effect of crgt cooling on modes of global vessel failure

EFFECT OF CRGT COOLING ON MODES OF GLOBAL VESSEL FAILURE OF A BWR LOWER HEAD

Walter Villanueva Division of Nuclear Power Safety

Royal Institute of Technology (KTH) Stockholm, Sweden

Chi-Thanh Tran Institute of Energy 6 Ton That Tung

Dong Da, Hanoi, Vietnam

Pavel Kudinov Division of Nuclear Power Safety

Royal Institute of Technology (KTH) Stockholm, Sweden

ABSTRACT An in-vessel stage of a severe core melt accident in a

Nordic type Boiling Water Reactor (BWR) is considered

wherein a decay-heated pool of corium melt inflicts thermal and

mechanical loads on the lower-head vessel wall. This process

induces creep leading to a mechanical failure of the reactor

vessel wall. The focus of this study is to investigate the effect of

Control Rod Guide Tube (CRGT) and top cooling on the modes

of global vessel failure of the lower head.

A coupled thermo-mechanical creep analysis of the lower

head is performed and cases with and without CRGT and top

cooling are compared. The debris bed heat-up, re-melting, melt

pool formation, and heat transfer are calculated using the Phase-

change Effective Convectivity Model and transient heat transfer

characteristics are provided for thermo-mechanical strength

calculations. The creep analysis is performed with the modified

time hardening creep model and both thermal and integral

mechanical loads on the reactor vessel wall are taken into

account. Known material properties of the reactor vessel as a

function of temperature, including the creep curves, are used as

an input data for the creep analysis.

It is found that a global vessel failure is imminent

regardless of activation of CRGT and top cooling. However, if

CRGT and top cooling is activated, the mode and timing of

failure is different compared to the case with no CRGT and top

cooling. More specifically, with CRGT and top cooling, there

are two modes of global vessel failure depending on the size of

the melt pool: (a) ‘ballooning’ of the vessel bottom for smaller

pools, and (b) ‘localized creep’ concentrated within the vicinity

of the top surface of the melt pool for larger pools. Without

CRGT and top cooling, only a ballooning mode of global vessel

failure is observed. Furthermore, a considerable delay (about

1.4 h) on the global vessel failure is observed for the roughly

30-ton debris case if CRGT and top cooling is implemented.

For a much larger pool (roughly 200-ton debris), no significant

delay on the global vessel failure is observed when CRGT and

top cooling is implemented, however, the liquid melt fraction

and melt superheat are considerably higher in non-cooling case.

INTRODUCTION The success of ex-vessel severe accident termination

and melt coolability in a deep pool of water located under the

reactor vessel of a Nordic Boiling Water Reactor (BWR) is

highly dependent on the melt ejection characteristics from the

lower head. It has been shown that melt fragmentation and

resulting characteristics of the debris bed ([1], [2], [3], [4]) as

well as coolability [5] are controlled by the mode of vessel

failure (rupture size and location, characteristic time of melt

ejection from the vessel, melt composition, amount of liquid

melt and its superheat, etc.). A non-coolable debris will dry out,

re-heat, and then re-melt, and consequently attack containment

base-mat, presenting a threat to containment integrity. In

addition, energetic molten fuel-coolant interaction, e.g., steam

explosion, in a deep reactor cavity is also sensitive to melt

discharge conditions [6] and is also known to create another

Proceedings of the 2012 20th International Conference on Nuclear Engineering collocated with the

ASME 2012 Power Conference ICONE20-POWER2012

July 30 - August 3, 2012, Anaheim, California, USA

ICONE20-POWER2012-54955

1 Copyright © 2012 by ASME

source of credible threat to containment integrity. Therefore, in

order to establish firmly that containment integrity can be

preserved, it is necessary to reduce uncertainties in the modes of

vessel failure.

