nuclear loca

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Lappeenranta University of Technology

Faculty of Technology

Department of Energy and Environmental Technology

Large Break Blowdown Test Facility Study

The subject of Masters thesis has been accepted on 27th of February 2008.

Supervisor: Riitta Kyrki-RajamakiInstructor: Heikki Purhonen

Lappeenranta, 26.3.2008

Arto YlonenYlasatamakatu 31 B 1380100 JoensuuFinland

Tel. +358 50 354 4346


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TIIVISTELMA

Lappeenrannan teknillinen yliopistoTeknillinen tiedekuntaEnergia- ja ymparistotekniikan osasto

Arto YlonenIson LOCA:n ulospuhallusvaiheen koelaitteistojen kartoitusDiplomityo200860 sivua, 33 kuvaa, 4 taulukkoa ja 1 liite

Tarkastajat: Riitta Kyrki-RajamakiHeikki Purhonen

Hakusanat: LBLOCA, ulospuhallusvaihe, koelaitteistotKeywords: LBLOCA, blowdown, test facilities

Tyon tarkoituksena on kerata yhteen tiedot kaikista maailmalta loytyvista ison LOCA:nulospuhallusvaiheen tutkimiseen kaytetyista koelaitteistoista. Tyon tarkoituksena onmyos antaa pohjaa paatokselle, onko tarpeellista rakentaa uusi koelaitteisto neste-rakenne-vuorovaikutuskoodien laskennan validoimista varten. Ennen varsinaisen koe-laitteiston rakentamista olisi tarkoituksenmukaista myos rakentaa pienempi pilottikoe-

laitteisto, jolla voitaisiin testata kaytettavia mittausmenetelmia. Sopivaa mittausdataatarvitaan uusien CFD-koodien ja rakenneanalyysikoodien kytketyn laskennan validoi-misessa. Naita koodeja voidaan kayttaa esimerkiksi arvioitaessa reaktorin sisaosien ra-kenteellista kestavyytta ison LOCA:n ulospuhallusvaiheen aikana. Raportti keskittyymaailmalta loytyviin koelaitteistoihin, uuden koelaitteiston suunnitteluperusteisiin se-ka aiheeseen liittyviin yleisiin asioihin. Raportti ei korvaa olemassa olevia validointi-matriiseja, mutta sita voi kayttaa apuna etsittaessa validointitarkoituksiin sopivaa isonLOCA:n ulospuhallusvaiheen koelaitteistoa.


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ABSTRACT

Lappeenranta University of TechnologyFaculty of TechnologyDepartment of Energy and Environmental Technology

Arto YlonenLarge Break Blowdown Test Facility StudyMasters thesis200860 pages, 33 figures, 4 tables and 1 appendix

Examiners: Riitta Kyrki-RajamakiHeikki Purhonen

Keywords: LBLOCA, blowdown, test facilities

The main goal of this work is to gather all suitable LBLOCA blowdown tests togetherand compare them and the results obtained. By reviewing existing measurement data itis easier to decide if a new test facility with a modern measurement system is needed forstructural analysis and computational fluid dynamics (CFD) code validation purposes.Measurement techniques and methods could be first tested with a smaller pilot facility.Suitable data is needed to facilitate the coupling of new CFD codes with structural

analysis codes (fluid-structure interaction analysis). These codes could be used, forexample, to evaluate structural integrity of the reactor core and its supporting struc-tures during a LOCA blowdown. The report concentrates on the existing experimentalknowledge of the blowdown, the design of the test facility and discusses some relatedgeneral issues. This work is not meant to replace any existing validation matrices, butit can be used as an additional source of information when searching for a suitable testfacility for LBLOCA blowdown validation purposes.


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Contents

Nomenclature 3

List of Acronyms 5

List of Figures 6

List of Tables 7

Acknowledgements 8

1 Introduction 9

2 Large Break Loss-of-coolant Accident 10

3 Phenomena during LBLOCA 133.1 Basic thermodynamic phenomena . . . . . . . . . . . . . . . . . . . . . 13

3.1.1 Evaporation due to depressurisation or heat input . . . . . . . . 133.1.2 Condensation due to pressurisation or heat removal . . . . . . . 133.1.3 Friction and pressure drop . . . . . . . . . . . . . . . . . . . . . 143.1.4 Pressure wave propagation . . . . . . . . . . . . . . . . . . . . . 15

3.2 Critical flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 163.2.1 Single-phase flow . . . . . . . . . . . . . . . . . . . . . . . . . . 163.2.2 Two-phase flow . . . . . . . . . . . . . . . . . . . . . . . . . . . 16

3.3 Phase separation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 213.4 Entrainment and de-entrainment . . . . . . . . . . . . . . . . . . . . . 213.5 Stratified flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22

3.6 CCF/CCFL phenomena . . . . . . . . . . . . . . . . . . . . . . . . . . 223.7 Heat transfer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 233.8 Pool formation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 243.9 Quench front propagation . . . . . . . . . . . . . . . . . . . . . . . . . 24

4 Pipe break mechanics 264.1 Longitudinal break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 264.2 Circumferential break . . . . . . . . . . . . . . . . . . . . . . . . . . . . 284.3 Pipe break simulations using Fluent CFD . . . . . . . . . . . . . . . . . 28

4.3.1 Effect of pipe length on depressurisation . . . . . . . . . . . . . 304.3.2 Effect of break opening time on depressurisation . . . . . . . . . 30

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5 Large Break test facilities 345.1 Separate Effects Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . 34

5.1.1 Edwards-OBrien pipe . . . . . . . . . . . . . . . . . . . . . . . 355.1.2 Bartaks pipe . . . . . . . . . . . . . . . . . . . . . . . . . . . . 35

5.1.3 CANON and Super CANON . . . . . . . . . . . . . . . . . . . . 365.1.4 Marviken . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 365.1.5 CSE blowdown experiments . . . . . . . . . . . . . . . . . . . . 385.1.6 Semiscale (Bettis Flask) . . . . . . . . . . . . . . . . . . . . . . 415.1.7 Piper . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 415.1.8 Battelle-Frankfurt RS-16B . . . . . . . . . . . . . . . . . . . . . 425.1.9 Hei Dampf Reaktor (HDR) . . . . . . . . . . . . . . . . . . . . 435.1.10 Other tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46

5.2 Integral Effects Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . 475.2.1 LOFT and SEMISCALE . . . . . . . . . . . . . . . . . . . . . . 47

5.2.2 LOBI . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 49

6 Test facility summary 50

7 Designing the new facility 527.1 Scaling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 527.2 Structures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 537.3 Instrumentation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 557.4 Experiments . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55

8 Conclusions 56

Appendices

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Nomenclature

T the degree of subcooling at inletM mass flowA flow areaa sonic velocity

c specific heatCRF carry-over fractionD diameterE fraction of liquid entrained as dropletsf friction factorG mass flux, mass flow rategc gravitational conversion factorh enthalpyK bulk modulusK friction factor

k slip ratioKl loss coefficientL initial defect lengthL lengthm mass per unit areaN experimental parametern thermal equilibrium polytropic exponentp pressureQ volumetric flowq peak linear heat rates entropy

t timeV core flooding rateV velocityV volumev specific volumev velocityw phase velocityZ quench front level

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Greek symbols void fraction quality critical pressure ratio

isentropic expansion factor cp/cv density stress sector from which the liquid film is removed

Subscripts0 stagnation, upstream reservoir, initial1 inlet1 single-phase3 branch

c criticalE entrainmentE equilibriume equivalentf liquidf water added to coref g evaporationg gasi index for area changel liquidm mixturep constant pressurep fracture propagations constant entropys pipe materialsg steam generationt throatT P two-phasev constant volumey yield

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List of Acronyms

CCF Counter Current Flow

CCFL Counter Current Flow Limitation

CFD Computational Fluid Dynamics

DBA Design Basis Accident

DEGB double-ended-guillotine break

DNB departure from nucleate boiling

ECCS Emergency Core Cooling System

FSI Fluid-structure interaction

HEM homogeneous equilibrium model

HFK Henry-Fauske model

HFM homogeneous frozen flow model

ISP International Standard Problem

LBLOCA Large Break Loss-of-coolant Accident

LWR light water reactor

NRC US Nuclear Regulatory Commission

OECD Organisation for Economic Co-operation and Development

RPV Reactor Pressure Vessel

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List of Figures

2.1 Expansion wave propagation and growth of core barrel deformations [1] 112.2 Phenomena during end of blowdown phase [2] . . . . . . . . . . . . . . 12

3.1 Schematic of flashing transient phenomena during blowdown [3] . . . . 143.2 Critical Pressure Ratio Data as a function of Length/Diameter Ratio [4] 173.3 Differences between introduced two-phase critical flow models . . . . . 183.4 Countercurrent flow of steam and liquid [3] . . . . . . . . . . . . . . . . 233.5 Pool formation in upper plenum during reflood [3] . . . . . . . . . . . . 243.6 Quenching during reflood [3] . . . . . . . . . . . . . . . . . . . . . . . . 25

4.1 Different rupture types considered in High Energy Line Breaks . . . . . 274.2 Predicted opening times of the break in three cases [5] . . . . . . . . . 274.3 Hoop stress in cylindrical shell . . . . . . . . . . . . . . . . . . . . . . . 284.4 DEGB locations, realistic break sizes and times [6] . . . . . . . . . . . . 294.5 Mesh/Geometry for 1.369 m pipe . . . . . . . . . . . . . . . . . . . . . 31

