materials referred to as "solid" include all compact materials

where there is no moderator material(4), even in accident

(3) For a material capable of sustaining a fission chain reaction, reactivity is the param- For a material capable of sustaining a fission chain reaction, reactivity is the param-For a material capable of sustaining a fission chain reaction, reactivity is the param-eter that gives the deviation from criticality, with positive values corresponding to supercriticality and negative values to subcriticality.

(4) Material containing light nuclei such as hydrogen (water, CH2, etc.), acts to slow neutrons, thereby increasing their likelihood of fission.

(5) Temperature variation effects influencing intrinsic neutron properties of materials (absorption, production, and slowing of neutrons).


2. 5

Powders

Criticality accidents may involve powder fissile materials in

the presence of a moderator capable of mixing with the

powder. Accidents of this type are studied regularly, even

though no accident or general accident experience involving

powder fuel materials has ever occurred. The main postulated

scenarios involve the accidental penetration of water in an

environment containing fissile material in powder form. here

again, accident phenomenology is more complex than in a

solution because the material can become a three-phase

system, and a distinction must be made between "expansion"

of fissile material grains and moderator expansion. Powder

wettability and inter-grain fluid migration kinetics are not

well known, and could induce grain movement, deforming the

system's geometry. As a result, a high degree of uncertainty

i s assoc iated with post-acc ident outcomes for these

configurations.

The overall complexity of these phenomena illustrates both

the importance and difficulty of apprehending orders of

magnitude for the consequences of criticality accidents.

Quantifying impact in this way has three key prerequisites:

thorough knowledge of the state of the art and particularly

past accident feedback; access to experimental data; and a

firm grasp of both the empirical and advanced computation

methods to estimate characteristic parameters of the source

term.

closed (vapor recondenses and returns to solution) or open

(evaporation or ejection of the solution eventually restores

a definitive subcritical state). Given all these phenomena, the

duration of solution criticality accidents is extremely variable,

ranging from a few seconds to a few hours.

Heterogeneous media

This type of material consists of nuclear fuel rod or plate

assemblies, or pieces of solid fissile material immersed in a

moderator, usually water. Criticality accident phenomenology

in this case is even more complex than in a solution, given

that the material can become a three-phase system with

vapor or radiolysis gas bubbles, or even a vapor film at the

solid-liquid interface. heat transfers must therefore take into

account three types of interface: solid-liquid, solid-gas, and

liquid-gas. In most cases, the material is initially under-

moderated, and as in solution accidents, several spikes can

occur, but the characteristic time depends largely on the

heterogeneity of the materials.

Power

Exponential power

increase

Bubble migration and

release

Time

1st spike

2nd spike FREE EVOLUTION

Pseudo-equilibriumOscillations

Solution heating +

Radiolysis gas formation

Figure 1



2. 5

established and systems are in place for detection, protection,

and termination of the chain reaction) and accidents in process

fac i l i t i es (where a l l conf igurat ions a re des igned for

subcriticality).

In general, several types of lessons have been learned from

criticality accidents.

Regarding their causes, most criticality accidents took place

during non-routine operations. None of them were exclusively

attributable to equipment defects, and none took place during

transport or in storage.

The accidents arose from a wide range of fissile materials,

varying in type (U, Pu) and enrichment. They usually involved

liquid materials (fuel cycle plants) and metal media (especially

in research facilities). None of the events involved fissile

material in powder form, and accidents with heterogeneous

media only occurred in research facilities.

The mechanical consequences of these criticality accidents

were very limited, aside from cases where a steam explosion

followed the power excursion.

In process accidents in particular, the total number of fissions

for the first spike, where this number could be reconstructed,

ranged from 3x1015 to 2x1017. The total overall number of

fissions ranged from 1015 to 2.7x1018, except for one accident

involving very large amounts of fissile material and a number

of fissions equal to 4x1019. Most of the 22 accidents did not

end after a single power spike. Duration ranged from a few

seconds to around 40 hours, with more than half the accidents

requiring human intervention to terminate the fission chain

reaction. Analysis has shown that the pre-accident scenarios

were diff icult to predict and several (more than two)

independent failures were reported.

Parallel to this detailed, systematic analysis of past accidents,

experimental research has been conducted since the 1960s

to investigate criticality accident phenomena and estimate

potential consequences.

Experimental data

Starting in 1967, France became one of the first countries to

initiate ambitious experimental programs aimed at building

knowledge on solution criticality accidents, using fissile

solutions of uranyl nitrate highly enriched in 235U. Various

the state of the art

Criticality accidents in safety analyses on French

nuclear facilities

To define mitigation measures and facilitate crisis management

for a criticality accident in a nuclear facility, the methodology

generally used in France centers on a conservative "bounding"

fission source term for the first hours of the accident (typically

5x1018 fissions). This fission source term is based on data

from experimental facilities (namely the French CRAC and

SILENE facilities at CEA/Valduc) and analysis of past criticality

acc idents in the wor ld . As a ru le , cr i t ica l i ty acc ident

configuration in a given facility has little influence on the

fission source term value. At 5x1018 fissions, this value

represents the maximum number of fissions for a two-hour

period, obtained during experiments in the CRAC and SILENE

facilities.

however, the bounding fission source term approach does not

take into account accident kinetics (time aspect) or the

conditions specific to the facility. It may lead decision-makers

to adopt excessive mitigation measures (large evacuation

zones, additional protection measures outside the site for

facilities with limited radiation protection). Consequently, a

" f iner" approach, one that integrates speci f ic fac i l i ty

configurations as well as accident kinetics, may represent an

improvement worth considering for the assessment of

consequences and measures.

Insights gained and research underway

Experience feedback on past criticality accidents

In-depth study of recorded criticality accidents in the world is

a precious source of information for apprehending accident

phenomenology.

Since 1945, 60 criticality accidents have been documented in

the world, most occurring in the US and the former USSR. Thirty-

eight took place in research facilities (critical reactors and

assemblies) and 22 in process facilities. In France, two accidents

occurred in experimental reactors, neither resulting in severe

irradiation. All 60 criticality accidents and the relevant information

available are described in a document published by Los Alamos

National Laboratory [McLaughlin et al., 2000].

A distinction is usually made between accidents in research

facilities (where criticality or near-criticality is purposely


2. 5

obtained on CRAC and SILENE, along with criticality accident

phenomenology. It was found that the total number of fissions

depends mainly on total reactivity insertion, and that the first

power spike characteristics depend mainly on the conditions

in which reactivity is introduced initially. Another important

parameter in the uncertainty relative to the fission source term

involves the calibration procedures applied prior to the

experiments. This point was therefore examined in detail.

Feedback effects, such as radiolysis and boiling, represent the

main phenomena that influence the sequence of events in the

accident over time. Knowledge of all these phenomena is crucial

to improving current accident codes, aimed at providing an

order-of-magnitude estimate for the fission source term, and

in developing more efficient tools.