Previously, a 2D coupled thermo-mechanical creep analysis has

been performed by Willschuetz et al. ([7], [8]) for the

FOREVER experiment (a scaled model of a pressurized water

reactor (PWR) vessel lower head) in order to investigate global

failure mode and predict failure time. A similar study by

Villanueva et al. [9] has focused on the analysis of the global

failure of a Nordic type BWR vessel wall under thermo-

mechanical load inflicted by a melt pool assuming that CRGT

and top cooling of the melt are available. It was found that the

amount of melt has an effect on the global vessel failure mode,

i.e., ballooning or localized creep. Another group of vessel

failure modes include failure of the local penetrations such as

nozzles of Instrumentation Guide Tube (IGT), CRGT, and the

pump. Previous work by Rempe et al. ([10], [11] addressed

different failure mechanisms in reactor pressure vessel (RPV)

lower heads for different PWR and BWR designs. They

performed structural analyses for the cases with metallic and

ceramic debris with the assumption that debris is unable to go

through the penetration. They found that vessel failure will not

occur in the case of metallic debris within 24 hours but vessel

failure is predicted in the case of ceramic debris by localized

thinning in the bottom of the vessel in less than 4 hours after

debris relocation. In the metallic case, heat from the 24 tons of

debris is primarily transferred to the coolant thus inflicting less

thermal load onto the vessel wall. While in the ceramic case

(120 tons), debris temperatures continue to rise because of

decay heat. A study by Ikonen [12] also addressed different

failure mechanisms in the lower head by a thermo-mechanical

analysis and supported the hypothesis that the pressure vessel

fails first at the welds of the IGT nozzles, but it is also pointed

out that there is an uncertainty in the interaction of the crust and

the nozzle that can cause a delay in failure. However, no CRGT

cooling was taken into account in all the above-mentioned

analyses and its potential as a severe accident mitigation

measure has not been investigated.

The goal of the present work is to investigate if CRGT cooling

(which also provides water for melt cooling from the top) can

have a significant effect on the modes of global vessel failure of

the lower head. The analysis is performed on a Nordic design

BWR lower head considering the cases with and without CRGT

cooling and debris top cooling at different melt pool depths.

The debris bed heat-up, re-melting, melt pool formation and

heat transfer are predicted with the Phase-change Effective

Convectivity Model (PECM) [13]. The PECM is implemented

on the platform of the Fluent code [14], and transient heat

transfer characteristics are provided for thermo-mechanical

strength calculations. The creep analysis is performed using the

commercial code ANSYS [15], taking into account both the

thermal and integral mechanical loads on the RPV. Material

properties of the reactor vessel are taken from Rempe et al.

[10]. The deformation of the vessel wall is assumed to have

negligible effect on the melt pool heat transfer, so only one way

coupling between PECM and ANSYS is employed.

In the next section, the modeling and simulation of the debris

bed and melt pool heat transfer is presented. Thermal transient

loads from these calculations are then used as boundary

conditions for the coupled thermo-mechanical creep analysis of

the vessel wall which is presented in Section II. This is followed

by concluding remarks in Section III.

I. Debris Bed and Melt Pool Heat Transfer

I.1. Heat Transfer Simulation Models

In the late phase of in-vessel accident progression, a debris bed

can form in the BWR lower plenum. If cooling is not adequate,

the debris bed heats-up and then re-melts. The debris bed heat-

up can be simulated by a conduction model. The re-melting

stage is simulated by the Effective Convectivity Model (ECM)

and the PECM [13].

Mechanical strength of the vessel wall depends on the thermal

loads imposed on its boundaries. In the present work the ECM

is used for prediction of transient thermal loads imposed on the

vessel wall from a decay-heated debris bed and melt pool

formed in the BWR lower head during a severe accident. The

ECM has been developed based on the concept of effective

convectivity proposed by Bui and Dinh [16]. The ECM is a

model for description of turbulent natural convection heat

transfer in an internally heated fluid volume. In the ECM

method, the convective terms of the energy conservation

equation are described using characteristic velocities; therefore

the need of solving Navier-Stokes equations is eliminated [13].

The characteristic velocities are determined using experimental

heat transfer correlations, namely the upward, sideward and

downward Steinberner-Reineke correlations [17]. The ECM is

implemented in the commercial code Fluent, to utilize all

advantages of a CFD commercial code solver such as the pre-

and post-processing.