4.6 Arrival of the rarefaction wave to the vessel in Case 2 (0.5 ms) . . . . . 314.7 Pressure at outlet in Case 1 simulations . . . . . . . . . . . . . . . . . . 324.8 Effect of pipe length on depressurisation . . . . . . . . . . . . . . . . . 324.9 Pressure at outlet in Case 2 simulations . . . . . . . . . . . . . . . . . . 334.10 Effect of break opening time on depressurisation . . . . . . . . . . . . . 33

5.1 Schematic of Edwards-OBrien blowdown pipe [7] . . . . . . . . . . . . 355.2 Schematic of Bartaks horizontal experimental channel [8] . . . . . . . . 365.3 Schematic of CANON facility [9] . . . . . . . . . . . . . . . . . . . . . 375.4 Schematic of Super CANON facility [10] . . . . . . . . . . . . . . . . . 375.5 Schematic of Marviken critical flow test facility [11] . . . . . . . . . . . 39

5.6 Schematic of CSE reactor simulator [12] . . . . . . . . . . . . . . . . . 405.7 Schematic of the vessel used for SEMISCALE test 711 [13] . . . . . . . 415.8 Schematic of Piper vessel and blowdown nozzle [14] . . . . . . . . . . . 425.9 Schematic of RS-16B blowdown facility and models[15] . . . . . . . . . 435.10 Schematic of HDR pressure vessel and core barrel [16] . . . . . . . . . . 445.11 Instrumentation of the long and short discharge pipes in HDR [17] . . . 465.12 LOFT test facility [18] . . . . . . . . . . . . . . . . . . . . . . . . . . . 475.13 LOFT reactor vessel instrumentation [18] . . . . . . . . . . . . . . . . . 485.14 Semiscale Mod-1 (left) and Semiscale Mod-2C (right) [18] . . . . . . . . 485.15 LOBI test facility [18] . . . . . . . . . . . . . . . . . . . . . . . . . . . 49

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List of Tables

4.1 FLUENT simulation settings . . . . . . . . . . . . . . . . . . . . . . . . 30

5.1 HDR preliminary and main phase test matrices . . . . . . . . . . . . . 45

6.1 Test facilities summarized . . . . . . . . . . . . . . . . . . . . . . . . . 51

7.1 Time-reducing scaling law [19] . . . . . . . . . . . . . . . . . . . . . . . 54

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Acknowledgements

I would like to express my gratitude to professor Riitta Kyrki-Rajamaki and my su-perior Heikki Purhonen for their support and guidance during the writing process ofthis thesis. I would also like to thank my family and friends for their support throughthe studies. Also big thanks belongs to all of my colleagues and especially to MarkkuPuustinen and Vesa Tanskanen for their support and expertise.

Lappeenranta, 26.3.2008

Arto Ylonen

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1 Introduction

Large Break Loss-of-coolant Accident (LBLOCA) is a Design Basis Accident (DBA) inlight water reactor (LWR). The design basis accident scenarios define the operationaldemands of the safety systems. No matter how unlikely these accident types may be,all the systems must be designed so that the safety of the plant is as high as reasonablyachievable. LBLOCA was for a long time considered the worst case scenario, but after

the Three Mile Island accident (1979) it was realized that also small and intermediatebreaks could cause very severe damage to the reactor core. Therefore the researchersmoved their focus to the smaller breaks for decades. However, the development of newcomputer codes for nuclear safety analysis has invoked the interest on LBLOCA issuesagain.

Large break experiments done in the 60s and 70s were mainly designed to studythe coolability and the long term cooling of the core. All main phases of the LBLOCAwere studied: blowdown, refill and reflood. Only a few test facilities were used to studythe first phase of the loss-of-coolant accident, the blowdown. Pipe rupture phenomena,depressurisation, critical flow, propagation of the pressure waves and possible defor-

mation of the core barrel are the main subjects of interest during the blowdown phasewhen designing adequate emergency cooling systems and guaranteeing the safety of theplant. [20, 21]

The main goal of this work is to gather all suitable LBLOCA blowdown tests to-gether and compare them and the results obtained. By reviewing existing measurementdata it is easier to decide if a new test facility with a modern measurement system isneeded for structural analysis and Computational Fluid Dynamics (CFD) code val-idation purposes. Measurement techniques and methods could be first tested witha smaller pilot facility. Suitable data is needed to facilitate the coupling of new CFDcodes with structural analysis codes (Fluid-structure interaction (FSI) analysis). These

codes would be used, for example, to evaluate structural integrity of the reactor coreand its supporting structures during a LOCA blowdown. To give some kind of over-all view of the LBLOCA the most important phenomena are described. The reportconcentrates on the existing experimental knowledge of the blowdown and the designof the test facility and discusses some related general issues. This work is not meantto replace any validation matrices, but it can be used as an additional source of infor-mation when searching for a suitable test facility for LBLOCA blowdown validationpurposes. [20]

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2 Large Break Loss-of-coolant Accident

The LBLOCA is a DBA in light water reactors and the double-ended-guillotine break(DEGB) of the largest primary piping system is limiting condition for the EmergencyCore Cooling System (ECCS) requirements. Although a LBLOCA is considered veryunlikely to happen, the safety systems must be designed to secure adequate cooling ofthe reactor core during the accident to prevent the meltdown of the core. Main course

of a LBLOCA is described next. [20]

In numerical analyses hypothetical loss of coolant accident is most often assumed totake place in the cold leg of the primary circuit. Sudden rupture of a pipe in the primarycircuit leads to the rapid depressurisation of the system, this is called the blowdownphase. During the depressurisation high-pressure coolant in the primary circuit is incontact with the low-pressure containment atmosphere. As a result so called rarefactionor expansion wave propagates through the primary circuit and the Reactor PressureVessel (RPV) (Figure 2.1). From the first milliseconds to approximately 50 ms of theaccident the depressurisation wave propagates in single-phase liquid flow (subcooledblowdown), later on in two-phase flow (saturated blowdown). The coolant flow in thereactor core is reversed in the beginning of the blowdown phase, but as the break flowbecomes saturated the flow in the core is restored. This takes only a couple of secondsfrom the beginning of the accident. The velocity at which the pressure wave propagatesthrough the system depends on the compressibility of the fluid and the elasticity of theprimary circuit piping and structures. Forces caused by the pressure waves are thehighest in the beginning of the blowdown phase. Break opening time has a significantinfluence on the blowdown forces. When decreasing the break size and increasing thedistance from the break, forces decrease. Loads induced by the blowdown forces couldcause deformation of the core barrel and fuel bundles which could lead to ineffectivecooling of certain parts of the core and end up to the partial meltdown of the reactor

core. [20, 21, 22, 23]

Critical flow rate, the highest rate at which the coolant leaves the circuit, is limitedby the sonic velocity or pressure wave propagation velocity. When the coolant is insubcooled state, the velocity is limited by the liquid sonic velocity. When the coolantreaches saturated state and flow at the break is in two phases, the sonic velocity isreduced significantly. As the break flow velocity equals the sonic velocity, the flow issaid to be choked or sonic flow. During the saturated blowdown the pressure wavesbecome damped and the coolant exits the primary circuit at critical flow rate in twophases. Damping results from the equalities and the opposite directions of the blow-

down pressure wave velocity (rarefaction waves) and break flow velocity. Pressure wave

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Figure 2.1: Expansion wave propagation and growth of core barrel deformations [1]

propagation velocity drops sharply when flashing begins and vapor is present. Velocityis strongly dependent on the void fraction . [3, 24, 25]

As the primary circuit pressure drops, boiling front propagates through the coolingsystem starting from the hottest regions of the reactor core (upper core and upperplenum). As a result of increasing void fraction, neutron moderation weakens andfission process stops. Reactor scram is triggered when certain signals happen (lowpressure or low coolant level etc). Though fission is stopped, the decay heat andgeneration of void result in departure from nucleate boiling (DNB). This could leadto high fuel cladding temperatures. The maximum allowed peak cladding tempera-ture for Zircaloy-clad is defined by US Nuclear Regulatory Commission (NRC) in 10CFR 50.46 regulation as 2200F (1204C). When cladding temperature increases above980C, zircaloy and steam reaction produces heat. As the cladding temperature ex-ceeds 1100C, the amount of heat generated in zircaloy-steam reaction is even moresignificant than the decay heat. [3, 20, 22]

After the reactor scram the primary system pressure continues decreasing. Whenit has dropped below the accumulator pressure, emergency coolant is discharged intothe primary system. As a result of the blowdown, the flow in the reactor core has beenreversed again. Steam water counter-current flow blocks ECC flow in downcomer and

turns it towards the broken loop instead of the reactor core (Figure 2.2). Heat storedin the RPV walls boils ECC water and strengthens the bypass phenomenon. In a coldleg break more emergency core coolant is needed than in a hot leg break, partitiallybecause of this phenomenon. When the break is located in the hot leg similar counter-current phenomenon doesnt happen. As the pressure difference between the vesseland containment decreases the amount of steam generated due to flashing is decreased.ECC water is now able to counter the upward steam flow and enter the lower plenumand the core, refill of the RPV starts. The second peak in fuel cladding temperatureoccurs now as the steam flow through the core decreases. As the water level reachesthe bottom of the core, reflood of the reactor core begins. ECC water gradually rewets

the fuel bundles and heat transfer increases. All two-phase flow regimes are presentin the core during the reflood period. Heat transfer rate strongly depends on the flow

