The journal of Nuclear Science and Technology - Volume 9, Number 4, December 2019
ISSN 1810-5408
Nuclear Science
and Technology
Volume 9, Number 4, December 2019
Published by
VIETNAM ATOMIC ENERGY SOCIETY
VIETNAM ATOMIC ENERGY INSTITUTE
NUCLEAR SCIENCE AND TECHNOLOGY
Volume 9, Number 4, December 2019
Editorial Board
Editor-in-chief
Tran Huu Phat (VINATOM)
Executive Editors
Vuong Huu Tan (VARANS)
Tran Chi Thanh (VINATOM)
Cao Dinh Thanh (VINATOM)
Hoang Anh Tuan (VAEA)
Editors
Nguyen Kien Cuong (VINATOM)
Nguyen Nhi Dien (VINATOM)
Nguyen Thi Kim Dung (VINATOM)
Ho Manh Dung (VINATOM)
Nguyen Nam Giang (VINATOM)
Trinh Van Giap (VINATOM)
Le Ngoc Ha (108 Military Central Hospital)
Phan Son Hai (VINATOM)
Le Hong Khiem (IOP)
Dao Tien Khoa (VINATOM)
Tran Hoai Nam (Duy Tan University)
Dang Duc Nhan (VINATOM)
Nguyen Hao Quang (VINATOM)
Nguyen Mong Sinh (VINATOM)
Tran Duc Thiep (IOP)
Dang Quang Thieu (VINATOM)
Le Ba Thuan (VINATOM)
Le Huy Ham (VAAS)
Nguyen Quoc Hien (VINATOM)
Le Van Hong (VINATOM)
Nguyen Trung Tinh (VARANS)
Tran Ngoc Toan (VINATOM)
Duong Thanh Tung (VARANS)
Nguyen Nu Hoai Vi (VARANS)
Nguyen Tuan Khai (VINATOM)
Pham Dinh Khang (VINATOM)
Science Secretary
Hoang Sy Than (VINATOM)
....................................................................................................................................................................................................................
Copyright: ©2008 by the Vietnam Atomic Energy Society (VAES), Vietnam Atomic Energy Institute
(VINATOM).
Pusblished by
Vietnam Atomic Energy Society, 59 Ly Thuong Kiet, Hanoi, Vietnam
Vietnam Atomic Energy Institute, 59 Ly Thuong Kiet, Hanoi, Vietnam
....................................................................................................................................................................................................................
Contents
Bubble behavior in liquid-gas two-phase flow behind a cross-shaped obstacle in a
vertical circular duct
K. Takase, G. Kawasaki, K. Ueta……………………………………………..………………. 01
Evaluation of the Potential for Containment Bypass due to Steam Generator Tube
Rupture in VVER-1000/V320 Reactor during Extended SBO sequence using
SCDAP/RELAP5 code
Nguyen Van Thai, Doan Manh Long, Tran Chi Thanh …………………………………… 09
Study on transmutation efficiency of the VVER-1000 fuel assembly with different
minor actinide compositions
Tran Vinh Thanh, Vu Thanh Mai, Hoang Van Khanh, Pham Nhu Viet Ha ……………. 18
Determination of in situ detection efficiency for IM-NAA of non-standard
geometrical samples
Nguyen Duy Quang, Trinh Van Cuong, Tran Tuan Anh, Ho Van Doanh, Nguyen Thi
Tho, Ho Manh Dung…………………………………………………………………………….
27
34
Evaluating uncertainty of some radiation measurand using Monte Carlo method
Bui Duc Ky, Nguyen Ngoc Quynh, Duong Duc Thang, Le Ngoc Thiem, Ho Quang
Tuan, Tran Thanh Ha, Bui Thi Anh Duong, Nguyen Huu Quyet, Duong Van Trieu......
Evaluation of image reconstruction algorithms in cone-beam computed
tomography technique
Tran Thuy Duong, Bui Ngoc Ha…………………………………………………………….. 41
Relative output factors of different collimation systems in truebeam STx medical
linear accelerator
Do Duc Chi, Tran Ngoc Toan, Robin Hill, Nguyen Do Kien …………………………….. 48
Nuclear Science and Technology, Vol.9, No. 4 (2019), pp. 01-08
Bubble behavior in liquid-gas two-phase flow behind
a cross-shaped obstacle in a vertical circular duct
K. Takase*, G. Kawasaki, K. Ueta
*Department of Nuclear System Safety Engineering, Nagaoka University of Technology, Japan
(Received 07 November 2019, accepted 26 December 2019)
Abstract: Grid spacers installed in subchannels of fuel assemblies for nuclear reactors can promote
heat transfer. However, the fluid velocity and bubble behavior are greatly affected as the cross-
sectional area of the flow passage changes. Therefore, the void fraction distribution behind the
obstacle that simulates the grid spacer shape simply was measured by using a wire mesh sensor
(WMS) system. Moreover, a two-phase flow analysis was performed to investigate the effect of the
obstacle on the bubble behavior in a vertical duct.
Keywords: Bubbler, Two-phase flow, Obstacle, Vertical duct, WMS, Experiment, Analysis.
simulated spacer was visually observed and
the void fraction and interfacial velocity
distributions just behind the simulated
spacer was measured.