In addition to providing a better understanding of criticality

accident phenomenology, the experimental results, particularly

for fissile solutions, have given rise to simple equations providing

a bounding fission source term estimate expressed as the total

number of fissions.

Empirical and simplified equations

Based on lessons learned from past criticality accidents

[McLaughlin et al., 2000] and results from fissile solution

experiments, various equations (some empirical, others based

on a simplified heat balance) were developed in France and

elsewhere to estimate the bounding number of fissions

generated by a criticality accident [Nakajima, 2003], without

requiring any specific details about the accident scenario.

IRSN inventoried and analyzed the various equations established

for fissile solutions, comparing them against experimental

results. The findings highlighted assumptions, some implicit,

that underlie these relationships and considerably limit their

range of validity. IRSN was then able to explicitly define the

range of validity for these equations and hence their area of

applicability.

This inventory also enabled IRSN to propose a new formulation,

based on existing relationships, to estimate a bounding value

for total number of fissions in homogenous fissile solutions.

The new relationship determines the maximum number of

fissions for two cases, assuming negligible heat loss. In the first

case, the chain reaction self-terminates (or is terminated) prior

to boiling (solution heated to 100°C maximum). In the second

case, the chain reaction self-terminates after the solution

reaches the boiling point (once boiling occurs, the solution is

assumed to evaporate down to the minimum critical volume

facilities capable of producing "controlled" criticial excursions

have been built around the world, for example:

GODIVA (U metal) and JEZEBEL (Pu metal) in the US;

ShEBA (solution of uranyl fluoride enriched to 5% 235U) in

the US;

KEWB (highly enriched uranyl sulfate solution) in the US;

CRAC and SILENE (highly enriched uranyl nitrate solutions)

at CEA/Valduc in France;

TRACY (solutions of uranyl nitrate enriched to 10% 235U) in

Japan.

Experiments have mainly focused on solutions representing

the greatest risk, based on feedback from past criticality

accidents.

As part of its mission, IRSN has engaged in an in-depth analysis

of CEA/Valduc experimental results and broadened theoretical

knowledge of the physical phenomena associated with criticality

accidents. For reasons involving access to detailed data, this

work has focused on the 72 experiments conducted on the

CRAC facility (designed to study the radiological consequences

of criticality accidents) from 1967 to 1972, then on experiments

conducted on the SILENE facility (a free neutron evolution

irradiation source, Figure 2) starting in 1974. This work

culminated in an initial summary document, bringing together

over 100 reports and articles describing experimental results

Figure 2



2. 5

Initiated in the mid-1980s through an EC-funded contract and

a partnership between AEA-SRD (UK) and IPSN (now IRSN),

CRITEX is structured around "simple" methods based on physical

laws and more empirical observations drawn from CRAC and

SILENE experimental results. It is particularly suitable for

cylindrical geometries like those of the CRAC and SILENE

vessels.

The area of validation for CRITEX currently includes some 30

experiments conducted in CRAC and SILENE as well as the

American ShEBA reactor. IRSN is currently expanding the

database to include the new CRAC and SILENE experiments,

and the Japanese TRACEY experiments. Unfortunately, all these

studies involve cylindrical equipment and uranium solutions,

limiting the area of applicability of the code.

CRITEX is currently undergoing in-depth analysis to identify

shortcomings and map out areas for improvement. The first

step consisted in an exhaustive review (to identify the various

procedures and variables, their functions, the physical models

used, etc.). The current step focuses on assessing the models

and their validity, and it is already apparent that major changes

will be needed to accurately take into account geometries other

than cylinders.

IRSN plans to analyze its other criticality accident codes as

well, first POWDER then ChâTEAU. however, unlike CRITEX,

there are no experiments validating the media simulated in

Vc[Φ] in the equipment considered, while in practice self-

shutdown occurs at an equal or greater volume, depending on

solution concentration). The bounding number of fissions can

be expressed as a function of volume (V) and solution density

(dsol) using the following relationships, with volume in liters

and density in kilograms per liter, where de is water density:

Nf = 1.3x1016 .V .dsol

(no boiling)

Nf = 1.3x1016 .V .dsol + 8x1016.[V-Vc(Φ)].de

(if boiling occurs).

The advantage of these equations, applicable to all geometries,

is that they require only minimum knowledge of the accident

scenario. however, they are not valid if forced cooling is used,

or if the solution recondenses after boiling. Nor can they be

used to determine time-dependent changes in nuclear power

and total number of fissions. Computer codes are therefore

indispensable, particularly in the post-accident management

of situations requiring intervention.

Computer codes

Most computer codes used for criticality accident applications

were developed for emergency response. The operational

requirements therefore center on the simplest possible

implementation, very short computing times, and order-of-

magnitude estimates of characteristic parameters of the fission

source term. Although relatively basic , the models used are

adequate for the targeted objective.

Far more complex tools, such as FETCh [Pain et al., 2001],

make it possible to perform much more rigorous analysis, given

the models used, but require prohibitive computing time and

resources.

Various criticality accident simulation codes have been

developed or co-developed in France, each designed for a

particular type of fissile material: CRITEX for fissile solutions,

POWDER for powders, and ChâTEAU for fuel rods immersed

in water. All three codes are based on a common architecture,

shown in Figure 3. Considered "simplified" computer codes,

they can be used to quickly estimate, for the defined geometry

(for instance, cylindrical for CRITEX), time-dependent changes

in the power, energy, and temperature of the material over a

limited period (first minutes of the accident). CRITEX is the

most frequently used code at present.

Similar codes have also been developed in other countries:

AGNES, TRACE, and INCTAC for solutions, and AGNES-P for

powders.

Figure 3

Accidental insertion of

reactivity

Reactivity balance

Point kinetics

equations

Power Energy

Temperature

FEEDBACK MECHANISMS

TEMPERATURE EFFECTS

VOID EFFECTS(radiolysis gas, vapor, etc.)

Doppler effect

Spectral effect

Expansion effect (density, leakage, etc.)


2. 5

also a need to assess the possibility of using criticality

accident codes to aid emergency response in the event of an

accident.

In a small number of cases, bounding fission source term

determination was based on robust arguments and

simplified equations, taking into account lessons learned

from past accidents as well as specific characteristics of the

facility in question. Here again, it is essential to accurately

define the area of applicability for these simplified

equations.

Past accident feedback and experimental programs indicate

that criticality accidents can vary in duration, potentially

exceeding several hours. In this regard, the Tokai-mura

accident underscored the need to consider ahead of time

which shutdown mechanisms could be implemented in case

of a criticality accident. Computer codes may serve as a

valuable tool in this context, as they integrate the temporal

aspect of accident progression.

Finally, few countries have computer codes or experimental

facilities dedicated to criticality accidents. Hence the need

for IRSN to foster cooperation with the few partners

possessing such capabilities, in an effort to improve its own

knowledge. The first step was to exchange technical

information with CEA/Valduc. Collaborative projects were

also forged with JAEA (Japan Atomic Energy Agency), and

since 2006, IRSN has been participating in the OECD/NEA

Expert Group on Criticality Excursion Analyses [OECD].