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Figure 2.2: Phenomena during end of blowdown phase [2]

regime. After the core is fully reflooded, long-term cooling must be arranged to removethe decay heat. [3, 20, 26]

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3 Phenomena during LBLOCA

There are many complicated phenomena present during a LBLOCA. Accurate mod-elling of these phenomena is needed to simulate the overall behaviour of the properreactor system. These two-phase flow phenomena are described in this chapter. How-ever, the main focus is on the equations and developed models, because quite completeoverall description of the phenomena already exists. This characterisation can be found

from OECD Code Validation Matrix. As the subject of this study is the blowdown testfacilities it is reasonable to emphasize those particular phenomena which are presentin the beginning of the LBLOCA. The idea is to introduce relevant phenomena asbackground for the test facility study. [27, 28]

Phenomena during the LBLOCA can be basically divided in nine main categories:basic thermodynamic phenomena, critical flow, phase separation, entrainment and de-entrainment, stratification, CCF/CCFL phenomena, heat transfer, pool formation andquench front propagation. [28]

3.1 Basic thermodynamic phenomenaBasic phenomena consist processes of evaporation and condensation, friction and pres-sure drops and pressure wave propagation. One could say that they are present duringmore complex phenomena. [27]

3.1.1 Evaporation due to depressurisation or heat input

As the system is depressurised and the pressure drops rapidly, saturation temperatureof the fluid decreases. If the system is initially at saturated state and depressurisationstarts, fluid flashes instantaneously. In the beginning of the LBLOCA depressurisationwave propagates through the primary system and fluid flashes starting from the hottestregions of the reactor core.

During the LOCA the other form of evaporation is also observed. Fluid evaporatesas a result of the heat input. Hot walls of the reactor pressure vessel or uncovered hotfuel bundles evaporate water and steam is generated. Evaporation phenomena duringblowdown are illustrated in Figures 2.2 and 3.1. [3, 27]

3.1.2 Condensation due to pressurisation or heat removal

As the pressure rises over saturation pressure, the steam starts to condensate. Similarcondensation follows when temperature drops below the saturation temperature due

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Figure 3.1: Schematic of flashing transient phenomena during blowdown [3]

heat removal. Some condensation phenomena can be seen in Figure 3.1. [27]

3.1.3 Friction and pressure drop

Friction and pressure drops have great importance in nuclear safety. Interfacial fric-tion, wall to fluid friction and pressure drops at discontinuities are the main types.Modelling of these phenomena is based on global measurements like pressure differ-ence, pressure and temperature. Correct modelling is needed to simulate many othertwo-phase phenomena.

Pressure drops and friction factors are easily obtained in case of single phase flow.Pressure drop in a flow channel can be calculated from the equation

Pfriction =

f L

De

V2

2gc+i

Ki

V2i2gc

(3.1)

Friction factor f can be solved from the well-known Moody friction factor chart as a

function of Reynolds number and relative roughness of the flow channel wall. Differentfactors K for contractions and expansions can be found from literature. Pressuredrop in different locations like bends and valves can be calculated in the same manneras pressure drop by flow area change, with loss coefficient Kl or as a pipe that hasequivalent pressure drop with length Le

Pfriction = KlV2i2gc

=

f LeDe

V2

2gc(3.2)

In case of two phase flow, calculation of the pressure drop is more complex and thereforeseveral correlations have been developed for different purposes. [22, 27]

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3.1.4 Pressure wave propagation

The main interest in this work is to determine how to study experimentally the effectsof pressure waves on the reactor core barrel during the LBLOCA blowdown. The

pressure waves may cause deformation of the core barrel which could result to weakenedcoolability of the reactor core or make the insertion of the control rods impossible.Therefore it is important to know the effects to ensure the safety of the reactor. [27]

3.1.4.1 Single-phase flow

In a single-phase case the propagation velocity of the pressure waves is simply thesonic velocity. Pressure pulse for the fluid velocity change is derived from momentumequation and as a result we get

p2 p1

V2 V1= a1 (3.3)

Sonic velocity for the single-phase fluid is defined using bulk modulus Kwhich describessubstances resistance towards compression. For example bulk modulus for water inatmospheric pressure and 20C is 2.19 109 Pa. Equation for the single-phase sonicvelocity is formed as

a1 =

K

=

V dp

dV=

1

2

dp

dv

s

(3.4)

The sonic velocity in water decreases as temperature rises. Increase in the pressure hasopposite effect on the sonic velocity. [22, 29]

3.1.4.2 Two-phase flow

In a two-phase case the calculation of pressure wave propagation velocity is more com-plicated as it is strongly dependent on the void fraction and two-phase flow pattern.Therefore, many different models have been developed during the years. One simplifi-cation is to assume two-phase flow as homogeneous mixture and then calculate averagedensity from equation

= g + (1 )l (3.5)

The frozen two-phase sonic velocity is derived by substituting equation (3.5) withequation (3.7). The propagation velocity in two-phase fluid is derived from sonic ve-locities in each phase as

a2TP =

ga2g+

(1 )

la2l

1

(3.6)

The equation (3.6) can be used when void fraction is low (Grolmes and Fauske). As itwas said earlier, the propagation velocity depends on the flow pattern or to be exacton the slip ratio. Complexity of the propagation velocity in two-phase medium makesit difficult to model two-phase flow systems. [22, 30]

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3.2 Critical flow

The flow is said to be critical, when the flow rate is as high as physically possible.It means that the downstream conditions dont affect the discharge rate anymore.

Critical flow or choked flow can also be defined in single-phase system with sonicvelocity, choking velocity equals sonic velocity. This means that pressure signals cantbe transmitted to higher pressure upstream anymore, because transmission speed cantexceed sonic velocity. In single-phase case the sonic and critical velocities are farmore easier to calculate than in two-phase case. Critical mass flow rate can be easilyobtained. When the flow is in two phases, the calculation procedure is quite different.In this chapter the calculation of the critical flow rate is described and some developedcritical flow models are introduced. Critical flow is an important topic, because itdetermines the severity of the LOCA. The effect of break opening time and breaklocation on critical flow and depressurisation will be discussed in the chapter 4. [21,

22, 27]

3.2.1 Single-phase flow

In case of the single-phase flow, sonic velocity and critical mass flow rate are simplyrelated as

a21 =gcv

2

(dv/dp)s, (3.7)

where a1 is sonic velocity, gc gravitational conversion factor, v specific volume and s in

subscript indicates that derivate is evaluated at constant entropy. Mass flow rate canbe calculated from

G =Q

A, (3.8)

Now by substituting volumetric flow Q and flow area A with sonic velocity using thedefinition of specific volume, we get

G2c = a21

2 =gc(1/)

22

(dv/dp)s=

gc (dv/dp)s

. (3.9)

Critical mass flow rate tells mass flow in time unit per area.[22]

3.2.2 Two-phase flow

Several two-phase critical flow models have been developed during the last 60 years.Couple of them are introduced here. More information can be found from CSNI re-port about two-phase critical flow models [31]. Two-phase models can be dividedinto two main categories, thermodynamical equilibrium and non-equilibrium models.Thermodynamical equilibrium means that pressure and temperature of the liquid andvapor phases are equal. Pressure and temperature are also linked by the Mollier dia-gram. The other way is to divide them to homogeneous and non-homogeneous models.

Homogeneity means that theres no slip between phases (phase velocities are equal,

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Figure 3.2: Critical Pressure Ratio Data as a function of Length/Diameter Ratio [4]

k = 1). The question is which model should be applied in each case. Homogeneousequilibrium model is not a bad way to calculate critical flow rates in long pipes whereequilibrium state can be achieved. When pipe is short more sophisticated modelsfor non-equilibrium flows must be applied to improve accuracy. Length to diameterratio is used to estimate the type of the critical flow (Figure 3.2). Mechanical andthermodynamical differences of later introduced critical flow models are presented in

simplified scheme (Figure 3.3). Phases are presented separately only for visual reasons.[21, 22, 31, 32]

3.2.2.1 Homogeneous equilibrium model (HEM)

The simplest analytical model is so called homogeneous equilibrium model (HEM)(Starkman et al. (1964)). The model assumes that velocities are equal for both phases,thermodynamical equilibrium exists and flow is isentropic and stationary. Also, asthe name says, fluid is considered as a homogeneous mixture whose thermodynamicalproperties are averages between the two phases. Subscript 0 refers to reservoir con-ditions or stagnation conditions. Flow would achieve stagnation values if it would be

brought isentropically to rest. Therefore, it can be said that the stagnation conditionsupstream are directly the initial conditions if the flow is at rest before the break. Thebasic balance equations for this model are

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Figure 3.3: Differences between introduced two-phase critical flow models

mwA = const. (3.10)

1

2w2 + h = h0 (3.11)

h = hf(p) + hfg(p) (3.12)

1m

= vm = vf(p) + vfg (3.13)

s0 = sf(p) + sfg(p) (3.14)

The critical mass flow rate can be solved by substituting equations (3.12), (3.13) and(3.14) to energy equation (3.11). Now by applying the equation (3.9), we get

Gc =1

v[2 (h0 h)]

1/2

=[2 (h0(p0) hf(p) (s0(p0) sf(p))/(sfg(p))hfg(p))]

1/2

vf(p) + (s0(p0) sf(p))/(sfg(p))vfg(p) (3.15)