I. INTRODUCTION
Clarifying two-phase characteristics
in a nuclear reactor core is important in
particular to enhance the thermo-fluid safety
of nuclear reactors. Moreover, correct data
on bubbly flow in subchannels with spacers
are needed in order to verify two-phase flow
models in conventional nuclear safety
analysis codes and validate predicted data
by current CFD codes like a direct two-
phase flow analysis code (Douce, et al.,
2010). Spacers installed in subchannels of
fuel assemblies have the role of keeping the
interval between adjacent fuel rods constant.
Similarly, in case of PWR the spacer has
also the role as the turbulence promoter.
When the transient event occurs in a nuclear
reactor, two-phase flow is generated by
boiling of water due to heating of fuel rods.
Therefore, it is important to confirm the
bubbly flow behavior around the spacer.
The purpose of this study is to make the
effect of the spacer affecting the bubbly
flow clear and obtain code validation data.
So bubble dynamics around the simply
II. EXPERIMENTAL METHOD
1. Experimental Apparatus
The experimental apparatus mainly
consists of
a
measuring section and
water/air supply lines and is shown in Fig.
1. The measuring section includes a vertical
flow channel with a diameter of 58 mm
made of an acrylic resin, inlet and outlet
plenums, and an air injection nozzle with
120 injection holes with a diameter of 0.6
mm. Water flows through the inlet plenum
into the flow channel and goes up through
the measuring section to the outlet plenum.
Air is supplied from the air injection nozzle
into the flow channel. As a result, water is
mixed with air and then a water-air two-
phase flow is formed as can be seen in Fig.
2. Bubbly flow conditions are determined
by both flow rates of water and air.
©2019 Vietnam Atomic Energy Society and Vietnam Atomic Energy Institute
BUBBLE BEHAVIOR IN LIQUID-GAS TWO-PHASE FLOW BEHIND A CROSS-SHAPED…
f58 mm
Wire
mesh
Sensor
Cross-
shaped
obstacle
High
speed
Camera
Air nozzle
Fig. 1. Outline of an experimental apparatus
Fig. 2. Observed rising bubbles in a flow channel
temperatures of water and air are measured at
the water/air supply piping lines.
2. Wire-Mesh Sensor System
The void fraction distribution at the
horizontal cross- section in the flow channel is
measured with a wire-mesh sensor (WMS)
system (Prasser et al., 1998; 2000; 2001).
Complicated behavior of an interface between
water and air of each bubble is visually
observed with a high speed camera (HSC).
Moreover, two-phase pressure loss is measured
with seven differential pressure transducers
installed in the arbitrary locations in the
vertical direction of the flow channel. In
addition, both flow rates, pressures and
The WMS can obtain the void fraction
distribution in a cross-section of the flow channel
by measuring the electric current from the
transmitter-side wire layer through the fluid to the
receiver-side wire layer. Both wire layers consist
of nine wires respectively and are installed in the
flow channel with the distance of 3 mm in the
flow direction. The wire diameter is 0.3 mm.
Appearance and measuring wires of the WMS
are shown in Fig. 3. A mearing method with
the WMS is as follows.
Layout of each wire of transmitter and
Clipped layout of
each wire
receiver in the vertical direction
(Dimension: mm)
Appearance of a wire-mesh sensor
Fig. 3. Appearance and layout of a wire-mesh sensor in a circular flow channel
2
K. TAKASE et al.
1)Current is given to the transmitter-side
wire layer;
Moreover,
the
cross-correlation
coefficient used to obtain the interfacial
velocity is obtained from the relationship
between both data measured by a couple of
the WMS layers and is expressed by the
following equation.
2) The current flows from the transmitter-
side wire layer through the fluid to the
receiver-side wire layer;
3)The voltage is measured at the location
where each wire intersects; and,
n
x x
y
y
t
t
t1
4)Void fraction between both wires
is obtained.
(3)
cor( )
n
2
2
x x
y
y
t
t
t1
The following equation is used for
the conversion from the voltage to the
void fraction.
Here, x and y represent time series void
fraction data measured with the WMS
installed upstream in the flow direction and
the WMS installed downward. The n shows
the number of comparison data. The cross-
correlation coefficient, cor() has the value
from -1 to 1, and it shows 1 when both data of
x and y have the positive correlation, -1 when
those are the negative correlation, and 0 when
those have no correlation.
UL Ui, j
(1)
i, j
UL UG
Here, is the local void fraction, (i, j)
is the position where two wires intersect in
the cross-section of the flow channel, U [V]
is the voltage measured by the WMS at two-
phase flow condition, Each of UL [V] and
UG [V] is the voltage measured by the WMS
at the single-phase flow condition of liquid
or gas.
On the other hand, gas-liquid interface
velocity is determined using the time-series
void fraction data measured by a couple of the
WMS layers which are installed upstream and
downstream of the flow channel. Two void
fraction data by both WMS layers are
compared. In Fig 4, time difference, , in case
that the cross-correlation coefficient between
both void fraction data shows the highest
match is identified. Finally, the gas-liquid
interface velocity, ub, is calculated by Eq. (2)
using S and . Here, S is the distance between
both receiver-side wire layers.