In 2007, IRSN also proposed a draft version for an ISO

standard on "fission source term" estimation for criticality

accidents. This standard will provide methodology and

recommendations for assessing the bounding number of

fissions, as part of a safety analysis focused on a postulated

criticality accident.

POWDER or ChâTEAU. For now, the models can only be

validated by comparisons between codes.

One of the fundamental questions examines the extent to

which results from these computer codes can be applied to

geometries other than those simulated. To lead off the

investigation into this difficult issue, IRSN proposed a case

study (parallelepiped equipment containing a highly enriched

uranyl nitrate solution) to the OECD/NEA(6) Expert Group on

Criticality Excursion Analyses [OECD].


Based on the state of the art in France, a fission source term

on the order of 5x1018 fissions is considered sufficiently

conservative for analyzing consequences in the initial hours

of a criticality accident. This value is generally used in all

new facilities posing a criticality accident risk. Certain

operators, however, have already considered the possibility

of using a less conservative maximum total number of

fissions, taking into account specific facility characteristics,

in cases where the direct irradiation consequences estimated

using the bounding value (5x1018 fissions) were significant

outside the facility and/or site.

For example, in safety demonstrations based on feedback

from the Tokai-mura accident (Japan, September 30, 1999),

operators attempted to use the CRITEx code, with varying

degrees of success, to determine a more realistic fission

source term than the value mentioned above (5x1018

fissions) for assessing potential accident consequences, or

defining evacuation areas or assembly stations.

Unfortunately, in most attempts of this sort, criticality

accident codes are used outside their areas of applicability.

Results must therefore be interpreted with caution, as the

uncertainty estimate for the calculated values is not

accurately known. Significant improvements must be made

in the codes to validate a code-based approach. There is

(6) OECD: Organisation for Economic Co-operation and Development NEA: Nuclear Energy Agency.



2. 5

References

IAEA, The criticality accident in Sarov, 2001 – http://www-pub.iaea.org/MTCD/publications/PDF/Pub1106_scr.pdf.

T.P. McLaughlin et al., A Review of Criticality Accidents, 2000 Revision, LA-13638 – http://www.csirc.net/docs/reports/la-13638.pdf.

K. Nakajima, Applicability of Simplified Methods to Evaluate Consequences of Criticality Accident Using Past Accident Data, ICNC 2003 – http://typhoon.jaea.go.jp/icnc2003/Proceeding/paper/8.4_171.pdf.

The OECD/NEA Expert Group on Criticality Excursion Analyses – http://www.nea.fr/html/science/wpncs/excursions/index.html.

C.C. Pain et al., Transient Criticality in Fissile Solutions — Compressibility Effects, Nucl. Sci. Eng., 138, p. 78-95 (2001).


2. 6

The risk of containment failure during a severe accident arises

from the multiple types of load applied to the structure, which

may exceed its design pressure. These pressure loads and the

accompanying temperature changes constitute the thermo-

mechanical load over time. Studies conducted as part of the

PSA2 project on 900 MWe PWRs aimed to assess containment

response to a quasi-static load, i.e. a pressure spike or slow

pressure rise [Raimond et al., 2004].

Linear simulations conducted on several severe accident

categories allowed researchers to identify the scenario that

would cause the most damage to the containment building.

Designated the "AF scenario", it consists of three phases

(Figure 1):

the pre-thermal load phase, corresponding to core degrada-

tion; instants P1 and P2 represent the beginning and end of

this phase, respectively;

the pressure and temperature spike phase, corresponding to

the adiabatic isochoric hydrogen combustion induced by core

oxidation; instant P3 represents this spike;

the slow rise in pressure and temperature, corresponding to

corium-concrete interaction, assumed to bring corium into

contact with sump water; instants P4 and P5 represent, respec-

tively, the beginning and end of this phase (assumed to last

100 hours, with final absolute pressure reaching 10.02 bar).

Several pressure and temperature levels for the spike phase

(P3) were studied. A spike limit pressure of 11.44 bar absolute

(2.61 times the containment design pressure) was selected,

corresponding to adiabatic isochoric combustion of 125% of

the maximum amount of hydrogen produced by core oxidation.

The spike was assumed to last 90 seconds (a 30-second rise

and 60-second fall), consistent with static simulation assump-

tions. These limit values selected for the AF scenario can be

Georges Nahas, Bertrand CiréeCivil Engineering and Structural Analysis Unit

The containment building serves as the third and final barrier against environmental release of radioactive

products from the reactor core. Its integrity, especially in accident situations, is critical to nuclear safety.

IRSN carried out a major project on containment integrity as part of the level 2 probabilistic safety study

(PSA2) on CPY-series 900 MWe pressurized water reactors (PWRs). This project took on the ambitious

scientific challenge of assessing the risk of containment leakage after a severe accident leading to core

meltdown. Reaching across several fields of expertise, the studies aimed to meet the following objectives:

identify the various severe accident scenarios and probabilities of occurrence, analyze the mechanical

behavior of containment, and assess containment integrity along with contamination risks for the imme-

diate environment under severe accident conditions [Raimond et al., 2004].

ANAlySIS oF the mechANIcAl behAvIoR oF containment on CPY 900 MWe PWRs under severe accident conditions



2. 6

used to determine the containment safety margins in a severe

accident context.

In order to quantify the effect of temperature on severe accident

loads and extrapolate results to other severe accident scenarios,

without requiring an unreasonable number of non-linear cal-

culations, the selected loads were limited to three scenarios:

AF scenario;

AS scenario (AF scenario without pressure/temperature

spike);

PL scenario (pressure only, no temperature loading).

multiscale approach

The study was based on deterministic best-estimate calcula-

tions, performed using non-linear finite element analysis. A

multiscale approach was adopted to apprehend structural

behavior at various levels of detail, i.e. the straight section of

the containment building, the equipment hatch, and the equip-

ment hatch closing system. This approach made it possible to

realistically represent the different thermomechanical phe-

nomena, while keeping computing time and cost within reason-

able limits (Figure 2).

This article presents the simulations run using the global com-

plete containment model, the quarter containment model, the

local equipment hatch penetration model, and the detailed

model [Cirée et al., 2007],[Nahas et al., 2007].

The simulations were optimized on the IRSN computation

server prior to complete containment simulation using the

CAST3M computer code [Verpeaux et al., 1989], which ran

within reasonable computing time (275 hours).

Global containment model

(simulation of initial containment state)

Severe accident simulation requires that knowledge on the

initial, pre-accident state of the structures, under the effects

of creep and shrinkage phenomena, is as close to reality as

possible. For the PSA2 project, the containment age was set

to 30 years. The prestressing reference simulation to establish

the condition of the building structure after 30 years of service

was run on a generic 900 MWe PWR containment building. The

reference reactor was Unit 3 at the Blayais nuclear power

plant.