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The model can be applied for long pipes, but it has some restrictions. It might give agood estimate for critical flow rates, but it predicts poorly the pressure at the blowdownpipe outlet (Moody). [22, 29, 31, 33]

3.2.2.2 Slip equilibrium model (Moody)

One well-known and widely used model was developed by Moody (1965). In his non-homogeneous equilibrium model he assumed thermodynamic equilibrium and adiabatic,frictionless and stationary flow. Void fraction is defined as

=1

1 + k 1

vfvg

(3.16)

For mass conservation Moody formed

G =

wg

vg =1

1

wf

vf (3.17)

and for energy conservation

h0 =

hg +

w2g2

+ (1 )

hf +

w2f2

(3.18)

All other equations are similar with the homogeneous equilibrium model. From theequations it can be derived that critical flow rate is

Gc =

2 h0 hf hfgsfg (s0 sf)k(sgs0)

sfgvf +

s0sfsfg

vg

2 s0sfsfg

+ sgs0k2sfg

1/2

(3.19)

Slip ratio k at choking point is assumed as

k =

vgvf

1/3(3.20)

Exponent 1/3 is obtained from data comparison. Fauske assumed exponent as 1/2.[24, 31, 34]

3.2.2.3 Homogeneous frozen flow model (HFM)

Starkman et al. (1964) developed one frequently in literature referred two-phase flowmodel. In their homogeneous frozen flow model (HFM) it is assumed that averagephase velocities are equal. Also no heat and mass transfer between the phases isassumed, therefore quality remains constant throughout the expansion. Isentropicallyexpanding vapor behaves like a perfect gas. Kinetic energy is due only to the vaporexpansion. Critical flow rate is defined by gas-dynamic principles. Applying all thebefore mentioned assumptions, equation for the critical flow rate can be formed as

Gc =

1

v

20vg0p0(

1)(1

1

)1/2

(3.21)

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The specific volume for two-phase condition is defined as

v = (1 0)vf0 + 0vg01/ (3.22)

In the equation (3.22) critical pressure ratio is denoted with which is

=

2

+ 1

1

(3.23)

when (1 0)vl0/(0vg0) 0.14

For the critical mass flow rate they formed the equation

Gc =

0vg

np+ (vg vl0)

(1 0)N

sgE slE

dslEdp

0cpg(1/n 1/)

p(sg0 sl0)

1/2

(3.25)

However, equation (3.25) contains two unknowns, critical pressure pc and critical massflow rate Gc. Thermal equilibrium polytropic exponent n is defined as

n =(1 )cl/cpg + 1

(1 )cl/cpg + 1/(3.26)

Therefore additional equation needs to be used in order to get solution for the problem.Momentum equation was formed

(1 x0)vl0(p0 pt) +0

1[p0vg0 ptvgt] =

[(1 0)vl0 + 0vgt]2

2G2c (3.27)

Solution for the critical mass flow rate can be obtained from equations (3.25) and(3.27). Pressure ratio can be expressed as

= 10

0

(1 ) +

1122t

+ 1

1(3.28)

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where

=pt

p0

(3.29)

=

1

n+

1

vl0vgt

(1 0)Npt

0(sgE slE)t

dslEdp

t

cpg(1/n 1/)

(sg0 sl0)

(3.30)

0 =0vg0

(1 0)vl0 + 0vg0(3.31)

t =0vg0

(1 0)vl0 + 0vgt(3.32)

vgt = vg0

1

(3.33)

This model can be used in cases of dispersed liquid and vapor flow in nozzles. If

theres separation of phases, it cant be used. Solution requires only the knowledge ofthe geometry and the upstream stagnation conditions. Henry and Fauske applied themodel also for orifices and short tubes. [22, 33]

3.3 Phase separation

During the LOCA it is important to know the quality of the flow in the break location.It determines the mass and energy removal from the system. Phase separation inbranches depends on the quality of the flow in inlet and on flow pattern. Experimentshave been conducted for different types of branches like Ts and Ys. Also gravity and

inertia forces cause phase separation. Phenomenon has significant influence on heattransfer and quenching of the core during LBLOCA, therefore accurate modelling isneeded.

For T-branches Whalley and Azzopardi proposed a correlation for the annular flowof high velocity in the horizontal and vertical lines.

3/1 =1

2( sin )/(M3/M1) (3.34)

where the angle defines the sector from which the liquid film is removed

= 360

M3(1 3)

M1(1 1)(1 E1)

1.2(D3/D1)

0.4

(3.35)

and M is mass flow, E1 fraction of the entrained liquid in inlet and D is diameter of theconduit. However, the correlation can only be verified with low-pressure data. [22, 27]

3.4 Entrainment and de-entrainment

Entrainment is a term for the phenomenon in which water droplets separate from the

liquid phase to the vapor phase. After entrainment droplets either fall back (horizontalflow), deposit in the liquid phase (vertical flow) or end up being carried over by the

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vapor. Carry-over is a term for the water droplets entrained in the vapor phase. Carry-under is used for the opposite case in which the vapor bubbles are entrained in theliquid phase. De-entrainment is the inverse phenomenon of entrainment, entrainedwater droplets are removed from the vapor. Entrainment and de-entrainment have big

influence on heat transfer in the core during LOCAs. Phenomena also have effect onECC performance during an accident. [22, 27]

These phenomena have been studied in different cases and some correlations havebeen developed for the calculations during the years. However, the lack of detailedinformation makes it possible only to form global correlations. One example of thecorrelations is the one developed for the reflood phase of the LOCA. During the refloodsignificant amount of water is carried over as the ECC water evaporates because ofdecay heat. When the average steam velocity is more than the carry-over velocity, thefollowing correlation can be used to estimate the entrainment

ME = (CRF)

Mf Msg (3.36)

where CRF is carry-over fraction, Mf mass added to core and Msg steam generatedby the decay heat. Empirical CRF factor is formed as

CRF = 0.9553 exp0.0013935(p/60)2

{1 exp[4.3127(q/1.24)]}

{1 exp[16.62(V /6)]} exp0.037161(T /140)2

1 + sin 3.1416

0.5 + (Z/6)0.025733(6/V)+0.13553(T/140)+0.0656(P/60)

2

(3.37)

where pressure is in psia, q peak linear heat rate in kW/ft, V flooding rate in in./s,subcooling at inlet in F and quench front level Z in ft.

As can be seen from the multiplication factors and exponents the correlation ishighly empirical and it cant be used in any other case than ECCS evaluation duringLOCA reflood. Other correlations can be found from literature, but all of them aremore or less experimental just like the one presented above. [22]

3.5 Stratified flow

Flow stratification is a term for the phenomenon in which gravitational forces haveseparated the phases in the flow channel. Frictional forces and heat transfer are differentfrom homogeneous flow, because the liquid phase interacts only with a part of the flowchannel walls. Phase separation in the branches can also be significantly different. [27]

3.6 CCF/CCFL phenomena

Counter Current Flow (CCF) is a phenomenon in which the subcooled emergency

core coolant fluid is in contact with the hot walls of the reactor pressure vessel duringLBLOCA refill phase. Steam is generated which rises upwards and reduces penetration

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Figure 3.4: Countercurrent flow of steam and liquid [3]

of ECC fluid to the core. Counter current flow phenomenon is presented in Figure 3.4.

Counter Current Flow Limitation (CCFL) sets limits for the ECC penetration anddetermines how effectively the reactor pressure vessel is refilled and how much of theECC fluid is bypassed directly to the broken loop. As the steam generation decreasesthe ECC fluid starts to fill the lower plenum. [3]

3.7 Heat transfer

Correct modelling of the heat transfer from fuel rods to coolant is essential whenpredicting the fuel and the peak cladding temperatures during the transients and ac-

cidents. Accurate modelling of the heat transfer during the LBLOCA blowdown is aquite complicated task due to different boiling modes. In the beginning of the blow-down the mode of the heat transfer is still single-phase convection and Dittus-Boelterand Rohsenow-Choi correlations can be used. As the nucleate boiling starts Chenscorrelation can be applied. Hsu and Beckner developed correlation for critical heatflux during rapid depressurisation. For heat transfer calculations after critical heatflux there are couple of correlations. Tong and Young correlation for transition fromcritical heat flux to film boiling. During the film boiling correlation by Bishop et al. ormodified Condie-Bengston can be applied. As the LBLOCA blowdown is very rapidevent, the accuracy of the heat transfer correlations is not so critical issue. Conserva-tive and underpredicting correlations that lead to the highest temperatures of the coreare well-founded if there are no other appropriate correlations available. [22]

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3.8 Pool formation

During the reflood phase of LBLOCA water droplets may entrain in the core regionand then again de-entrain at the upper plenum support plate. This phenomenon can

cause two-phase pool formation at the upper plenum (Figure 3.5). In some reactordesigns emergency core coolant is also injected to the upper plenum. Due to the steamgenerated in the reactor this water may form a pool at the upper support plate of thecore. This phenomenon is also referred as CCFL. [3]

Figure 3.5: Pool formation in upper plenum during reflood [3]

3.9 Quench front propagation

Several correlations have been derived to calculate quench front propagation and rewet-ting of the fuel pins. Some of the parameters that have effect on the quench frontpropagation are wall temperature, coolant temperature and subcooling degree, pres-sure, wall material etc. Couple of the most used parameters in correlations are Pecletand Biot numbers, and dimensionless temperature. Effect of a single parameter on thisphenomenon is usually hard to observe. Also spacer grids have effect on quenching(Figure 3.6). [3, 36]