S
Fig. 4. Example of time variation of void fraction
ub
(2)
data at the different vertical positions
3
BUBBLE BEHAVIOR IN LIQUID-GAS TWO-PHASE FLOW BEHIND A CROSS-SHAPED…
Table II. Experimental Conditions of
3. Experimental Conditions
The present experiments
Superfacial Velocity of Gas in Mixture, JG
were
performed under the conditions of room
temperature, atmospheric pressure and non-
heated isothermal flow. Working fluids are
water and air. Superficial velocities of water
and air, JL and JG, are set to be 0.13-1.0 m/s
for JL, and 0.025-0.4 m/s for JG. Experimental
conditions of JL and JG are shown in Table I
and Table II.
Case
JL1
JL2
JL3
JL4
JL (m/s)
0.13
0.25
0.5
1.0
In order to investigate the influence of
an obstacle in the flow channel to the void
fraction distribution, the obstacle with the
dimensions of the thickness 5 mm, the height
50 mm and the horizontal distance 58 mm as
shown in Fig. 5 was installed in the position
of 20 or 100 mm upstream from the WMS.
The shape of the obstacle simply simulates
that of the grid spacer and shows a cross-
shape, but the dimensions are different.
Table I. Experimental Conditions of Superfacial
Velocity of Water in Mixture, JL
Case
JG1
JG2
JG3
JG4
JG (m/s)
0.025
0.05
0.1
0.2
Fuel rod
Grid spacer
Dimensions of a cross-
shaped obstacle
Layout of fuel rods
with a grid spacer
Fig. 5. Appearance of a cross-shaped obstacle which simulates the shape of the grid spacer simply
small bubbles are distributed around the
wall. On the other hand, the force which
includes the large lift force, etc. generates
to the center region of the flow channel in
case of the large bubbles. As a result, many
large bubbles are distributed to the center
region.
III. RERULTS AND DISCUSSION.
The time-averaged void fraction
distribution in the radial direction in the
flow channel is indicated in Fig. 6. Here, in
case of Fig. 6 (a) JL is L2 and in case of
Fig. 6 (b) JL is L4. When JG is low, the void
fraction distribution shows the wall-peak
distribution. On the other hand, when JG is
high, it becomes the core-peak distribution.
The force which includes the lift force, etc.
generates towards the wall direction in case
of the small bubbles. Therefore, many
Figure 7 is the bubble behavior observed
with HSC. Here, the superficial velocity
condition in Fig. 7 (a) are L3 and G1. Similarly,
that in Fig. 7 (b) is L3 and G4. In case of Fig. 7
(a) small bubbles are observed. Meanwhile,
4
K. TAKASE et al.
since coalescence of small bubbles is promoted
slug flow with a shape of the bullet is observed
in case of Fig. 7 (b).
with increasing the superficial gas velocity, the
0.5
0.5
G1
G2
G3
G4
G5
G1
G2
G3
G4
0.4
0.3
0.2
0.1
0
0.4
0.3
0.2
0.1
0
-20
-10
0
10
20
-20
-10
0
10
20
Radial position [mm]
Radial position [mm]
(a) In case of JL=JL2
(b) In case of JL=JL4
Fig. 6. Radial void fraction distributions obtained by the conditions of JL2 and JL4
(b) In cases of JL3 and JG4
Fig. 7. Observed bubble behavior at JL3, JG1 and JG4
(a) In cases of JL3 and JG1
Void fraction distributions in the flow
channel with and without the obstacle when
JL=L3 and JG=G1 are indicated in Fig. 8.
Here, since measuring data in the circular
flow channel are shown as a square figure,
four comers in the square figure are the
outside of the measuring area. The void
fraction is shown with the color contour.
Here, blue is 0, red is 0.2 and other colors
show the value between 0 and 0.2. The void
fraction distribution of Fig. 8 (a) which is the
no obstacle case shows the wall peak because
Gris the small bubble (i.e., bubbly flow)
condition. On the other hand, Fig. 8 (b) and
(c) are the results at each position of 20 or
100 mm behind the obstacle. The void
fraction becomes high just behind the
obstacle at Fig. 8 (b). The reason for this is
that the bubbles separated at the leading edge
of the obstacle coalesce just behind the
obstacle. Furthermore, the void fraction
becomes uniform at Fig. 8 (c). This is
because the distance of 100 mm is long and
the distribution is flattened.
5
BUBBLE BEHAVIOR IN LIQUID-GAS TWO-PHASE FLOW BEHIND A CROSS-SHAPED…
Fig. 8. Void fraction distributions at JL3 and JG1
Fig. 9. Void fraction distributions at JL3 and JG5
(a) No obstacle
(b) 20 mm behind
(c) 100 mm behind
Fig.10. Three-dimensional void fraction distributions when JL=JL3 and JG=JG5
Void fraction distributions in the flow
channel with and without the obstacle when
JL=L3 and JG=G5 are indicated in Fig. 9.
Here, blue is 0, red is 0.4 and other colors
6
K. TAKASE et al.
show the value between 0 and 0.4. The void
from the channel center to the side wall; r0 and
r are the radius and radial distance from the
center of the circular flow channel.
fraction distribution of Fig. 9 (a) which is the
no obstacle case shows the core peak because
G5 is the large bubble (i.e., slug flow)
condition. Figures 9 (b) and (c) are the results
at the positions of 20 and 100 mm behind the
obstacle. In Fig. 9 (b) the void fraction
becomes low just behind the obstacle and high
at the region except for the area just behind
that. This is because large bubbles are divided
by the obstacle. In Fig. 9 (c) the void fraction
becomes high at the core region because the
divided bubbles coalesce again.