The position and tensioning of prestressing tendons is unsym-

metrical, making it necessary to simulate the complete contain-

ment (360°).

The mesh used simulates the containment concrete, passive

reinforcement, metal liner, protective concrete on the basemat,

and, with simplified modeling, the equipment hatch shell with

sleeve, flanges, and head. Internals are also modeled in a simpli-

fied manner, but all prestressing tendons are modeled with

precision, including their geometry and deviations, especially

around the equipment hatch and the two personnel airlocks.

The model also simulates ground effect and backfill effect.

Figures 2 and 3 depict the meshes used for this simulation.

The concrete containment is subject to its own weight and

prestressing from the tendons, which is calculated taking into

account the nine phases of tensioning over a nine-month

period. Prestressing calculations integrate tendon tension loss

caused by friction (linear and angular) and relaxation (2.5%

at 1000 hours), tension loss due to tendon anchor slip, and

instantaneous tension loss due to elastic concrete shrinkage.

Vertical tendon loads are calculated with tensioning at only

one end, except for tendons deviated around the equipment

hatch, which are tensioned at both ends. Loads for horizontal

and dome tendons are calculated with tensioning at both

ends.

For each tendon tensioning phase as well as the 30-year service

period, concrete shrinkage and creep were estimated in accor-

dance with regulations, using equations from BPEL 1999 (regu-

latory document on limit-state design of prestressed concrete).

These parameters were introduced at each simulation stage

as "initial strain" loads, dependent on concrete drying, load

age, and stress field. The calculation of prestressing and creep

at 30 years served as the basis for the entire mechanical study

as well as the simulations performed using the various

models.

Metal linerTmax Containment concrete Tmax Pressure

Temperature (C°) Absolute pressure (bar)250

200

150

100

50

00

50 000100 000

150 000200 000

250 000300 000

350 000400 000

450 000500 000

Time (s)

30

28

26

24

22

20

18

16

14

12

10

8

6

4

2

0

P1P2

P3

P4

P5

Figure 1 Changes in containment pressure and temperature for the AF scenario.


2. 6

were carried out in linear transient mode. Limit conditions

between the containment concrete and the steel liner assume

they are separated by a layer of air layer to take into account

thermal resistance between the concrete and steel. The air

layer is necessary to simulate the temperature jump between

the two materials, despite their physical continuity. The metal

liner is in fact used as formwork during containment construc-

tion (Figure 3).

Like the complete model, the quarter containment model

simulates the containment concrete, prestressing tendons,

passive reinforcement, metal liner, protective concrete on the

basemat, internal structures, and the equipment hatch shell

Quarter containment model

(severe accident simulations)

Thermomechanical simulations for severe accident conditions

were carried out using a mesh representing one quarter of the

containment, to reduce computing time.

Prestressing and creep calculations performed for the complete

containment model were projected on the "quarter contain-

ment" model, before severe accident pressure and temperature

loads were applied. These calculations used a best-estimate

approach that ignored variability in material characteristics.

For the AF and AS scenarios, thermal calculations to define the

temperature field at various instants in time during loading

Global complete containment model

Global quarter containment model

Local penetration model

Detailed sleeve/flange/head model

Figure 2 Nested multiscale models: global complete containment model, global quarter containment model, local penetration model, sleeve/shell/flanges/head detailed model.

Figure 3 Global complete containment meshes of prestressing tendons, reinforcement, metal liner, and equipment hatch.



2. 6

Comparing simulation results between the three scenarios

(AF, AS, and PL) makes it possible to assess the temperature

effect for the accident load applied (Figure 5). For example,

comparing the AF and AS curves in Figure 5 shows that contain-

ment response is nearly reversible in the straight section (67.5°

vertical plane at 22.9 m).

Overall stability of the containment is maintained by the

integrity of the prestressing tendons.

Maximum equivalent plastic strain in the metal liner during

the AF scenario after the pressure spike (P4) is greater than

that obtained during the spike (P3) caused by thermal loading

(Figure 6).

Results from these three accident loads can be used to

extrapolate mechanical containment behavior to other sce-

narios, since pressure was shown to be the main factor driving

the mechanical phenomena.

Leak paths are formed by any tears in the metal liner and

cracks in the prestressed concrete containment wall.

Calculated strain values for the metal liner are far below the

fracture strain values for steel, and according to the results

obtained, the liner should remain leaktight, with no signs of

tearing.

with sleeve, flanges, and head. Figure 2 depicts the mesh used

for this simulation.

The concrete is modeled by eight-node, linear, solid finite

elements, using a non-linear diffuse cracking behavior law

(Ottosen model). Prestressing tendons and reinforcements are

modeled by two-node rebar finite elements, with a non-linear

elastoplastic behavior law integrating isotropic strain hardening.

Tendon and reinforcement meshes, independent of the concrete

mesh, are linked by kinematic relationships. The metal liner is

modeled by shell elements with a non-linear elastoplastic

behavior law integrating isotropic strain hardening. The ground

is simulated using a superelement. Basemat uplift is possible,

depending on severe accident loading [Verpeaux et al., 1989].

Analysis of non-linear computation results

Analyzing results for the three severe accident simulations (AF,

AS, and PL) led to the following observations.

The AF simulation confirmed zones identified as the most

sensitive areas of the containment building (Figure 4), particularly

around the equipment hatch and also the gusset zone, which

sustained crosswise cracking toward the prestressing tunnel.

AF scenario at9.06890E+04 sec

Amplitude1.00E+020.0

W1 in 67.5° plane W1 in 0° plane at 90689 sec

1.00E-02

9.50E-03

9.00E-03

8.50E-03

8.00E-03

7.50E-03

7.00E-03

6.50E-03

6.00E-03

5.50E-03

5.00E-03

4.50E-03

4.00E-03

3.50E-03

3.00E-03

2.50E-03

2.00E-03

1.50E-03

1.00E-03

5.00E-04

> -4.50E-03< 1.77E-02

0.0

Figure 4 Containment deformation amplified 100 times and concrete cracking in equipment hatch axis and in straight section during AF scenario spike (P3).


2. 6

offered differing explanations of what caused the initial liner

tears, such tears are always caused by strain localization. The

absolute pressure level leading to containment loss was around

10 bar in both tests [hessheimer et al., 2006].

PCCV test results and their numerical simulations were analyzed

as part of the benchmark exercise for the OECD International

Standard Problem No. 48 (ISP48), in which IRSN participated.

Simulations were run with CAST3M, using the same approach

as for the PSA2 project [International Standard Problem,

2004].

Neither IRSN simulations nor those run by the other participants

predicted these tears at an absolute pressure of 10 bar, even

though the various geometrical non-linearities were taken into

account [International Standard Problem, 2005].

At this level of containment pressure loading, measured cir-

cumferential strain of the metal liner in the straight section

of the barrel is 0.17%, and the calculated equivalent plastic

strain is around 0.3-0.5%. This value falls below the limit values

determined by liner characterization tests performed after the

mockup tests.