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Figure 3.6: Quenching during reflood [3]

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4 Pipe break mechanics

Break characteristics determine the consequences of the pipe rupture. Break charac-teristics consist of fracture orientation, break opening time and break area. Normallybreak is estimated to open in 1 millisecond to have certain conservatism in the safetycalculations. Amplitude of the rarefaction wave is a function of the opening time.Therefore, when the break opening time is longer the dynamic loads are reduced, but

exist for a longer period of time. [5]

4.1 Longitudinal break

Baum summarizes in his article effects of different break characteristics [5]. He describestests with the gas-pressurised pipe to determine the opening time of the axial break. Heintroduces a simple model to predict the minimum break opening time. Using the initialcrack length and breach tip velocity ratio L/vp (three cases) and the characteristic time (equation (4.4)) for the rupture process, the opening time can be predicted (Figure4.2). In the figure the opening time is presented as the function of the characteristictime. Baum describes the charateristic time as a term which presents the effect ofthe pipe diameter and how the pipe is stressed (hoop stress). For smaller and highlystressed pipes opening times are naturally much shorter. [5]

The break opening time to full area is derived from the idealised breach. Breacharea as a function of time can be expressed as

A =2P0vpt

3

3m+

P0Lt2

m(4.1)

in which the first term presents area generated by the movement of the crack tip and the

second term area generated by the horizontal movement of the free edges. When initialdefect length L is large it has more effect on the break area growth than the crack tippropagation has until t > (3L)/(2vp). The ductile fracture propagation velocity vp canbe defined as the propagation velocity of a displacement wave in a plastic membrane

vp = (y/s)1/2 (4.2)

The hoop stress before the rupture is 0 = P0R/h in which h is the wall thickness andR is the radius of the pipe (Figure 4.3). Mass per unit area is m = sh. Substitutingthese assumptions to (4.1) and using full break area (200%) as A, we get equation

23 = 23

t3 + (L/vp)t2 (4.3)

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Figure 4.1: Different rupture types considered in High Energy Line Breaks

0.0001

0.001

0.01

0.0001 0.001 0.01

Openingtime(s)

alpha (s)

00.3 ms

2 ms

Figure 4.2: Predicted opening times of the break in three cases [5]

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Figure 4.3: Hoop stress in cylindrical shell

in which

= (R/vp)(y/0)1/3 (4.4)

The equation introduced by Baum (4.3) is simple and quite well-founded way toapproximate the minimum break opening time. The model suggests smaller breakopening times than were measured in the large scale experiments. The reason for thisis the conservative assumption that the rupture pressure P0 remains the same. Baumsummarizes that break opening times are larger than 1 ms when exceeds 0.5 ms. [5]

4.2 Circumferential break

Circumferential break is more familiarly known as a guillotine break. The break is con-servatively assumed to open instantaneously (1 ms) in the safety calculations. However,there has been a discussion that the conservative assumption of 1 milliseconds is un-realistic and instead of it the best estimate break opening time could be used. Someassumptions for the realistic break opening time in different locations of the primarycircuit can be found in Figure 4.4. Instead of using the conservative assumption thebreak opening time of 20 ms would be used in the case of cold leg guillotine break. [6]

4.3 Pipe break simulations using Fluent CFD

Series of simulations with Fluent CFD software were done for this report. Purposeof these simulations is mainly to illustrate the effect of different break parameterson depressurisation. Using of Fluent to LBLOCA blowdown simulations may soundquestionable. However, in the beginning of the blowdown the flow is mainly in one-phase. Therefore it is possible to use Fluent for the sensitivity studies.

Preliminary simulations were done to test the effect of different solver settings and

calculation meshes. After these simulations the solver settings for the final simulationswere decided. The pressure based solver was found out to be the fastest and the results

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Figure 4.4: DEGB locations, realistic break sizes and times [6]

were the same as with the density based solver. Coupled method was chosen as thepressure-velocity coupling scheme. Realizable k- turbulence model was applied to thesimulations. Courant number was calculated using the speed of sound. Due to relativelylarge speed of sound small time steps had to be used (t = 3 106 s). Because of smalltime steps the simulation time was selected as 6 milliseconds. However, the reflectionof the rarefaction wave from the constant pressure inlet boundary condition happensmuch earlier. Pressure recovery in the vessel is a result of this issue. Two main cases

were simulated using Fluent. The effect of pipe length on depressurisation was studiedin Case 1 and the effect of break opening time in Case 2. Three different geometrieswere used in Case 1. In total 6 different Fluent cases were simulated. The center line ofthe pipe was defined as the rotational axis (x-axis). One of the geometries is presentedin Figure 4.5. As an example, arrival of the rarefaction wave to the vessel is presentedin Figure 4.6. Other geometries were similar with an exception that the pipe lengthwas different. The pipe radius (100 mm) and the downcomer width (140 mm) andlength (1000 mm) were the same in all geometries. The location where pressure acrossthe core barrel wall was measured is marked in Figure 4.5 with a red circle (referredin figures of curves as KP9). General and case specific settings of the simulations are

summarized in Table 4.1.

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Table 4.1: FLUENT simulation settings

Solver settingsSolver Pressure Based, Axisymmetric

Pressure-Velocity Coupling CoupledFormulation and Time Implicit, UnsteadySimulated time 6 millisecondsTime Step 3e-06 sIterations per Time Step 150Discretization Pressure: 2nd Order

Density: 2nd Order UpwindMomentum: 2nd Order UpwindTurbulent Kinetic Energy: 2nd Order UpwindTurbulent Dissipation Rate: 2nd Order Upwind

Velocity formulation AbsoluteUnsteady formulation 1st-Order ImplicitGradient Option Green-Gauss Cell BasedViscous Model Realizable k-, Standard Wall Functions

Case specific settings Case 1 Case 2Mesh size [cells, quadrilateral] 15627, 22819, 28039 22819Initial Gauge Pressure [bar] 109 109Inlet Gauge Pressure [bar] 109 109Outlet Gauge Pressure [bar] 31.99, 52.39, 57.29 52.39Pipe length [m] 0.5, 1.369, 2 1.369

Break Opening Time [ms] 1 0.5, 1, 1.5, 2Water Temperature [K] 541 541

Water propertiesDensity [kg/m3] Pressure variable (UDF)Viscosity [kg/ms] Constant, 9.998e-05Speed of Sound [m/s] Pressure variable (UDF)

4.3.1 Effect of pipe length on depressurisation

Setting of the outlet boundary condition was the most difficult part in these simulations(Figure 4.7). It was decided to set the outlet pressure as the pressure in the criticalflow situation (Fauske diagram, Figure 3.2). Other option would have been to set breakopen to atmospheric pressure. In Case 1, the pressure difference across the core barrelwall behaved as could be expected (Figure 4.8). When the pipe is shorter, the pressuredifference is greater and the depressurisation of the vessel begins earlier.

4.3.2 Effect of break opening time on depressurisation

The effect of break opening time was studied in Case 2. The HDR geometry (Figure

4.5) was chosen as the simulation case and the outlet boundary condition was differentin each simulation (Figure 4.9). The pressure difference across the core barrel wall

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Grid (Time=0.0000e+00) FLUENT 6.3 (axi, pbns, rke, unsteady)Feb 27, 2008

Figure 4.5: Mesh/Geometry for 1.369 m pipe

Contours of Absolute Pressure (pascal) (Time=1.6470e-03) Mar 06, 2008FLUENT 6.3 (axi, pbns, rke, unsteady)

1.10e+07

5.34e+065.53e+065.72e+065.91e+066.09e+066.28e+066.47e+066.66e+066.85e+067.04e+067.23e+067.42e+067.60e+067.79e+067.98e+068.17e+06

8.36e+068.55e+068.74e+068.93e+069.11e+069.30e+069.49e+069.68e+069.87e+061.01e+071.02e+071.04e+071.06e+071.08e+07

Figure 4.6: Arrival of the rarefaction wave to the vessel in Case 2 (0.5 ms)

behaved as could be expected (Figure 4.10). When the break opening time is shorterthe shape of the rarefaction wave becomes much steeper. This results in greater pressuredifferences across the wall.

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3e+006

4e+006

5e+006

6e+006

7e+006

8e+006

9e+006

1e+007

1.1e+007

0 0.0005 0.001 0.0015 0.002 0.0025 0.003 0.0035

[Pa]

TIME [s]

Outlet pressure 0.5mOutlet pressure 1.369mOutlet pressure 2m

Figure 4.7: Pressure at outlet in Case 1 simulations

-2e+006

-1.5e+006

-1e+006

-500000

0

500000

0 0.0005 0.001 0.0015 0.002 0.0025 0.003 0.0035

[Pa]

TIME [s]

KP9 0.5mKP9 1.369mKP9 2m

Figure 4.8: Effect of pipe length on depressurisation

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5e+006

6e+006

7e+006

8e+006

9e+006

1e+007

1.1e+007

0 0.0005 0.001 0.0015 0.002 0.0025 0.003 0.0035

[Pa]

TIME [s]

Outlet pressure 0.5msOutlet pressure 1msOutlet pressure 1.5msOutlet pressure 2ms

Figure 4.9: Pressure at outlet in Case 2 simulations

-1.5e+006

-1e+006

-500000

0

500000

1e+006

1.5e+006

0 0.0005 0.001 0.0015 0.002 0.0025 0.003 0.0035

[Pa]

TIME [s]

KP9 0.5msKP9 1msKP9 1.5msKP9 2ms

Figure 4.10: Effect of break opening time on depressurisation

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5 Large Break test facilities

Large Break tests have been conducted in several laboratories and with numeroustest facilities. The tests carried out before the Three Mile Island accident (1979)were mainly designed to study phenomena and overall behaviour of the system duringLBLOCA. However, after the accident it was realized that also small breaks couldcause severe damage to the reactor core. The main focus in the studies moved to

smaller breaks and leaks and the majority of the test facilities was designed to studythese accidents.