The flow velocity was JL=L3 and JG was
changed to three kinds of G1, G3 and G5. The
open symbol shows the result in the case of no
obstacle. When the flow rate is low, the void
fraction distribution in the horizontal cross
section of the flow channel becomes almost
uniform, and shows the trend of the core peak as
the flow rate increases. On the other hand, the
solid symbol shows a case where there is an
obstacle, and since the large bubble is divided as
the flow rate increases, a high void fraction is
seen near the wall of the flow channel.
Figure 10 shows the results of three-
dimensional void fraction distribution in the
case of such large bubbles. Here, Fig. 10 (a) is
the result when there is no obstacle. Figure 10
(b) and (c) are the results at the positions of 20
and 100 mm behind the obstacle, respectively.
Here, Fig. 10 (a) is found to be the slug flow.
Also, Fig. 10 (b) shows that large bubbles are
divided into four. In Fig. 10 (c), the appearance
of becoming large bubbles is seen.
Velocity distributions of the bubble
interface in the flow channel with and without
the obstacle when and JG=G5 arc indicated in
Fiq.12. Here, blue is 0 m/s, red is 1 m/s and
other colors show the velocity between 0 and
1 m/s. The velocity distribution of Fig.12 (a)
which is the no obstacle case shows the core
peak because G5 is the slug flow condition.
Fig.12 (b) and (c) arc the results at each
position of 20 or 100 mm behind the obstacle.
In Fig.12 (b) the velocity distribution of the
bubble interface shows four peaks in the region
where no obstacle exists. Moreover, the
Figure 11 shows the void fraction
distributions in the radial direction of the flow
channel. Here, the vertical axis represents the
local void fraction and the horizontal axis
represents the dimensionless radial distance
Without an obstacle:
JG＝JG1
JG＝JG3
JG＝JG5
With an obstacle:
JG＝JG1
JG＝JG3
JG＝JG5
Fig. 11. Radial void fraction distributions against different JG at
the position of 20 mm behind the obstacle when JL=JL3
7
BUBBLE BEHAVIOR IN LIQUID-GAS TWO-PHASE FLOW BEHIND A CROSS-SHAPED…
velocity distribution of the bubble interface
becomes high in the core region as well as the
tendency of the void fraction distribution, as
can be seen in Fig. 12 (c).
Fig. 12. Gas-liquid interface velocity distributions in the flow channel
with and without the obstacle when JL=JL3 and JG=JG5
IV. CONCLUSIONS
5) The gas-liquid interface velocity can
be predicted from the distribution of the void
fraction by using the equation of the cross-
correlation coefficient.
Bubbly flow behavior, void fraction
distribution and interfacial velocity distribution
around the simulated spacer installed in the
flow channel were investigated experimentally
and at non-heated and isothermal water-air
two-phase flow condition. The results of the
present study arc summarized as follows:
REFERENCES
[1]. Prasser, H. M., Böttger, A. and Zschau, J., “A
New Electrode-Mesh Tomograph for Gas-
Liquid Flows,” Flow Measurement and
Instrumentation, 9, 1998.
1) When JG is low, the void fraction
distribution in the flow channel without the
obstacle
becomes
the
wall-
peak
[2]. H. M. Prasser, H. M., Krepper E. and Lucas,
D., “Fast Wire-Mesh Sensors for Gas-Liquid
Flows and Decomposition of Gas Fraction
Profiles According to Bubble Size Classes”,
Second Japanese-European Two-Phase Flow
Group Meeting, Tsukuba, Japan, September,
25-29, 2000.
distribution. When JG is high, it becomes
the core-peak distribution;
2) When JG is low, the void fraction
becomes high just behind the obstacle. The
reason is that the bubbles separated at the
leading edge of the obstacle coalesce just
behind the obstacle;
[3]. Prasser, H. M., Scholz, D., and Zippe, C.,
2001, “Bubble Size Measurement Using Wire-
Mesh Sensors”, Flow Measurement and
Instrumentation, 12, 4, 2001.
3) When JG is high, the void fraction
becomes low just behind the obstacle and high
at the region except for the area just behind
that. This is because large bubbles are divided
by the obstacle;
[4]. Douce1, A., Mimouni1, A.., Guingo1, M.,
Morel2, M., J. Laviéville1, J. and Baudry1, C.,
“Validation of Neptune CFD 1.0.8 for
Adiabatic Bubbly Flow and Boiling Flow”,
CFD4NRS-3, September, 14-16, Washington
DC, USA, 2010.