To describe the crust formation phenomena, the ECM was

extended to the PECM. The ECM and PECM have been

validated against a set of experiments which cover a broad

spectrum of physical phenomena involved in melt pool

formation heat transfer, and wide range of Rayleigh number.

The ECM and PECM have been demonstrated as sufficiently

accurate and computationally efficient tools for 3D simulations

of melt pool formation heat transfer for a BWR accident

analysis [18].


I.2. Computational aspects

I.2.1. Material properties

A homogeneous melt pool is considered in this study. The melt

is assumed to be an oxidic corium with material properties

adopted from Theofanous et al. [19]. Scenarios of core damage,

relocation and quenching in the lower head water pool, which

can lead to formation of heterogeneous (oxidic and metallic

debris) or formation of stratified melt configuration (e.g. with

layers of metallic and oxidic melts) require further investigation

and are beyond the scope of this work.

I.2.2. Geometry and mesh

PECM simulations are performed for a 3D slice of BWR lower

plenum filled with decay-heated corium (see Figure 5 and

Figure 7 as references). The slice geometry is a segment of

BWR lower plenum that includes 8 CRGTs and bounded from

below by the vessel wall. It is assumed that the IGTs are melted

and plugged by corium melt during gradual re-melting of the

debris, and do not have an influence on melt pool heat transfer.

The CRGTs may be used as a flow path for water injection to

the reactor. In such a case, the CRGTs are protected from

thermal attack by cooling from inside. The water is assumed to

be ejected from the CRGTs providing a water layer atop of the

debris bed.

Two cases are considered in the present paper for both

scenarios with and without CRGT cooling; in the first case, the

maximum debris bed height is 0.7 m, and in the second case the

maximum bed height is 1.9 m. The slice is meshed with

hexahedron and tetrahedron 3D elements. To better capture the

dynamics of phase-change boundaries, Adaptive Mesh

Refinement (AMR) technique in Fluent is employed. The debris

bed computational domain is limited to about 1.5 million cells.

I.2.3. Boundary conditions

With CRGT cooling, isothermal boundary conditions (water

saturation temperature) are applied on the debris bed top and

inner CRGTs walls. For the other surfaces, Neumann boundary

conditions are applied. The external surface of the vessel wall is

covered with insulation so a small heat flux (20 W/m2) is

allowed.

For the cases without CRGT cooling, radiation heat transfer is

applied on the top debris bed surface with T∞ equal to

saturation temperature and the emissivity coefficient is set to

0.8. Zero heat-flux boundary conditions are applied on the

CRGT inner surfaces.

Initial temperature is set to 450 K and the solidus and liquidus

temperatures of debris are assumed to be 2750 K and 2770 K,

respectively. The solidus and liquidus temperatures of steel

(material of the vessel wall and CRGTs) are 1671 K and 1727 K

respectively [10].

I.3. PECM Calculations: With and Without

CRGT and Top Cooling

In this section, PECM calculations of 2 different debris bed

heights, H = 0.7 m and 1.9 m are presented. The debris bed

mass is considerably smaller in the 0.7 m case, around 30 tons,

compared to the 1.9 m case which is ~200 tons (considered the

maximum amount of melt that can get relocated to the lower

head). In both cases, the effect of CRGT and top cooling is

investigated. It is noted that some results for the case with

CRGT and top cooling have been presented in [20].

Figure 1. Temperature snapshot from PECM calculation for H =

0.7 m debris bed at t = 3.06 h, (a) with CRGT and Top cooling,

(b) without CRGT and Top cooling.

Snapshots of temperature profile is shown in Figure 1 for

H = 0.7 m at time t = 3.06 h comparing the cases with and

without CRGT and top cooling. For ease of discussion from

this point onwards, when we say ‘with cooling’ or just ‘cooling’

in a context, we meant ‘with CRGT and top cooling’. The

(a) with CRGT and Top cooling

(b) without CRGT and Top cooling


maximum temperature is a bit lower for the case with cooling,

which is 2708 K, while it is 2761 K for the case without

cooling.