The main goal of this work is to review LBLOCA tests and more precisely the testswhich were designed to study phenomena during the blowdown phase of the accident.During the rapid depressurisation in the blowdown phase, pressure waves propagatethrough the primary system. Forces induced by these waves could cause deformationof the core barrel. Therefore, the interest is in the tests which were designed to studydepressurisation, critical flow or fluid-structure interactions during LBLOCA.

The reviewed tests are divided in two categories, to separate effects tests and in-

tegral effects tests. Separate effects tests are designed to investigate individual orlocalized phenomena. Integral effects tests are quite opposite of the separate effectstests. They aim to simulate the overall behaviour of the reactor system and the in-teractions between the different components. The relatively large scale of the mostintegral test facilities makes the detailed measurement of the two-phase phenomenaquite impossible. Therefore the data acquired in the separate effects tests is needed inthe detailed modelling of the two-phase phenomena and code validation. [27]

Separate effects tests are reviewed in section 5.1 and integral effects tests in section5.2.

5.1 Separate Effects Tests

Some separate effects test facilities have been constructed to study LBLOCA blowdown.However, the vast majority of them have been designed for critical flow rate studies.In this report Edwards-OBrien pipe (5.1.1), Bartaks pipe (5.1.2), CANON and SuperCANON (5.1.3) and Marviken (5.1.4) depressurisation facilities are described. As it wassaid earlier deformation questions are important during LBLOCA blowdown. There-fore couple of test facilities with core internals have been built to study the effects ofblowdown on the internal structures of the reactor. CSE (5.1.5), SEMISCALE aka

Bettis Flask (5.1.6), Piper (5.1.7), Battelle-Frankfurt RS-16B (5.1.8) and Hei DampfReaktor (HDR) (5.1.9) are described here in more detail.

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Figure 5.1: Schematic of Edwards-OBrien blowdown pipe [7]

5.1.1 Edwards-OBrien pipe

Edwards and OBrien conducted a test in 1973 which is known as International Stan-dard Problem (ISP) 1. Their experimental test facility was designed to study criticalflow and contained a straight pipe whose length was 4.1 meters and internal diame-ter 73 millimeters (Figure 5.1). They conducted a series of depressurisation tests inwhich the pipe was initially filled with water. In the ISP test the pipe was pressur-ized to 7 MPa and the water temperature was 240C. The desired initial state wasobtained by first removing the air from the pipe and then filling it with demineralized

water. The leakage test was made by pressurizing the pipe without heating to overdepressurisation state. After that the pipe was heated to get the wanted temperaturedistribution. Depressurisation was initiated by using a glass rupture disc (thickness0.5 in. and diameter 3.5 in.) which was shot with a pellet gun. Opening time of thebreak is estimated as 1 ms using a thin conductive silver layer on the surface of therupture disk. Flow area was reduced by about 13%, because the glass disc did not fullybreak. Instrumentation system of the pipe consisted of 7 pressure and 7 temperaturemeasurements during the transient. Also local void fraction was measured in two lo-cations with x-ray densitometers. Axial blowdown forces were measured using a loadcell which was firmly attached to concrete. [7, 37, 38]

It was found in these tests that thermal non-equilibrium conditions are significantduring the fast depressurisation transients. During the years these tests have been themost popular benchmark experiments when testing code capabilities. [38]

5.1.2 Bartaks pipe

One blowdown experiment series was conducted by Bartak. Bartaks facility was ascaled model of the soviet WWER-type pressure vessel with one inlet and one outletnozzle (Figure 5.2). The length of the blowdown pipe was 1.7 meters and the diameter88 millimeters. The blowdown transient was initiated by using a rupture disc assembly.

The assembly consisted of two rupture discs. The pressure between these discs was 50%

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Figure 5.2: Schematic of Bartaks horizontal experimental channel [8]

of the system pressure. The outer rupture disc was broken by increasing the pressure

between the discs. After that the inner rupture disc broke as well. Total 13 tests werecarried out with different initial conditions. Maximum initial system pressure was 12MPa and temperature 300C. Findings from these tests were similar as in Edwards-OBrien experiments, thermodynamic non-equilibrium effects are significant. [8]

5.1.3 CANON and Super CANON

Series of blowdown experiments were performed with the CANON (Figure 5.3) andSuper CANON blowdown (Figure 5.4) facilities. Blowdown experiments from highinitial temperature (320C at maximum) and pressure (15 MPa) were carried out with

the latter facility. The length of the Super CANON blowdown pipe was 4.389 metersand the diameter 102.3 millimeters. The length of the CANON blowdown pipe was3 meters and the diameter 102.3 millimeters. However, the initial pressure was muchlower (3.2 MPa) in CANON and therefore the initial temperature was lower as well(230C at maximum). A rupture disc assembly (two discs) was used as a break devicein both facilities. Diameter of the break area ranged from 30 to 100 millimeters. In theSuper CANON experiments absolute pressure was measured in 7 locations along thepipe and temperature in 6 locations. Also void fraction was measured in one locationusing neutron scattering technique. In the CANON experiments absolute pressure andtemperature were measured from 4 locations and void fraction with neutron absorption

technique. Experimental results gathered in these experiment series have been used forcode assessment purposes (CATHARE). [10, 39, 40]

5.1.4 Marviken

The Marviken power plant or R4 reactor in Sweden was supposed to operate as aboiling heavy water direct cycle reactor, but instead it served as a critical flow testfacility. The plant was never used as a nuclear reactor so it was an ideal large scale testfacility for critical flow experiments. The test facility was operated in the years 1972-1985. Containment response, critical flow, jet impingement and level swell tests werecarried out in 1977-1981. The critical flow test project started in 1976 and the tests

were carried out in the years 1978-1979. The tests are well documented and reports are

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Figure 5.3: Schematic of CANON facility [9]

Figure 5.4: Schematic of Super CANON facility [10]

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available in the OECD Nuclear Energy Agency data bank. Also an info sheet aboutMarviken experiments can be found from the Code Validation Matrix. [11, 39]

5.1.4.1 Overview of Marviken experiments

As the Marviken power plant was originally designed as a full scale reactor it providedexcellent opportunities for plant scale testing (Figure 5.5). The net volume of thepressure vessel was 425 m3, design pressure 57.5 bar and temperature 272C. Innerheight of the vessel was 24.55 meters and diameter 5.22 meters. The vessel was modifiedby removing parts of the internals and installing a discharge pipe vertically to thebottom. Test nozzle was attached to the other end of the discharge pipe. Break devicewas a rupture disc assembly with two rupture discs. The volume between the discs waspressurized to half of the system pressure. In the reports it is referred as the balancingpressure. The volume between the discs was pressurized by using one tank and thebreak was initiated by using an other tank with higher pressure. [11]

In total 27 tests were performed in the critical flow test series. The test parameterswere varied so that the main objective of the series was reached. The objective was todetermine how critical mass flow rate depends on different parameters like upstreamstagnation pressure and enthalpy, nozzle length and diameter and air concentration.Steam dome pressure was slightly varied and therefore the saturation temperaturevaried as well. The mass of water (saturated and subcooled), nozzle size and subcoolingdegree were also varied. Oxygen content in the water was measured in part of the testsand in test 27 it was lowered significantly and results were compared with test 21.[11, 41]

5.1.4.2 Instrumentation

The pressure vessel and the discharge line had vast instrumentation. The pressure ves-sel had six absolute pressure transducers and several differential pressure transducers.Also temperatures were measured from the cooling lines for the water cooled pressuretransducers. Inside the vessel 28 temperature measurement points were used to obtainthe temperature profile. The discharge line had several differential and absolute pres-sure transducers. Temperature was measured from many points as well. For densitymeasurement a three-beam gamma densitometer was used. [41, 42]

5.1.5 CSE blowdown experiments

Containment Systems Experiment (CSE) reactor simulator was instrumented to mea-sure pressures, temperatures, water level, void fraction in the exit duct, mass of fluidin vessel, forces and stresses during loss of coolant accident. The reactor simulator wasdesigned to operate in temperatures up to 600F and pressures up to 2750 psig (approx-imately 316C and 191 bar). The volume of the vessel was 150 ft3 (appr. 4.2 m3). Theblowdown pipe was attached to the lower part of the vessel and the break was initiatedusing a double rupture disc assembly (Figure 5.6). The core was simulated with a quitesimilar approach as in the Piper facility by using a perforated platedummy core. The

object of the blowdown experiments was to gather data for code comparison. Specialinterest was in blowdown forces which are subjected to the simulated core. [12, 43]

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Figure 5.5: Schematic of Marviken critical flow test facility [11]

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Figure 5.6: Schematic of CSE reactor simulator [12]

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Figure 5.7: Schematic of the vessel used for SEMISCALE test 711 [13]