In contrast, calculations correctly simulated the structural

failure mode test and the tear observed at the end of this test.

This can be explained by modeling uncertainty and assumptions

taken into account for the calculations. While metal liner tearing

is a local phenomenon occurring at the weld scale, the calcula-

tions are conducted on an overall scale, where finite element

size represents roughly twenty to thirty centimeters. To repro-

duce any liner tears, the models must be at the same scale as

the phenomenon and take into account discontinuities pre-

To assess the risk of containment failure (metal liner, pre-

stressed concrete wall), before analyzing and interpreting the

results of the preceding studies based on the deformations

obtained, it was first necessary to obtain experimental results

to define acceptability criteria for the non-linear calculations.

Simulation results were therefore compared to experimental

results obtained using mockups, namely PCCV (NUPEC – NRC

– Sandia National Laboratories), in order to correlate, when

possible, the type of failure and the associated leak rates. A

group of experts participated in this analysis in order to define

these criteria.

Analysis of mockup test results

Results from representative tests constitute an important

element in validating CAST3M simulations. The challenge is to

find tests representative of the load conditions in question

[hessheimer et al., 2006].

PCCV (NUPEC – NRC – Sandia National Laboratories) is a 1:4-

scale mockup of a prestressed concrete containment with metal

liner. Pressure tests in dry air at ambient temperature were

carried out at Sandia National Laboratories, followed by a

structural failure mode test using water.

The PCCV tests resulted in liner tears, with significant leakage,

for absolute pressure values of around 10.7 bar (2.5 times the

design pressure) [International Standard Problem, 2004].

Another test in the 1:6-scale RCCV mockup of a reinforced

concrete containment with metal liner (NRC – Sandia National

Laboratories), conducted under pressure load conditions, pro-

duced similar results. Although analyses following the tests

AF scenario AS scenario PL scenario

Displacement (m)0.06

0.05

0.04

0.03

0.02

0.01

0

–0.01

–0.020 1 2 3 4 5 6 7 8 9 10 11 12

Absolute pressure (bar)

AF scenario AS scenario PL scenario

Equivalent plastic strain1.3x10-2

0

1x10-3

2x10-3

3x10-3

4x10-3

5x10-3

6x10-3

7x10-3

8x10-3

9x10-3

1x10-2

1.1x10-2

1.2x10-2

0 1 2 3 4 5 6 7 8 9 10 11 12Absolute pressure (bar)

Figure 5 Radial displacement at +22.9 m in 67.5° plane. Figure 6 Maximum equivalent plastic strain in metal liner.



2. 6

Based on the single "quarter containment" simulation, several

"penetration" accident simulations were carried out, with

variations in certain parameters such as the mesh, mechanical

bolt characteristics, rebars, limit conditions, and bolt tightening

torque. These sensitivity studies estimated the uncertainty

related to modeling, computing, and materials as being around

15%.

Flanges are modeled with shell elements, as are the liner, gus-

sets and collars, and sleeve/shell/flanges/head system. Rebars

and tendons along with the 44 equipment hatch closing bolts

are modeled by two-node rebar elements (Figure 7). Mesh

topology (geometric node position, discretization) is indepen-

dent of concrete mesh topology. Mechanical connections with

the concrete (sleeve, tendon, and rebar anchoring) are repre-

sented by linear relationships. Unilateral (simple contact)

relationships simulate the "non-interpenetration" of the flanges,

including the possibility of disjunction [Verpeaux et al., 1989].

Given the number of linear and unilateral relationships, their

management must be optimized as part of the numerical

resolution.

In the local model, three bolt types were considered (RCC-M

data):

E24 steel bolts currently used on 900 MWe PWR contain-

ments in France (diameter: 33 mm; yield strength: 238 MPa;

tightening torque: 69 and 140 MPa);

Z6 CNU 17.4 steel bolts (diameter: 33 mm; yield strength:

729 MPa; tightening torque: 369 MPa);

40 CNDV 0703 steel bolts (diameter: 24 mm; yield strength:

852 MPa; tightening torque: 273 MPa).

To obtain local model limit conditions, displacement fields

from the global simulation (quarter containment) are pro-

jected on the local model contour at each time step, based

on the multiscale approach. This method is validated by

comparing results from differently sized local models and by

running basic test cases.

Given the lack of experimental mechanical data on aged seal

behavior, the studies do not take into account the seal between

the two flanges, therefore limiting the possible outcome to

disjunction of the flanges. Initial tests conducted by IRSN

show that confinement of the seal to its housing at high

temperature, due to its strong thermal expansion coefficient,

can substantially impair reversibility.

Even though the most realistic approach possible was used,

certain simplifications were adopted for the penetration

simulations. These simplifications are considered second-order

elements relative to the main modeling elements.

sented by each weld and anchor, along with any cracks in the

concrete, using tools capable of simulating strain localization

in the structure.

Transposing PCCV results to the PSA2 containment calculations

led to the following criterion: maximum plastic strain obtained

by non-linear calculations in the straight section must be less

than 0.30% ± 0.15%.

Above this value, liner tearing is highly probable, as a result of

strain localization. Strain around the tear corresponds to values

(around 10%) determined by the liner characterization tests.

The criterion retained is thus relevant to containment modeling,

rather than the material.

More specifically, this leakage criterion, related to the modeling

used and the mesh fineness, takes into account uncertainty

inherent to the models and assumptions.

Generalizing this criterion to cases of thermomechanical loading

limits the mechanical aspect of strain to the 0.30% value

[International Standard Problem, 2005].

The 0.3% metal liner strain value recommended by the expert

group on the basis of the mockup test results corresponds to

a containment pressure of around 10.5 bar absolute for the AF

scenario, and 9.75 bar for the pressure-only PL scenario. hence,

the mean containment failure pressure is assumed to be around

10 bar absolute (2.25 times the design pressure).

Local equipment hatch model

The quarter containment model, which incorporates prestressing

tendons and passive rebars as well as non-linear mechanical

behavior laws, requires significant computing time, even though

spatial discretization of the geometry is relatively coarse. A finer

model was therefore adopted to study behavior in sensitive

zones such as the equipment hatch, particularly the risk of

flange disjunction in the containment closing system, resulting

in direct leakage to the atmosphere. This model represents the

exact geometry of the flanges, along with the bolts joining them

together. Featuring the same elements as the global model

(concrete, metal liner, reinforcement, and prestressing tendons

covering an area of the barrel 10.60 m wide and 23.40 m high,

the shell, flanges and bolts, gussets, and collars anchoring the

shell in the concrete, etc.), the local model also applies the same

thermomechanical loads and material behavior laws. In addition,

prestressing, shinkage, and creep from the global model are

projected on the local model. An initial iterative calculation

"rebalances" the structure and achieves the initial mechanical

state of a 30-year-old containment, as in the global model.


2. 6

sleeve/concrete junction modeling.