Figure 2. Melt fraction snapshot from PECM calculation for

H = 0.7 m debris bed at t = 3.06 h, (a) with CRGT and Top

cooling, (b) without CRGT and Top cooling.

Snapshots of the melt fraction profile is shown in Figure 2 for

H = 0.7 m at time t = 3.06 h. The debris bed is still mostly solid

with the cooling case, which is only 5 % melt fraction located in

the bulk, compared to 18 % melt fraction with the non-cooling

case. In other words, the debris bed re-melts faster in the non-

cooling case as expected, which can also be seen in Figure 2b in

the regions between CRGTs.

Details of the melting of the CRGTs are shown in Figure 3 that

corresponds to H = 0.7 m. At about 1.94 h in Figure 3a, the

CRGTs start to melt with a direction from the center region of

the CRGT to its top and bottom, and from the center CRGT to

the periphery of the debris bed. At 2.22 h shown in Figure 3b,

the melting progresses in the same way and 5 out of the 7

CRGTs submerged in the debris bed has partially melted. Then

later at 2.78 h (Figure 3c), the melting follows the same trend

and 6 out of 7 of the CRGTs have larger melted factions than

previously.

The surface-averaged temperature profile of the bottom area of

the debris bed in contact with the vessel wall inner surface is

shown in Figure 4. The cases with and without cooling are

compared for H = 0.7 m at different times. At t = 3.06 h (Figure

4a) for the case with cooling, the average temperature is about

900 K and then decreases to about 468 K as it reaches the top.

Without cooling, the temperature is about 1200 K (with smaller

difference between minimum and maximum values compared to

the cooling case) and then decreases to about 706 K as it

reaches the top. At a later time t = 3.61 h, the temperature

profiles follow the same trend. With cooling, the temperature is

about 1000 K and decreases to about 475 K. Without cooling,

the temperature has smaller variations along the vessel wall

with a peak around 1400 K and decreases to about 757 K.

Figure 3. Melting of the CRGTs for H = 0.7 m debris bed at

times t = 1.94 h, 2.22 h, and 2.78 h (no CRGT and top cooling

provided).

(a) t = 1.94 h

(b) t = 2.22 h

(c) t = 2.78 h

(a) with CRGT and Top cooling

(b) without CRGT and Top cooling


Figure 4. Distribution of averaged temperature on the internal

vessel wall: Comparison of cases with and without CRGT and

Top cooling for H = 0.7 m debris bed at times (a) 3.06 h, and

(b) 3.61 h.

Snapshots of temperature and melt fraction for a larger debris

bed H = 1.9 m at t = 3.06 h are shown in Figure 5. With

cooling, the maximum temperature is 2762 K which is almost

equal to the non-cooling case at 2761 K. However, the melt

fraction in the cooling case is only 10 % compared to 29 % in

the non-cooling case. Parts of the CRGTs submerged in the

debris bed have mostly melted (more than 90 %) at this time. In

addition, the temperature and melt fraction in the periphery of

the debris bed is not affected whether there is cooling or not

since that region is not covered by CRGTs.

The surface-averaged temperature of the bottom area in contact

with the inner surface of the vessel wall is shown in Figure 6 for

H = 1.9 m. With cooling at t = 3.06 h, the average temperature

is about 900 K similar to the 0.7 m case but increases with a

peak around 1140 K in the region not covered by CRGTs, and

with a minimum temperature of 497 K at the surrounding tip of

the debris bed. Without cooling, the average temperature is

about 1200 K, even higher than the maximum temperature in

the region not covered by CRGTs, about 1147 K. At a later time

t = 3.61 h, the temperature follows the same trend. With

cooling, the temperature is about 1000 K and then increases

with a peak of 1296 K before it decreases to 513 K at the

surrounding tip. Without cooling, the temperature is about

1370 K then flattens to 1300 K in the region not penetrated by

CRGTs and finally decreases to 776 K at the tip.

In comparison, melting of the CRGTs starts at about 1.8 h in

both the 0.7 m and 1.9 m cases without cooling. But it is

considerably faster in the 1.9 m case in which about 90 % of the

CRGTs have melted after 2.6 h while it takes 1.4 h more in the

0.7 m case to reach 90 %.