5.1.6 Semiscale (Bettis Flask)

The Water Reactor Research Test Facility (Idaho National Laboratories, Test AreaNorth) includes the Semiscale Facility, the Blowdown Facility, and the Full Area SteadyState Testing (FAST) Facility. The Blowdown Facility was constructed to study thebehaviour of the instruments and components during the depressurisation phase ofthe LBLOCA. The facility was used from 1965 to 1968 and three test series wereconducted, named as 500, 600 and 700 series. The whole vessel was initially filledwith subcooled water and the depressurisation was initiated by using a rupture diskassembly. Schematic of the vessel used in Semiscale test series 700 and test 711 showsthe places of the instrumentation and the scale of the facility. (Figure 5.7). [13, 18, 44]

5.1.7 Piper

Piper facility was designed to study different aspects of blowdown (Figure 5.8). Eval-uation of loads on internal structures was one studied subject. The facility was mainlyused to simulate BWR blowdown as the initial content of the vessel was steam-watermixture. However, initial void varied from 0 to 0.9 in the performed tests. The pres-sure range was from 1 to 9 MPa and the subcooling degree from 0 to 150 C. Blowdownwas initiated using the double rupture disc assembly, probably with similar workingprinciple as in Bartaks experiments. Break opening time varied between 3 and 5 ms.Internal structures of the vessel were simulated by a plate with a central hole. [14]

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Figure 5.8: Schematic of Piper vessel and blowdown nozzle [14]

5.1.8 Battelle-Frankfurt RS-16B

A series of blowdown experiments (indicated as DWR tests) was carried out in Battelle-Institute in Frankfurt. These RS-16B experiments were designed to study blowdownloads on pressurised water reactor internals. The tests were carried out in 1975. Thetest facility simulated a reactor pressure vessel with internals (Figure 5.9). The heightof the vessel was 11.3 meters and the diameter 0.8 meters. The outside diameter of thecore barrel was 0.62 meters and the length 7.37 meters and it was attached at the topto a grid plate. This plate was rigidly attached to the vessel. Electrical heaters andsome passive rods simulated the core at the lower part of the core barrel. The mostsignificant difference between this facility and later introduced HDR facility was thediameter scale. The heights were almost identical, but the diameters in this facilitywere much smaller than in the HDR facility. Also the way to simulate the reactor corediffered as well. Because of the scale, thickness of the core barrel was less than in theHDR facility, the upper one third of the core barrel had a thickness of 8 millimeters andthe lower two thirds 18 millimeters. As an example it can be said that in test DWR5the initial pressure was 14.2 MPa and temperature 301C at maximum. The dischargepipe had a length of 0.35 meters and a diameter of 0.142 meters. Break was initiatedusing a rupture disc and the break opening time was approximately 1 ms. The largelength-to-diameter ratio of the vessel resulted in highly one-dimensional behaviour inthe lower parts of the vessel. [15, 45]

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Figure 5.9: Schematic of RS-16B blowdown facility and models[15]

5.1.9 Hei Dampf Reaktor (HDR)Hei Dampf Reaktor (HDR) was originally built as a research reactor and was locatedin Growelzheim near Karlstein am Main, in West Germany. However the reactor hadto be shut down in 1971 only after a short period of operation. The reason for thiswas constructional faults in the fuel elements. After the shutdown the reactor got anew research purpose as a test facility for large scale experiments when Project HDRwas established in 1973. The main goal of these experiments was to determine theimportance of the multidimensional fluid-structure interaction phenomena during theblowdown phase of a loss of coolant accident. Fluid-structural tests were performedearlier in a couple of smaller scale test facilities. For the code validation purposes larger

scale tests were needed. In the end of the 1970s and in the beginning of the 1980sseries of large scale blowdown experiments were performed in HDR facilities.

5.1.9.1 Overview of HDR experiments

HDR had a near full scale reactor pressure vessel (Figure 5.10). For the blowdowntests the reactor internals had to be modified. The original core barrel was too thickfor deformations to happen. It was replaced with a new one. This new core barrelwas carefully designed and fluid-structure interaction codes were used in determiningthe wall thickness of the barrel (23 mm). To get clear boundary conditions for thesimulations the core barrel was rigidly clamped at the top in most of the experiments

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Figure 5.10: Schematic of HDR pressure vessel and core barrel [16]

and the core itself was simulated with a mass ring attached to the bottom of the

core barrel. The weight of the mass ring was approximately half of the core mass.In pressurised water reactors the horizontal movement of the lower core barrel end islimited by the core barrel bumpers. In HDR facility there were also such bumpers,but clearances were so large that impacts were not expected to happen. The bumperswere there as an additional safety feature to prevent overloading of the barrel clamping.[1, 17, 46]

Issues that had to be solved during planning of HDR experiments included corebarrel clamping (rigid or non-rigid?), designing the new core barrel, air removal fromthe system, the effects of initial temperature distribution, simplification of the core andplanning the tests according to the pre-calculations to avoid plastic deformations.

HDR-RPV-I blowdown test program was divided in two phases, to preliminarytest phase and main test phase. Blowdown experiments V29.2, V31 and V31.1 wereperformed during the preliminary test phase. For obvious reasons testing was startedwith initial conditions which would result in low blowdown loads. Blowdown loads wereincreased in test V31. Test V31.1 was carried out to test repeatability. The preliminarytest phase was also needed to test instrumentation and measurement techniques andto get experience of the facility and its behaviour. Test matrix for the main test phasewas finalized during preliminary tests. [17]

During the main test phase four blowdown tests were performed: V31.2, V32, V33and V34. Instrumentation was increased significantly from the preliminary test phase.

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Table 5.1: HDR preliminary and main phase test matrices

Test Pressure Subcooling in downcomer Upper Core DowncomerNo. (bar) (C) (C) (C)

V29.2 90 30 293 273V31 110 50 308 268V31.1 110 50 308 268V31.2 110 50 308 268V32 110 78 308 240V33 110 78 308 240V34 110 78 240 240

Test V31.2 was a replication of test V31 to examine how well experiments can be

reproduced after a long time. Subcooling degree was increased in test V32. This testis also German Standard Problem number 5. In test V33 the effect of reduced breakarea (25% flow area) was studied and the initial conditions were identical with V32.Test V34 was done with non-rigid clamping of the core barrel. Non-rigid clampingallowed axial movement of the core barrel and impacts between the lower end of thecore barrel and the pressure vessel. Initial conditions of all tests are summarized inTable 5.1. Originally, more isothermal tests were supposed to be done in the mainphase, but the test matrix was changed so that test V34 was the only one. The lengthof the blowdown nozzle was 4.524 meters in V29.2 and 1.369 meters in all others. [47]

More in depth analysis of the HDR test results can be found from Wolfs arti-cles [17, 47]. From the preliminary tests Wolf concludes that the repeatability of theexperiments and the reliability of the measurements were excellent despite some diffi-culties in density and displacement measurements. The results of tests indicated thatFSI reduces blowdown loads to the core barrel, three-dimensional effects exist duringblowdown, higher subcooling and shorter blowdown pipe increase the core barrel loads,maximum loads occur during subcooled blowdown, the core barrel displacements arequite small and that only substantial displacements of the pressure vessel have beenmeasured. From the main test phase Wolf summarizes that higher subcooling increasesthe blowdown loads, the reduction of the break area to 25% reduces loads by 10 to 40%and the non-rigidly clamped core barrel have significant effect on core barrel dynamics,

but hardly any on fluid dynamics. Many pre- and post-analyses have been done withseveral fluid-structure interaction codes. Comparisons between the measurement dataand calculations can be found from the several articles. [17, 47, 48]

5.1.9.2 Instrumentation

The instrumentation of the HDR facility was quite extensive. Instrumentation in thedischarge pipe during the preliminary test phase consisted of several fast pressure trans-ducers, a two-beam gamma densitometer, two dragbodies and many pairs of pressureand temperature sensors (Figure 5.11). In order to repeat experiments initial temper-

ature distribution was measured. Measurement was done first with 122 temperaturesensors and during the blowdown 25 temperature sensors were used. By doing so it

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Figure 5.11: Instrumentation of the long and short discharge pipes in HDR [17]

was possible to reliably interpolate the temperature distribution inside the HDR vessel.The vessel itself had very vast instrumentation. In the core barrel there were absolutepressure transducers, differential pressure transducers and temperature sensors to mea-sure fluid dynamic quantities. To obtain structural quantities strains, accelerations anddisplacements were measured. The most significant problem in the use of HDR data

for CFD validation is the lack of pressure data inside the core barrel. [ 17]

5.1.10 Other tests

Some other critical flow and break flow studies that have been referred in literature andtechnical reports are LOFT-Wyle Transient Fluid Calibration test facility (WSB03Retc.), Carofano-McManus, Cumulus, Deich, Fincke-Collins and Neussen. Some othercritical flow tests are shortly described in the research report made by Yildiz et al.[49,50]

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5.2 Integral Effects Tests

Integral test facilities together with the well planned separate effects tests have been thebase for the thermal-hydraulic code validation. Integral test facilities have been built

to study different aspects of LOCA. Some facilities were used only to simulate smalland intermediate breaks while couple of facilities were designed to study all LOCAbreach windows. LOFT, SEMISCALE (5.2.1) and LOBI (5.2.2) facilities are brieflydescribed next.