Detailed model

Modeling the flange connection is one of the most complicated

aspects of the study, and generally the most sensitive in terms

of flange disjunction, which appears to be the predominant

leakage mode for the structure during the pressure rise. The

difficulty results from the choice of shell elements for the local

model. This led researchers to develop a detailed model, with

The main results are as follows:

the choice of bolts (cross-section, yield strength) is the

critical parameter in the mechanical study, with considerable

repercussions on the degree of disjunction (Table 1);

the spike in pressure and temperature (P3) has relatively

little impact on pressure-dependent disjunction values; flange

disjunction is thus largely influenced by containment ovalization

and buckling around the sleeve, neither effect being very sensi-

tive to temperature;

regardless of the scenario and parameters, as pressure sub-

sides, the flanges only close partially, due to bolt yielding and

lack of reversibility in the concrete containment deformations

around the equipment hatch (Figure 8);

the disjunction profile along the sleeve circumference is

more or less constant, with a disjunction length of around 4

m (for the half-circumference), and the leak cross-section is

almost proportional to maximum flange disjunction;

decreased prestressing has relatively little impact on disjunc-

tion during the pressure rise (slightly earlier disjunction);

there is generalized concrete cracking during the pressure

and temperature spike (P3);

disjunction is relatively unaffected by changing the bolt

tightening torque, the concrete behavior law parameters, and

Penetration7164 nodes 5676 elements

Sleeve

Reinforcement

Collars

Gussets

Flanges

Bolts

Figure 7 Concrete penetration, liner/sleeve/flanges/head, tendons, sleeve, flanges, and bolts.

Potentialleak

areas

Maximumflange

disjunction

0.1 cm2

6-7 mm

1 cm2

33-41 mm

10 cm2

264-288 mm

50 cm2

1066-1103 mm

AF scenarioE24steel bolts

5.13 6.03 7.26 9.57

AF simulationZ6 CNU 17.4steel bolts

5.13 6.64 9.36 > 12.00

AF simulation40 CNDV 07.03steel bolts

2.75 5.50 7.20 10.19

Table 1 Absolute disjunction pressure and maximum disjunction as a function of calculated potential leak area.



2. 6

projected on the detailed model contour. An initial iterative

calculation (prior to the accident simulation) balances the

structure and the bolt tightening torque.

Using solid elements and more accurate flange, shim, and clamp

geometries, the detailed model makes it possible to define

more realistic limit conditions for contact and friction between

the flanges, with the imposed displacement provided by simula-

tion of the same scenario using the local equipment hatch

model. In this way, the detailed model provides new insights

into the closing system's behavior. For example, it revealed the

predominant effect of shear stress on bolts leading to failure,

for moderate pressures, with a significant risk of compromising

closure integrity. The normal load profile between the flanges

is maximal at the clamp, and non-uniform around the circum-

ference, which explains the moderate effect of changing tight-

ening torque values and the friction coefficient:

bolt yield strength is reached from the beginning of the AF

scenario, after bolt tightening and establishment of steady-state

mode (depending on the tightening torque selected);

ultimate tensile strength for the bolts is reached at 8 bar

absolute;

plastic strain in bolts increases considerably during the

pressure and temperature spike, eventually reaching values that

have no physical significance;

opening at the seal is slight (less than 0.1 mm), despite

the following main characteristics:

solid elements are used to model the metal sleeve, flanges,

bolts, and hemispheric head, to avoid the problems posed by

shell and beam elements in defining limit conditions;

a much finer mesh is used, providing a more accurate rep-

resentation of real geometry (changes in thickness, weld bevels,

etc.);

the gussets, collars, concrete, and rebars are not represented

in this model, based on the assumption that concrete imposes

its displacements and deformations on the less rigid metal

components.

The thermal calculation for the AF scenario was performed on

the detailed sleeve model. The same behavior laws were used

as in the equipment hatch simulation. The total linear element

mesh, covering half the circumference (for reasons of sym-

metry), consists of 61,973 elements and 76,920 nodes

(Figures 9 and 10).

Unilateral relationships simulate attachment of the two flanges

and contact with the bolts. Play can be introduced in these

relationships. Unilateral contact with sliding and friction char-

acterizes the connection between the shims, clamp, and the

opposite flange. Implementation of the detailed model is similar

to local model implementation. Limit conditions for imposed

displacements are taken from the local model simulation and

AF30 E24_238 AF30 Z6 CNU 17.4 AF30 40 CNDV 07.03

0

1.8x10-2

0

1.6x10-2

1.4x10-2

1.2x10-2

1x10-2

0.8x10-2

0.6x10-2

0.4x10-2

0.2x10-2

2 4 6 8 10 12

2.5x10-3

2x10-3

1.5x10-3

1x10-3

0.5x10-3

0

EPSP UX

Relative pressure (bar)

1.67%

1.10%

0.48%

0 2 4 6 8 10 12Relative pressure (bar)

2.2 mm

1.9 mm

0.95 mm

E24

40 CNDV 07.03

Z6 CNU 17.4

E24

40 CNDV 07.03

Z6 CNU 17.4

Max: 2.23x10-3

Min: 0

Max: 1.71x10-2

Min: 0

Figure 8 Plastic strain and maximum disjunction under seal as a function of relative pressure (in red: e24 steel bolts; in blue: Z6 CNU 17.4 steel bolts; in turquoise: 40 CNDV 0703 steel bolts).


2. 6

noticeable disjunction at the outer perimeter of the flanges.

This model shows:

the complex flange deformation mechanisms, excluding any

interpolation, and the strong coupling between flange ovaliza-

tion and buckling;

the limited impact of axisymmetrical stress on the sleeve in

terms of disjunction and bolt shearing (pressure loading ,

prestress-induced sleeve pinching), and the considerable non-

axisymmetrical sleeve strain imposed by the containment during

the AF accident scenario (pressure and thermal loading), which

result in flange buckling and ovalization;

for bolts subjected to shear stress, the importance of bolt

selection (cross-section, grade), which had a direct influence

on the results (Figure 11).

Ignoring the play between flanges and bolts, in the case where

both flanges can slide relative to each other, bolt yielding occurs

at low pressures (3.2, 3.8, and 5.5 bar absolute for E24,

40 CNDV 0703, and Z6 CNU 17.4 bolts respectively). These low

pressure values, due to bolt shear stress, are highly sensitive to

bolt/flange play. But given the equipment hatch closing problems

on certain reactors and to remain conservative, it appears justi-

fied to minimize the flange-bolt play considered in the

calculations.

The shear criterion reached at low pressures (without play) is

indicative of the weak closing system and the risk of bolt failure

along a large portion of the circumference. The margin added

by around 3 mm of play, realistic in terms of incipient bolt

yielding, results in disjunction irreversibility during hydrogen

combustion, at the estimated pressures of 6.2, 6.8, and 8.5 bar

Shim in contact14 mm thk

Shim with 2 mm of play

Flange, head side Flange, penetration side

Shim with 2 mm of play

Clamp in contact

14 mm thk

Double O-ring seal

Figure 10 Mesh of single flange (with shims, clamp, and bolts) and schematic cross-section of both flanges.