II. Vessel Wall Thermo-mechanical Creep

Analysis

In this section, coupled thermo-mechanical creep analysis of the

pressure vessel wall is presented. The analysis takes into

account the debris bed heat transfer calculated by PECM that is

presented and discussed in the previous section.

II.1. Computational aspects

II.1.1. Material properties

The reactor pressure vessel that is considered is made of steel

SA533B1. The material properties that are all functions of

temperature such as density, elastic modulus (linear isotropic),

thermal conductivity (isotropic), specific heat, and coefficient

of thermal expansion are taken from Rempe et al. [10]. The

Poisson’s ratio is 0.3.

II.1.2. Axisymmetric model geometry and mesh

An axisymmetric model of a Nordic design BWR is considered

for the coupled thermo-mechanical creep analysis, see Figure 7.

Full transient analysis is implemented in ANSYS 13.0 [15]

where a strong structural-thermal coupling is supported. The 2D

geometry is meshed with 800 quadrilateral elements and 2731

nodes with an average edge length of 0.04. The element type

used is Quad Plane223 which is a 2D 8-node-coupled-field

(structural-thermal) solid.

(a) t = 3.06 h

(b) t = 3.61 h


Figure 5. Snapshot of PECM calculation for H = 1.9 m debris bed at t = 3.06 h. Temperature and Melt Fraction profiles, respectively,

(a)-(b) with CRGT and Top cooling, (c)-(d) without CRGT and Top cooling.

a) b)

Figure 6. Distribution of averaged temperature on the internal vessel wall: Comparison of cases with and without CRGT and Top

cooling for H = 1.9 m debris bed at times (a) 3.06 h, and (b) 3.61 h.

(a) (b)

(c) (d)


II.1.3. Boundary conditions

The thermal load from the debris bed is calculated using PECM

which is discussed in Section I. Other boundary conditions and

constraints are shown in Figure 7 which includes the weight

load, gravity, internal pressure (3 bars), and external pressure (1

bar). Debris bed of heights 0.7 m (that corresponds to a 30-ton

debris) and 1.9 m (that corresponds roughly to a maximum 200-

ton debris) are considered.

a)

b)

Figure 7. (a) 2D axisymmetric geometry and mesh with 850

quadrilateral elements and 2901 nodes, and (b) schematics of

mechanical load on the reactor vessel.

II.1.3. Creep model

For our analysis, a modified time hardening (primary) creep

model is chosen in ANSYS. Coefficients for the model are

generated using the experimental creep data for SA533B1 from

Rempe et al.[10]. Please see [9] for complete details along with

the validation test.

A yield or creep limit is not used in this study (see [9] for more

discussion). Instead we identify a range of percent strain that we

consider as reliably predicted by the model which is strains

within 20 %. This is partly based on the experimental creep data

and the creep model that we used. The time scale to reach 20 %

strain is one of our main interests since at this time the structure

is close to its mechanical failure.

II.2 Modes and Timing of Vessel Failure

Magnified views of von Mises creep strains for H = 0.7 m are

shown in Figure 8. The progression of creep is slower in the

cooling case. With cooling at time t = 4.72 h, the maximum

strain is about 0.02 with high strains distributed over the region

covered by debris, and the bottom center of the vessel wall has

displaced to about 0.06 m (Figure 8a). Then at a later time

t = 4.92 h in Figure 8b, the maximum strain has increased to

0.19 and the bottom center also displaced further to 0.26 m and

displacements get smaller near the top of the pool which

indicates a ‘ballooning’ type of failure [20]. The rest of the

vessel has negligible displacements. Moreover, this ballooning

mode of failure intensifies just minutes later (not shown) with

the maximum creep strain gone beyond 0.20.

Without cooling for H = 0.7 m (see Figure 8c), a maximum

creep strain of about 0.02 is reached at an earlier time

t = 3.22 h, in comparison to the cooling case. The bottom center

has displaced to about 0.06 m and high creep strains are also

fairly distributed over the region covered by debris. A little bit

later at t = 3.5 h, the maximum creep strain has increased to

0.18 (see Figure 8d) with the bottom center being displaced to

about 0.22 m and displacements get smaller upon reaching the

top of the pool which also indicates a ballooning type of failure.