4) Behind the obstacle, the distribution
of the void fraction at the channel cross section
is flattened as the distance from the obstacle
increases; and,
8
Nuclear Science and Technology, Vol.9, No. 4 (2019), pp. 09-15
Evaluation of the Potential for Containment Bypass due to
Steam Generator Tube Rupture in VVER-1000/V320 Reactor
during Extended SBO sequence using SCDAP/RELAP5 code
Nguyen Van Thai1,*, Doan Manh Long2, Tran Chi Thanh2
1 Department of Nuclear Engineering and Environmental Physics, School of Engineering Physics,
Hanoi University of Science and Technology
2 Vietnam Atomic Energy Institute
(Received 05 November 2019, accepted 30 December 2019)
Abstract: A severe accident-induced of a Steam Generator (SG) tube releases radioactivity from
the Reactor Coolant System (RCS) into the SG secondary coolant system from where it may escape
to the environment through the pressure relief valves and an environmental release in this manner is
called “Containment Bypass”. This study aims to evaluate the potential for “Containment Bypass”
in VVER/V320 reactor during extended Station Blackout (SBO) scenarios that challenge the tubes
by primarily involving a natural circulation of superheated steam inside the piping loop and then
induce creep rupture tube failure. Assessments are made of SCDAP/RELAP5 code capabilities for
predicting the plant behavior during an SBO event and estimates are made of the uncertainties
associated with the SCDAP/RELAP5 predictions for key fluid and components condition and for
the SG tube failure margins.
Keywords: Containment Bypass, SBO, SCDAP/RELAP5.
failure of steam generator tubes prior to the
failure of one of other components (hot leg
piping, pressurizer surge-line piping, and the
reactor vessel) leads to discharge of some
fission products into the steam generator
secondary system from where they may be
discharged to the environment through the
pressure-relief valves. This sequence is
potentially more risk-significant since it
involves a containment bypass scenario. The
relative timing of these structural failures
therefore affects the event sequence and
whether the containment is bypassed.
I. INTRODUCTION
In 1990s, U.S. Nuclear Regulatory
Commission (NRC) and the Russian Federal
Nuclear and Radiation Safety Authority
Gosatomnadzor (GAN) agreed to work
together in BETA Project to perform a
probabilistic risk assessment (PRA) of Unit I
of the Kalinin Nuclear Power Station (KNPS),
which is a VVER1000/V-338. Analysis within
the BETA Project involves different levels of a
wide-scope PRA and the PRA model
represents the set of accident sequences
following the initial events (IEs) up to the end
state of each sequence. The reports indicated
that the most significant release category for
offsite consequences is containment failure as a
result of an IE with a leak from the primary to
the secondary circuit [1]. In this aspect, steam
generator tubes comprise a majority of the
reactor coolant system pressure boundary and
This paper present thermal-hydraulic
evaluations of VVER-1000/V-320 during
extended
SBO
event
using
the
SCDAP/RELAP5 systems analysis code. In
general, the design features of VVER-1000/V-
338 are similar to the standard VVER-1000/V-
320, excepted of the main circulation loops
©2019 Vietnam Atomic Energy Society and Vietnam Atomic Energy Institute
EVALUATION OF THE POTENTIAL FOR CONTAINMENT BYPASS DUE TO STEAM…
configuration which is equipped with main
gate valves on the cold and hot legs (CLs and
HLs) [2]. The potential for “Containment
Bypass” is associated with natural circulation
of superheated steam inside the piping loop
that might induce creep rupture tube failure.
SCDAP/RELAP5 predictions provide key fluid
and components condition for estimation of the
SG tube failure margins.
The turbine-driven auxiliary feedwater
(TDAFW) system is assumed to independently
fail, so no MFW or auxiliary feedwater (AFW)
system is available;
A station battery life of four hours is
assumed; after that time all automatic and
operator control of the pressurizer PORVs and
SG secondary system PORVs is lost;
During the initial portion of the accident
scenario, buoyancy-driven coolant-loop natural
circulation carries hot water from the core
through the SGs, transferring heat to the SG
secondary water inventory. The SG water
inventory is boiled and the steam is released
through the SG PORVs. The secondary water
inventory declines and is eventually fully
depleted since the MFW and AFW systems are
not operative. After that time, the core decay
power heats and swells the RCS water,
increasing its temperature and pressure. The
basic physical processes during this regard the
transport of hot steam from the core outward
into the other regions of the reactor vessel
(RV) and coolant loops. Two coolant loop
natural circulation flow patterns that may be
encountered subsequent to the uncovering and
heat-up of the reactor core, based upon whether
or not the loop seals (the cold leg piping
connecting the outlets of the SGs to the inlets
of the RCPs) remain liquid-plugged.
II. DESCRIPTIONS OF ACCIDENT
SCENARIO AND PLANT BEHAVIOR
The low-probability SBO base case
accident event scenario results in a severe
accident because none of the systems that
normally provide core cooling are assumed to
be operable nor is any alternate equipment
(e.g., security-related mitigation methods)
assumed to be available [4]. The accident event
is initiated by a loss of off-site alternating
current (AC) power, which immediately results
in reactor and turbine trips and the coast-down
of the four reactor coolant pumps (RCPs).
Also, assumptions are made as follows:
The diesel-electric generators fail to
start and all AC power sources are lost;
Letdown flow is isolated and the
pressurizer level control and RCP seal injection
functions of the charging system are lost;
The high-pressure and low-pressure
safety injection systems are unavailable as a
result of the AC power loss;
If liquid is cleared from a loop seal
(along with the region of the RV lower plenum
that extends above the bottom of the core
barrel, the "downcomer skirt"), hot steam is
transported from the core through the HL, SG
tubes and cold legs in the normal direction of
flow (i.e., that seen during plant operation).