Figure 9 Mesh of sleeve/flange system (including shims and bolts) at head end.



2. 6

containment integrity. In addition to the simulations, mock-

up test results were examined by an expert group to define a

seal criterion adapted to the finite element calculations.

Based on the mockup test results, the expert group recom-

mended a containment failure pressure of around 10 bar

absolute (2.25 times the design pressure). The calculations

assume an idealized liner without considering possible weld

defects or corrosion damage, since these phenomena are

very difficult to simulate numerically. A safety coefficient

should therefore be defined based on knowledge of these

defects to determine the containment failure pressure.

Researchers were able to extrapolate results from

the three severe accident scenarios simulated (AF, AS, and

PL) to other severe accident scenarios in the PSA2 project,

given that the mechanical phenomena are mainly pressure-

driven.

For the equipment hatch, depending on the possibilities of

inter-flange sliding, the local and detailed models highlight-

ed two complementary containment failure modes involving

tensile stress and bolt shearing, subject to threshold effects

and dependent on bolt choice and the specific conditions in

each facility (initial play, flange surface condition, friction,

etc.). Regardless of the failure mode, these studies confirmed

the weakness of the current flange closing system, which

uses 33-mm-diameter E24 steel bolts. EDF has decided to

change the grade and diameter of equipment hatch bolts to

increase the accident failure pressure to at least 8 bar

absolute.

Other weak parts of the containment not considered in these

studies, e.g. other penetrations, also need to be investigated

for severe accident conditions, since failure pressure depends

on these elements as well.

Finally, results of the PSA2 project played a key role in the

900 MWe PWR safety assessments prior to the third series of

ten-year inspections, by providing the demonstrations

required to support IRSN's analysis.

absolute for E24, 40 CNDV 0703, and Z6 CNU 17.4 bolts

respectively. Decreasing the cross-section is therefore detri-

mental to bolt strength, whereas greater yield strength con-

tributes to mechanical strength (Figure 11).

Variation related to modeling choices and material character-

istics was found to be around 15%, based on sensitivity studies.

This is much lower than the variation linked to the flange

tightening configuration, in particular flange/bolt play.

conclusion

Assessing integrity of the containment building, which

serves as the third and final barrier, is a critical safety issue,

given that environmental release in a severe accident context

is partly driven by containment leakage.

The wide range of competence applied to this project is evi-

dence of the complexity of the problem and the diversity of the

parameters involved. Furthermore, the innovative multiscale

approach and the structural calculations using non-linear finite

element analysis, beyond the scientific challenge they represent,

also demonstrate that numerical simulations can be used to

assess containment integrity.

The non-linear calculations effectively simulated mechanical

containment behavior in severe accident conditions and

made it possible to detect sensitive points in the structure. In

900 MWe PWR containments, the internal metal liner ensures

Figure 11 Flange surface deformations at 10 bar absolute, amplified 100 times (AF scenarios).

Deformation with E24 bolts

AF PREC 30COL BO_E24

déformée des facesdes brides 90683,193 9

AF PREC 30COM BO_Z6CNUI74_N

déformée des facesdes brides 90683,193

AF PREC 30COL BO_E24_ser69déformée des faces

des brides 90683,193 9

Deformation with Z6CNU17.4 bolts

Deformation with 40CNDV0703 bolts


2. 6

References

B. Cirée, G. Nahas, 2007, Mechanical Analysis of the equipment hatch behaviour for the French PWR 900 MWe under severe accident. h01/3 - Proc. SMiRT, Toronto, Canada.

M.F. hessheimer, R.A. Dameron, 2006, Containment Integrity Research at Sandia National Laboratories. NUREG/CR-6906 SAND2006-2274P.

International Standard Problem No. 48, Containment capacity, 2004, Phase 2 Report Results of Pressure Loading Analysis, Organization for Economic Co-operation and Development, Nuclear Energy Agency, Committee on the Safety of Nuclear Installations. NEA/CSNI/R(2004)11.

International Standard Problem No. 48, Containment Capacity, 2005, Synthesis Report, Organization for Economic Co-operation and Development, Nuclear Energy Agency, Committee on the Safety of Nuclear Installations. NEA/CSNI/R(2005)5/Vol.1, 2 and 3.

G. Nahas, B. Cirée, 2007, Mechanical Analysis of the containment building behavior for the French, PWR 900 MWe under severe accident. h05/5 - Proc. SMiRT, Toronto, Canada.

E. Raimond, B. Laurent, R. Meignen, G. Nahas, Cirée B., 2004, Advanced modelling and response surface method for physical models of level 2 PSA event tree. CSNI-WG-Risk-Workshop level 2 PSA and severe accident management, Köln, Germany.

P. Verpeaux, A. Millard, T. Charras, A. Combescure, 1989, A modern approach of large computer codes for structural analysis. Proc. SMiRT, Los Angeles, USA.



Pascal Guillou William Le Saux

Fire Experimentation Laboratory

Ten countries and two French partners,

EDF and the French armament procurement

agency DGA, are participating in the experi-

mental PRISME program, focused on fire

propagation in elementary, multiple-enclo-

sure scenarios applied to nuclear facilities.

Launched by IRSN in 2005 and coordinated

under the auspices of the OECD, PRISME

consists of several test campaigns con-

ducted over a five-year period. Prisme Door

investigated heat and smoke propagation

through open doors between rooms. This

series of six tests carried out in IRSN's DIVA

facility ended on March 29, 2007. Some

3500 measurements were taken, as

described in the test reports, available to

program partners at the PRISME website.

IRSN has already applied the initial results

as part of a comparative exercise between

the various simulation codes used by the

partners.

Analysis of the results began with heat

release rate calculations. For the first time

in tests of this type, several approaches will

be compared: a "mechanical" approach, in

which the pyrolysis rate is measured using

a specific weighing device, and the effective

heat of combustion is calculated from

results of the Prisme Source test campaign,

conducted under the SATURNE hood; a

"chemical" approach, based on measuring

oxygen, carbon dioxide, and carbon mon-

oxide concentrations; and a "thermal"

approach, based on temperature measure-

ments and wall heat fluxes. The heat release

rate calculations were presented to program

partners at the OECD PRISME meeting held

on October 17 and 18, 2007.

InITIAL RESULTS of the Prisme Door campaign2.7



Georges Repetto, Nathalie Seiler, Valia Guillard and Pierre Ruyer

Laboratory of Studies and Experimental Analysis of Core Degradation

Following the in-core fuel management

changes (higher burnup rates) EDF plans

to implement, IRSN, in its expert capacity,

will be revising studies on Loss-of-Coolant

Accidents (LOCAs), the design-basis acci-

dent involving a line break in the reactor

coolant system. During this type of tran-

sient, water evaporation in the reactor

vessel can cause the fuel rods to dry out

and overheat, potentially resulting in

thermomechanical swelling and failure of

the fuel cladding. This swelling can sig-

nificantly clog part of the core, which can

in turn compromise cooling as safety

systems inject water into the core during

the reflooding phase. Once these systems

are activated, water supplied to the reac-

tor coolant system should gradually fill

the core. An important safety objective is

to make sure clogged zones (filled with

fuel fragments displaced by a process

observed in irradiated fuel experiments)

can be cooled and reflooded.