Thus, for a smaller debris bed H = 0.7 m the global vessel

failure is delayed by about 1.4 h if CRGT and top cooling is

implemented, but the failure is the same, a ballooning mode of

global vessel failure.

For a larger debris bed H = 1.9 m, von Mises creep strains are

shown in Figure 9. The progression of creep in time is the same

in both cooling and non-cooling cases but the displacement

behavior and distribution of strains are different. With cooling

at t = 3.19 h in Figure 9a, the maximum creep strain has reached

about 0.01 with a maximum displacement of 0.04 m. At a later


time t = 3.47 h (Figure 9b), the maximum creep strain has

increased to about 0.16 with creep strains localized in the

region right below the pool top surface. The maximum

displacement is 0.07 m and displacements are almost uniform

(small deformations) in the region below the localized creep.

This localized creep mode of failure [20] intensifies just

minutes later (not shown) with the maximum creep strain gone

beyond 0.20. Note that the outer periphery of the melt pool

(Figure 5b) is not penetrated by cooled CRGTs. So, it is not

surprising that a considerably higher temperature of the vessel

wall near the top is observed. The effect of higher thermal load

in the region right below the top pool surface along with the

effect of considerably bigger mechanical load leads to a

localized creep mode of failure.

(a) t = 4.72 h (with CRGT and Top cooling) (b) t = 4.92 h (with CRGT and Top cooling)

(c) t = 3.22 h (without CRGT and Top cooling) (d) t = 3.5 h (without CRGT and Top cooling)

Figure 8. Magnified view of the von Mises creep strains: Comparison between the cases with and without CRGT and Top cooling for

H = 0.7 m.

Without cooling for H = 1.9 m, a maximum creep strain of

about 0.01 is reached at time t = 3.19 h (Figure 9c), similar to

the cooling case. But the high creep strains are fairly distributed

over the region covered by debris, unlike the cooling case. In

addition, the bottom center of the vessel wall has displaced to

about 0.07 m. At a later time t = 3.47 h (Figure 9d), the trend is

similar. The maximum creep strain has increased to about 0.16

with a maximum displacement of 0.28 m at the bottom center.

The overall displacement behavior is similar to the 0.7 m case

indicating a ballooning mode of failure in contrast to the

localized creep mode of failure in the 1.9 m with cooling.

However, the timing of vessel failure is the same for both

cooling and non-cooling case.

Table 1 summarizes the debris bed conditions evaluated from

the PECM calculations (Section I) at the time (t1) where 20 %

maximum creep strain has been reached and also conditions

30 minutes after this time. The amount of liquid melt and melt

superheat is generally lower when CRGT and top cooling is

implemented.

With cooling for H = 0.7 m, the amount of liquid melt at

t1 = 4.9 h is 18 tons (out of the maximum 30-ton debris) with a

melt superheat of 160 K. This does not change significantly

after 30 minutes where the melt increases to 20 ton with a melt

superheat of 168 K. In comparison to the non-cooling case, the

amount of liquid melt is 16 tons at t1 = 3.5 h with a melt

superheat of 33 K. This changes significantly after 30 minutes

where the melt increases to 24 tons and the melt superheat

increases to 177 K.


For H = 1.9 m, the time to reach 20 % maximum creep strain is

t1 = 3.5 h in the cooling case (just a bit longer than the t1 = 3.4 h

in the non-cooling case). The amount of liquid melt at this time

is 68 tons (out of the maximum 200-ton debris) with a melt

superheat of only 23 K. After 30 minutes, the liquid melt

increases significantly to 112 tons along with the melt superheat

to 103 K. For the non-cooling case, the amount of melt at t1 is

144 tons, much higher than the cooling case. However, its

superheat is only 25 K similar to the cooling case. After

30 minutes, the liquid melt increases to 183 tons and the melt

superheat increases considerably to 278 K.