This flow pattern transports the hot steam
directly (without benefit of mixing) through all
of the SG tubes, leading to SG tube failure
The accumulator systems (four HAs)
are available for injecting coolant into the cold
legs should the RCS pressure fall below the
initial accumulator pressure, 5.9 MPa;
The main feedwater (MFW) flow stops
and the motor-driven auxiliary feedwater
(MDAFW) system is unavailable as a result of
the AC power loss;
10
NGUYEN VAN THAI et al.
prior to HL or pressurizer surge line failure.
However, if a loop seal remains liquid-
plugged, the more complex flow pattern
develops instead. Hot steam is transported
through the upper portion of the HL cross
section to the SG inlet plenum, where it is
mixed with cooler steam emanating from
circulations set up within the SGs, with some
of the tubes flowing in the normal direction
and the remaining tubes flowing in the reverse
direction. The mixing process within the SG
inlet plenum determines the temperatures of
the steam entering the SG tubes and the steam
that is returned to the RV through the lower
portion of the HL cross section. Fluid mixing
in the SG inlet plenum buffers the entry of hot
steam into the SG tubes, thus delaying SG tube
failure and making it more likely that some
other component (HL, pressurizer surge line or
RV) will be the first to fail.
The issues of primary interest for
containment bypass are: (1) do the loop seals in
all coolant loops remain liquid plugged and if
so (2) does the fluid mixing in the SG inlet
plenum sufficiently slow the SG tube heat-up
process so that the HL, pressurizer surge line
or RV will fail prior to a SG tube? [4]. The
analyses in this paper address these issues.
Fig. 1. RCS Layout of VVER-100/V-320 [3]
TURBINE
196
194
MSIV
187
MSIV
287
MSIV
387
MSIV
487
186
286
386
486
490
190
489
BRU-A1
189
BRU-A1
182
184
484
482
REG
REG
REG
REG
180
BRU-In1
191
BRU-In1
491
192
Valve Valve Valve Valve
185 285 385 485
492
480
Steam
dome
Steam
dome
169
168
167
469
468
467
Separator
Separator
Volume above
perforated sheets
Volume above
perforated sheets
508
177
477
ACC
ACC
ACC
ACC
152
452
PORV
Volume between
tube bundles and
perforated sheets
Volume between
tube bundles and
perforated sheets
506
511
466
176
175
174
136
166
125
124
123
122
121
116
416
415
414
413
412
411
410
436
435
476
JNG80
JNG60
087
JNG70
088
JNG50
086
165
115
465
464
463
462
461
425
424
423
179
178
173
172
171
135
134
133
479 475
089
164
114
434
433
432
431
430
478
473
474
504
450
150
163
113
162
112
422
421
132
131
130
472
471
161
111
502
110
56
54
100
400
3
102
61-62
1-2
49-50
45-48
41-44
51-52
63-64
3-4
402
8
7
4
3
2
1
1
2
4
142
144
5-6
7-8
444
442
141
441
11
1
11
36
37-40
9
9
9-11
30 32 33 31
26 28 29 27
10-12
34
4
1
2
3
1
6
6
9
8
7
6
5
2
3
3
140
440
35
4
3
2
1
1
1
4
2
1
4
5
6
23-25
22-24
18-20
5
6
19-21
15-17
14-16
Fig. 2. Nodalization scheme of VVER-1000/V-320 Plant
11
EVALUATION OF THE POTENTIAL FOR CONTAINMENT BYPASS DUE TO STEAM…
zero, the structure has not experienced any
creep damage. If the value is one, the structure
has failed due to creep damage. Two different
theories (Larson Miller and Manson-Haferd)
are applied in SCDAP/RELAP5 which are
dependent upon the structural materials and the
range of stress and summarized in Table I.
III. PLANT NODALIZATION AND
PARAMETERS SETUP WITH
SCDAP/RELAP5
The SCDAP/RELAP5 computer code
calculates the overall RCS thermal-hydraulic
response for severe accident situations that
include core damage progression and RV heat-
up and damage. The computer code is the result
of a merging of the RELAP5 and SCDAP
computer codes. Prediction of structural failure
are made based on the structure configuration,
its material properties and the fluid conditions
that are locally present. The code also includes
models for calculating the creepture failure of
structural components and these are used to
predict failure time for the hot legs, pressurizer
surge line, and SG tubes [5].
Table I. Equations for the time to creep rupture [5]
Since SG tubes are made from heat-
resistant steel (08X18H10T) which is similar to
heat-resistant Stainless Steel, equations for
creep rupture time of 316 Stainless Steel in
Table i are selected in this study. Estimated
stress in SG tubes (cylindrical shaped
structure) is calculated as follows:
The
SCDAP/RELAP5
VVER-
1000/V320 plant model represents the fluid
volumes and structures in the core, RV and
primary and secondary coolant system regions
in the plant as shown in Fig 1. The nodalization
diagrams for the final SCDAP/RELAP5
VVER-1000/V-320 four-loop plant model are
provided in Fig. 2.