5.2.1 LOFT and SEMISCALE

Loss Of Fluid Test facility (LOFT) was constructed in the Test Area North at IdahoNational Engineering Laboratory. The building was done between 1965 and 1975.LOFT facility was volumetrically scaled (coolant volume to power ratio) and it had

a real nuclear core which made it very unique in the field of the thermal-hydraulicresearch (Figure 5.12). The facility had two loops, one intact and one broken loop.The intact loop was scaled and designed to simulate three unbroken loops of a genericfour-loop PWR during LOCA. The broken loop was not closed, hot leg and cold leghad quick opening valves which initiated the LOCA. The coolant was discharged tothe pressure suppression system to prevent contamination and to simulate containmentpressure. In total 38 experiments ranging from large to small breaks were carriedout in the LOFT facility. Instrumentation was quite extensive, because the facilityhad normal PWR instrumentation and in addition vast temperature, pressure andmechanical measurements. The number of measurements and some main dimensions

are presented in Figure 5.13. [18, 51]

Figure 5.12: LOFT test facility [18]

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Figure 5.13: LOFT reactor vessel instrumentation [18]

First model of SEMISCALE integral test facility became operational in September1974. It was scaled to LOFT facility and its original purpose was to study LBLOCA. Itwas a full height and volume scaled facility. The test facility had an electrically heatedcore so it was easier to operate than the LOFT facility. Westinghouse 4-loop PWRwas used as a reference plant in the further modifications of the facility. Those latermodifications were also a full height and volume scaled (1/1705.5). [18]

Figure 5.14: Semiscale Mod-1 (left) and Semiscale Mod-2C (right) [18]

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5.2.2 LOBI

LOop Blowdown Investigation facility (LOBI) was originally designed to study largebreak loss of coolant accidents. The project began in the early 1970s when there

was a special need to research the ECCS performance in large power reactors. Thetest facility was build in Joint Research Center in Ispra and the German BIBLISB power reactor was used as a reference plant. The LOBI-Mod1 test facility wascommissioned in December 1979 and the experimental program was concluded in June1982 (Figure 5.15). During that period of time 28 experiments were carried out ofwhich 25 were LBLOCA tests. The facility was scaled using the power to volume scalingcriterion (volume ratio 1/700) except that the pressurizer had to be scaled differently toaccommodate the internal heaters. The downcomer scaling in the original test facilityis said to be the only major exception. It was made intentionally larger to avoid ECCbypass, but later on this was found out to be a mistake. The downcomer was replacedwith a new one which had a 12 millimeter gap instead of 50 millimeters. During thefirst phase of the LOBI project some experiments were also made to evaluate testrepeatability and the effects of the break size and geometry on the course of LBLOCA.[18, 52]

After the first phase of the LOBI project the facility was modified for the smallbreak studies. The experimental work continued in April 1984 and the last experimentwith LOBI-Mod2 was done in June 1991. [52]

Figure 5.15: LOBI test facility [18]

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6 Test facility summary

LBLOCA blowdown phase has been studied from many angles. There have been simplepipe facilities which have been used to study depressurisation and critical flow. Coupleof the test facilities have been more reactor like and have been used to study alsostructural forces and loads. Third category is the integral test facilities which simulatethe whole primary circuit and have been used for thermal-hydraulic studies. All the

test facilities described before are summarized in table 6.1 with the estimated scale(S=Small, M=Medium, L=Large) and studied phenomena.

The task of this study was to find out whether there are better test facilities forfluid-structure code validation than HDR or not. After the very wide search for thefacilities it can be concluded that none of the found facilities can compete with theresults gained from the HDR experiments. The next question is: could we somehowimprove measurements from the HDR experiments to gain some additional knowledgeof the fluid-structure interaction phenomena with a new test facility. General issuesand problems related to designing and scaling a new test facility are introduced inchapter 7.

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Table 6.1: Test facilities summarized

Test facility Scale Studied phenomena

Separate Effects TestsEdwards-OBrien pipe S critical flow, depressurisationBartaks pipe S critical flow, depressurisationCanon S critical flow, depressurisationSuper Canon S critical flow, depressurisationMarviken L critical flow, depressurisationCSE M critical flow, depressurisation, loads, forcesSemiscale (Bettis Flask) S critical flow, depressurisation

Piper S critical flow, depressurisation, loads, forcesBattelle-Frankfurt RS-16B L critical flow, depressurisation, loads, forcesHDR L critical flow, depressurisation, loads, forces

Integral Effects TestsLOFT L thermal-hydraulic effects during LBLOCASemiscale L thermal-hydraulic effects during LBLOCALOBI L thermal-hydraulic effects during LBLOCA

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7 Designing the new facility

The facilities like HDR and Battelle-Frankfurt facility were nearly full scale and builtto study fluid-structure interactions during the blowdown phase of LBLOCA. Buildingof such a large scale facility is expensive. The size of the existing laboratories mayalso set some limitations. Therefore one solution is to build a small scale facility witha lot of instrumentation. The initiative to design a new test facility came from the

code developers. They are searching for suitable test data to validate fluid-structureinteraction calculations done with a coupling of CFD and structural analysis codes.Some calculations have been done with the HDR geometry, but there are some flawsin the facility that restrict the use of experimental data for code validation purposes.For obvious reasons (tests were done in the 70s and 80s) the facility was not designedto validate CFD codes. For that purpose there must be a dense grid of pressure andtemperature measurements. In this case the pressure measurements would be enough.When the task is to validate the coupling of CFD and structural analysis codes extensivestructural measurements are also needed. Certain important aspects have to be takeninto account when designing a new separate effects test facility; scaling, structural

questions, instrumentation and planned test program. These questions are considerednext each in their own sections.

7.1 Scaling

In order to get results that could be scaled to the full scale of the reference plant, thetest facility should be designed using appropriate scaling laws. It is essential to choosethe right scaling law to capture the nature of the phenomena correctly. If the purposeof the experiments is to study a highly multidimensional phenomenon, dimensions ofthe test facility must be well chosen. During the blowdown phase the phenomena are

three-dimensional. For example in Battelle-Frankfurt blowdown experiments it wasfound out that in the lower parts of the pressure vessel the behaviour of the systemwas mainly one-dimensional due to the small width to height ratio. The problem withsmall scale or volumetrically scaled facilities is that phenomena are quite distorted.One example of this is the difference in ECC bypass phenomenon between UPTF andSEMISCALE experiments.

There are two traditional ways to scale a test facility: time-preserving volumetricscaling and time-reducing length scaling. Volumetric scaling is not an option for thisfacility, because the essential phenomena would become distorted. Nahavandi et al.

introduced three scaling laws: time-reducing length scaling, time-preserving volumetricscaling and time-preserving idealized model/prototype scaling. The most applicable of

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these laws to this case would be the time-reducing length scaling. The problem withthis scaling law is that accelerations are scaled with the inverse of length scale thatmeans 1/. However, they concluded that

For the transient problems such as the blowdown phase of the loss-of-coolant accident (LOCA), steam break, and water hammer phenomena, bodyforces due to the acceleration of gravity play an insignificant role comparedto the pressure differential, and distortion in the model is insignificant.However, for transient problems involving density wave phenomenon, suchas steam generator hydrothermal instability and the reflood phase of theLOCA, the elevational head in the downcomer due to gravity is consider-able compared to the pressure differential. For these problems, the modelpredictions using the above similarity rationale will be significantly distorted(i.e., different from the prototype behavior).

This means that the scaling law in question would be applicable to this particular case.The development of the time-reducing scaling law begins from the basic equations:continuity, momentum and energy equations. As a result it is obtained that model andprototype are similar if

= 1 (7.1)

where is the time-scale and is the length scale. This means that the time-scaleshould equal the length-scale (time-reducing). The time-reducing scaling law is pre-sented in table 7.1.[19]

As the wall temperature, power input, number of rods and rod diameter are notsubjects of interest, the law adapts quite nicely to blowdown phase study. Area andvolume are simply developed from the length scale. For example hydraulic diameterfor round tube equality can be derived as

Dh =4A

U=

42D2

4

D

=2D2

D=

(D)2

D

= D = D

= D DD

=

(7.2)

Dimensions of the prototype are marked with the star (*). A test facility scaled withthis law is a model of the prototype. The scale of the test facility is another issue andis discussed in the next section. [19]

7.2 Structures

Structural design of the test facility is a very challenging task. Pre-calculations should

be made to evaluate forces on the core barrel. The scale of the facility should be opti-mized using the results of these calculations and restrictions set by the instrumentation.

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Table 7.1: Time-reducing scaling law [19]

Model/Prototype Time ReducingLength

Time Hydraulic diameter 1Area 2

Volume 3

Velocity 1Acceleration 1/Properties (pressure, temperature, etc.) 1Heat generation/unit volume/unit time 1/Power 2

Rod diameter 1

Number of rods 2

Heat flux 1/Wall temperature DistortedLimitation Turbulent flow

Negligible buoyancy

The task is not the detailed designing of the test facility, but to give an overall viewwhat the test facility could be like and what structural decisions could be applied tothis case. One subject of the future studies could be the assessment of break opening

time effects. The most challenging task in this case would be the designing of such abreak assembly by which different break opening times could be produced and repro-duced reliably. Code developers have expressed that the shape of the dynamic loadshould be correct. Due to the time-reducing scaling law the break opening time shouldbe scaled with the dimension ratio (the length scale). This means smaller openingtime. However, the physical restrictions of actual break components set the value forthe smallest achievable break opening time (possibly 1 ms). This means that the con-servative scenario of the LBLOCA may not be studied as the break opening time wouldhave to be much smaller than 1 ms in a small scale test facility. In the best-e

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