Without cooling, the welds of the CRGT nozzles melt at around

2.1 h (see Table 1) for both 0.7 m and 1.9 m cases. At this time,

the debris bed is still in its solid form.

(a) t = 3.19 h (with CRGT and Top cooling) (b) t = 3.47 h (with CRGT and Top cooling)

(c) t = 3.19 h (without CRGT and Top cooling) (d) t = 3.47 h (without CRGT and Top cooling)

Figure 9. von Mises creep strains: Comparison between the cases with and without CRGT and Top cooling for H = 1.9 m at times

t = 3.19 h and 3.47 h.

Table 1. Debris bed conditions at time t1 where 20 % maximum creep strain has been reached and then after 30 minutes.

H

[m]

Cooling Time at max ~20 %

creep strain, t1 [h]

Time of CRGT

nozzle weld

melting, t2 [h]

Amount of liquid melt

at t1 (and after 30 min)

[ton]

Average melt superheat at

t1 (and after 30 min) [K]

0.7 Yes 4.9 - 18 (20) 160 (168)

No 3.5 ~2.1 16 (24) 33 (177)

1.9 Yes 3.5 - 68 (112) 23 (103)

No 3.4 ~2.1 144 (183) 25 (278)


III. Conclusions

The effects of CRGT and top cooling of a debris bed on the

modes of global vessel failure of a BWR lower head have been

investigated. Two cases with different debris heights are

considered; H = 0.7 m corresponds to a roughly 30-ton debris

and H = 1.9 m corresponds to a roughly 200-ton debris

(assumed maximum inventory). It is found that in the 0.7 m

case, global vessel failure is delayed by about 1.4 h if CRGT

and top cooling is implemented.

In the 1.9 m case, timing of the global vessel failure is relatively

the same (with no significant delay) but the mode of failure is

different. With CRGT and top cooling, the mode of failure is a

localized creep while it is ballooning in the case without CRGT

and top cooling. The localized creep is concentrated within the

vicinity of the top surface of the melt pool because this region is

not penetrated by CRGTs. For the case without CRGT and top

cooling, the timing of global vessel failure is relatively the same

in both the 0.7 m and 1.9 m cases, specifically, 20 % creep

strain has been reached around 3.5 h. Amount and superheat of

liquid melt is considerably higher in non-cooling case for 1.9 m

bed.

Furthermore, melting of the CRGTs starts at about 1.8 h in both

the 0.7 m and 1.9 m cases without cooling. But it is

considerably faster in the 1.9 m case in which about 90 % of the

CRGTs have melted after 2.6 h while it takes 1.4 h more in the

0.7 m case to reach 90 %. However, the welds of the CRGT

nozzles melt around 2.1 h in both the 0.7 m and 1.9 m cases. At

this time, the debris bed is still in its solid form. Predicted

amounts of liquid melt and melt superheat can provide

important information for the consideration of ex-vessel

accident progression.

In the current analysis, the mechanical strength of the

considered 2D axisymmetric lower head has not been adjusted

to take into account the CRGT penetrations in the vessel wall.

In the next step of this work, full 3D modeling of global

deformation of the vessel taking into account penetrations will

be considered and compared to the current 2D analysis to

investigate in details the influences of the (i) reduction of the

overall mechanical strength of the lower head due to the

penetrations and (ii) reduction of the mechanical load due to

smaller weight of the wall with penetrations. Another extension

of the work is to investigate the effect of CRGT and top cooling

on modes of failure of the Instrumentation Guide Tube (IGT).

ACKNOWLEDGMENTS This work is supported by the Swedish Nuclear Radiation

Protection Authority (SSM), Swedish Power Companies,

European Commission (SARNET-2), Nordic Nuclear Safety

Program (NKS), and Swiss Federal Nuclear Safety Inspectorate

(ENSI) under the APRI- MSWI program at the Royal Institute

of Technology (KTH), Stockholm, Sweden.

REFERENCES [1] P. Kudinov, A. Karbojian, W. Ma, and T.-N. Dinh “The

DEFOR-S Experimental Study of Debris Formation

with Corium Simulant Materials”, Nuclear

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effect of crgt cooling on modes of global vessel failure

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