(P r P ro )
i
rio rio
(2)
In this work, creep rupture failure
calculations are performed for the hot average SG
tubes in which pressure and temperature
conditions of fluid and heat structures in four SGs
are used as inputs for the analyses. A parameter
that measure creep damage is calculated each time
step for each structure being monitored for creep
rupture by following equation:
Fig. 3. Model ofSG ubes
IV. RESULTS AND DISCUSSION
A. Plant behaviors
t
Dc
t t
Dc
t
(1)
tr
t
The SCDAP/RELAP5 plant model was
run to a steady solution. The plant model using
the nodalization scheme to establish full-power
steady-state conditions from which the SBO
transient accident sequence is initiated
(Figures.4&5).
Where Dc
problem time t, t is time step at current
problem time and tr is time required for the
t
is creep damage at
t
structure to fail by creep rupture at current state
of temperature and stress. If the value of Dc is
12
NGUYEN VAN THAI et al.
Fig. 4. Steady-state conditions (Pressure) at primary
Fig. 6. Transient condition of reactor power and
and secondary loops
pressure in PRZ
Fig. 7. Transient condition of mass flow rate
through the primary loops and SG levels
Fig. 5. Steady-state conditions (Mass Flow Rate
and Temperature) at primary and secondary loops
Buoyancy-driven coolant-loop natural
circulation carries hot water from the core
through the SGs, transferring heat to the SG
secondary water inventory until the SG heat
sink is lost around 11000s (Fig. 7). It can be
seen that after the SG heat sink is lost, the
pressure the cooling afforded by system heat
loss to containment and RCP shaft seal leak
flow is insufficient to remove the RCS heat
load, causing the RCS and PRZ pressure to
increase (Fig. 6). After SG dry-out, the RCS
pressure increase is limited by multiple cycling
of the PORVs and by two cycles of the SRVs
during the period with the most-challenging
RCS pressurization conditions. This challenge
The SBO base case event sequence was
simulated with SCDAP/RELAP5, starting from
time 250s when the loss of off-site power
occurs. Figure 5 showed the transient condition
of reactor power and pressure level in PRZ.
After a short period when the PRZ
pressure initially falls and rises slightly due to
the effects of the reactor and turbine trips, the
PRZ pressure declines in response to the
cooling provided by heat removed to the SGs.
The RCS fluid mass lost through the
pressurizer PORVs and SRVs and through the
RCP shaft seal leakage paths depletes the
RCS inventory.
13
EVALUATION OF THE POTENTIAL FOR CONTAINMENT BYPASS DUE TO STEAM…
is presented when the increasing temperatures
cause the RCS fluid to swell, completely filling
the pressurizer with water. The fuel
temperature increases after the heat sink is lost
and the onset of fuel rod oxidation as well as
fuel rod cladding rupture occurred as an
unvoidable consequence (Fig. 8).
occur after the time when SG tube structural
failures is experienced.
Fig. 9. Structure temperature and creep
damage index
V. CONCLUSIONS
The
paper
presents
preliminary
Fig. 8. Temperature of fuel pins and hot legs
assessments the VVER-1000/V-320 plant
behavior during an SBO event with the
SCDAP/RELAP5 for key fluid and
components condition and for the SG tube
failure margins. Further analyses will be
performed to evaluate the potential for
“Containment Bypass” in VVER-1000/V-320
reactor during extended Station Blackout
(SBO) scenarios.
B. Creep Rupture Analysis of SG Tubes
Creep rupture model allows one to
specify a "stress multiplier" in following
meaning: a multiplier of 1.0 provides a creep
rupture failure prediction based on no
degradation of the structural strength of the
material and multipliers greater than 1.0
represent degraded structural strengths. In this
study, a stress multiplier from 1.0 to 5.0 (in
increments of 0.1) are used for SG tube
rupture prediction to investigate the spectrum
of material strengths, from undergraded to
highly-degraded.
ACKNOWLEGEMENTS
This work was supported by the
Ministry of Science and Technology (MOST)
R&D Project with contract number
05/HĐ/ĐTCB signed on 05/01/2018.
Typical results on Larson-Miller creep
rupture damage index for the hottest SG tubes
is presented in Fig. 9. It was found that the
average-tube predicted failure margin is 2.3
and the hottest-tube predicted failure margin is
1.7. Also, SG tube failure margin are
insensitive to variations in the fuel damage
progression behavior primarily because these
.REFERENCES
[1]. Federal Nuclear and Radiation Safety
Authority of the Russian Federation, Kalinin
VVER-1000 Nuclear Power Station Unit 1
PRA (Beta Project), NUREG/IA-0212,
published by US NRC, 1992.
14
NGUYEN VAN THAI et al.
[2]. Design, Safety Technology and Operability
Features of Advanced VVERs, OKB
“GIDROPRESS’’, 2011.
[4]. C.D., Fletcher et al. SCDAP/RELAP5 Thermal-
Hydraulic Evaluations of the Potential for
Containment Bypass During Extended Station
Blackout Severe Accident Sequences in a
Westinghouse Four-Loop PWR. NUREG/CR-
6995, published by US NRC, 2010.
[3]. S.,
Pylev.
Assessment
Study
of
RELAP5/MOD3.2 Based on the Kalinin NPP
Unit-1 Stop of Feedwater Supply to the Steam
Generator No. 4, NUREG/IA-0167, published
by US NRC, 1999.