IRSN teamed with EDF to launch an

R&D program dedicated to LOCAs in PWR

cores containing advanced high-burnup

fuel.This program is based on three lines

of research: building knowledge through

analytical experimentation, simulation

and comprehensive experimentation to

study coupling between phenomena, and

conducting an exhaustive survey to ensure

that all the phenomena involved have

been taken into account.

Grounded in a multiscale, multiphysics

approach, the simulation aspect focuses

on developing two new computing tools:

a three-dimensional fuel bundle simula-

tion code (DRACCAR) for studies directly

supporting expert assessment, and a ther-

mohydraulic simulation code (NEPTUNE)

developed in collaboration with CEA, EDF,

and AREVA, for apprehending elementary

phenomena.

The aim of DRACCAR is to model an entire

fuel assembly, in order to assess clogging

and cooling in severely deformed rods,

taking into account mechanical and ther-

mal interactions between rods.

In the NEPTUNE platform, NEPTUNE-

CMFD (Computational Multiphase Fluid

Dynamics) is a numerical simulation tool

that resolves averaged two-phase ther-

mohydraulic flow equations at local scales

us ing three-dimensional geometry.

Starting in 2006, IRSN took several initia-

tives to prepare for this new tool, begin-

ning with the analysis of existing models,

followed by development and validation

of new, more advanced models. Averaged

equations require significant modeling

work to take into account phenomena

that occur at the inclusion scale (steam

bubble or water droplet).

Research launched by IRSN in this area

includes:

studies to assess NEPTUNE-CMFD capa-

bilities, focusing on:

dispersed flow dynamics through simula-

tion of a series of adiabatic water-air bubble

flow experiments, conducted in the Topflow

facility at Forschungszentrum Dresden-

Rossendorf (FZD), in cooperation with

another German organization, Gesellschaft

für Anlagen und Reaktorsicherheit (GRS);

LOCAL-SCALE THERMOHyDRAULICS R&Dto support LOCA studies2.8



simulation of Debora boiling flow tests

conducted by CEA;

advanced dispersed flow modeling that

takes into account local size variation

(polydispersion) in an inclusion popula-

tion as applied to boiling flows, in part-

nership with the Fluid Mechanics Institute

of Toulouse (IMFT);

a thesis project involving experimental

study, modeling, and numerical simula-

tion of heat and mass transfers resulting

from droplet impact on a hot wall, con-

ducted in collaboration with Lemta

(Laboratory of Theoretical and Applied

Mechanics, CNRS joint research unit in

Nancy);

a proposed thesis project to investigate

two-phase turbulence in the clogged core

zone using the LES model, conducted in

cooperation with the CNRS Promes

Laboratory (Processes, Materials, and

Solar Energy) in Perpignan.


112 Scientific and Techincal Report 2007 - IRSN

Franck ArreghiniFuel in Accident Situations

Studies Department

In early 2007 IRSN began a new

research program, in partnership with

AREVA, to study the thermomechanical

behavior of spent fuel assemblies stored

underwater at the La hague reprocessing

plant, in the event of an accident causing

the emptying of a storage pool.

Scheduled to last three years, this

program will investigate the accident

conditions in which fuel assemblies, con-

ditioned in storage baskets arranged in

relatively tight arrays, may be subject to

overheating, deformation, or even loss

of seal. These phenomena are related to

residual heat, due mainly to fission prod-

ucts created during fuel irradiation in the

reactor. The final goal of this program is

to check that all emergency cooling mea-

sures implemented by the operator in

case of storage pool accidents will in fact

guarantee that stored fuel is not exposed

to unsafe thermal and mechanical

loads.

Thermomechanical studies conducted

using the ICARE-CAThARE code, devel-

oped by IRSN, will concentrate on real-

istic accident scenarios involving partial

dewatering of the fuel assemblies.

A nEW RESEARCH PROGRAMto study fuel assembly behavior at La Hague storage pools in the event of accidental dewatering

2.9

Accidents dans les installations nucléaires


2 . 10

Key dAteS Dissertations defended and other major eventsdISSeRtAtIoNS deFeNded

March 8, 2007 christian mun defended a thesis entitled

"Ruthenium chemistry in containment in a

severe accident situation" (Cadarache,

Bouches-du-Rhône). On June 14, 2007,

Christian Mun received the Jean Bourgeois

award from the French Nuclear Energy Society

(SFEN) for his thesis work.

September 7, 2007 Julien lamome defended a thesis entitled

"Study of steam explosion initiation and

acceleration" (INSTN, CEA Saclay).

November 28, 2007 yannick Pizzo defended a thesis entitled

"Using the number of mass transfers to char-

acterize a diffusion flame along a solid fuel

in natural convection conditions" (IUSTI,

Château-Gombert Research Park, Marseille).

otheR mAJoR eveNtS

May 2007 IRSN was well represented at ICNC 2007

(International Conference on Nuclear Criticality

Safety), in Saint Petersburg, Russia.

June 2007 2nd European Review Meeting on Severe

Accident Research, ERMSAR-07.

This meeting was organized by Forschungs-

zentrum Karlsruhe Gmbh (Germany) and

supported by the European Commission

through the SARNET excellence network on

severe accidents, launched in 2004 as part of

FP6 (Sixth Framework Program for Research

and Technological Development).

July 2007 End of the Pu Criticality and Temperature

experimental program, which aimed to verify

the positive temperature effect in low-con-

centration plutonium nitrate solutions; in

total, 13 approach-to-critical experiments

were conducted from October 2006 to July

2007.

September 2007 IRSN has been tasked by the European

Commission with coordinating the safety

work group of the Sustainable Nuclear Energy

Technology Platform (SNE-TP), an EC initia-

tive to map out the key themes in European

nuclear research between now and 2020.

Specifically, the SNE-TP program aims to

define and implement a Strategic Research

Agenda (SRA), to be drafted by the end of

2008.

The SRA has a matrix structure that includes

thematic work groups in charge of a specific

reactor design or problem, along with four

cross-functional groups on materials, simula-

tion, safety, and fuel. IRSN is coordinating the

safety group.

The agenda will influence the next three

Framework Programs for Research and

Technological Development.

October 2007 The preliminary safety report for CABRI

(IRSN test reactor for fuel safety) was submit-

ted to the French Nuclear Safety Authority.

Factory assembly began for test systems and

equipment, in view of the qualification test

on the pressurized water loop (CIPQ).

December 2007 The European excellence network SARNET

(Severe Accident Network), which is coordi-

nated by IRSN and brings together 52 insti-

tutions or organizations and 350 researchers,

was extended for an additional six months.

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