[5]. C. Allison et al., “SCDAP/RELAP5/MOD2
Code Manual”, Volume: I, II and III,
NUREG/CR-5237, EGG-2555, June 1989.
15
Nuclear Science and Technology, Vol.9, No. 4 (2019), pp. 16-26
Study on transmutation efficiency of the VVER-1000 fuel
assembly with different minor actinide compositions
Tran Vinh Thanh1, Vu Thanh Mai2, Hoang Van Khanh1, Pham Nhu Viet Ha1*
1Institute for Nuclear Science and Technology, Vietnam Atomic Energy Institute,
179 Hoang Quoc Viet str., Cau Giay dist., Hanoi 100000, Viet Nam
2Hanoi University of Science, Vietnam National University,
334 Nguyen Trai, Thanh Xuan, Hanoi 10000, Viet Nam
E-mail*: [email protected]
(Received 06 November 2019, accepted 30 December 2019)
Abstract: The feasibility of transmutation of minor actinides recycled from the spent nuclear fuel in
the VVER-1000 LEU (low enriched uranium) fuel assembly as burnable poison was examined in our
previous study. However, only the minor actinide vector of the VVER-440 spent fuel was considered.
In this paper, various vectors of minor actinides recycled from the spent fuel of VVER-440, PWR-
1000, and VVER-1000 reactors were therefore employed in the analysis in order to investigate the
minor actinide transmutation efficiency of the VVER-1000 fuel assembly with different minor
actinide compositions. The comparative analysis was conducted for the two models of minor actinide
loading in the LEU fuel assembly: homogeneous mixing in the UGD (Uranium-Gadolinium) pins and
coating a thin layer to the UGD pins. The parameters to be analysed and compared include the
reactivity of the LEU fuel assembly versus burnup and the transmutation of minor actinide nuclides
when loading different minor actinide vectors into the LEU fuel assembly.
Keywords: VVER-1000 LEU fuel assembly, burnable poison, minor actinide transmutation.
I. INTRODUCTION
environment, separation and transmutation of
the plutonium and MAs in the used fuel are
esssential [2]. It has been realized that the
transmutation of these actinide into either
short-lived fission products or valued fissile or
stable isotopes can be accomplished in fast
reactors, subcritical reactors or thermal
reactors [1,3-7].
The radioactive waste and spent nuclear
fuel discharged from nuclear power plants
causes a big issue for the countries holding
such nuclear installations. It is widely
recognized that a light water reactor (LWR)
with electric capacity of 1000 MWe, on
average, produces 20-30 metric tonnes of spent
nuclear fuel annually, which consist of
approximately 95 wt% uranium, 1 wt%
plutonium, 4 wt% fission products and minor
actinide (MA) [1]. In the spent fuel, only with
a very small amount of MAs, they still
dominate the decay heat load to the repository
and cumulative long-term radiotoxicity to the
environment. To lessen the burden for disposal
and storage of spent nuclear fuel as well as to
reduce its cumulative radiotoxicity to the
The VVER-1000 reactor (the Russian
Pressurized Water Reactor, PWR) is
nowadays operated in various East
European and Asian countries [8,9]. In
addition to the Western PWRs that have
been extensively studied for their MA
transmutation capabilities [10-14], the
VVER-1000 is also considered as
a
potential candidate for transmutation of
actinide in the spent fuel stock-pile and
©2019 Vietnam Atomic Energy Society and Vietnam Atomic Energy Institute
TRAN VINH THANH et al.
various methods of loading and burning
insignificant change in the reactivity of the fuel
assembly while providing considerable MA
transmutation rates.
transuranic elements in the Western PWRs
may be adopted similarly to the Russian
VVERs. In the past studies, transmuting the
MAs in the burnable poison rods [15,16] or in
some other locations in the PWR fuel
assemblies has been found technically feasible
and recommended as potential transmutation
methods for LWRs, in particular the unique
advantage of loading MAs to partially replace
the excess reactivity control functions of
gadolinium and boric acid.
II. CALCULATION METHOD
The VVER-1000 LEU fuel assembly
specified in the OECD VVER-1000 LEU and
MOX (mixed oxide) Assembly Computational
Benchmark [18] is utilized in the present
investigation to examine the possibility of MA
transmutation as burnable poison in the
VVER-1000 reactor. The configuration of the
LEU fuel assembly are shown in Fig. 1. The
LEU assembly consists of 300 fuel pin cells
with 3.7wt% 235U, 12 UGD pin cells with
3.6wt% 235U and 4wt% Gd2O3, 18 water filled
guide tubes for control insertion and one
central water filled instrumentation tube. The
LEU fuel assembly is modeled by the SRAC
code. The one-sixth of the LEU fuel assembly
modeled by the PIJ module of SRAC is
presented in Fig. 2; the burnup calculations
were performed with the BURN-UP module of
SRAC; and the 107 energy groups based on
the ENDF/B-VII.0 nuclear data library were
used in the SRAC calculations.
In a previous study [17], the feasibility
of MA transmutation in VVER-1000 LEU fuel
assembly [18] as burnable poison was studied
and the results showed that the total MA
transmutation rate of ~20% could be obtained.
However, only the MA vector of the VVER-
440 spent fuel was considered. In the present
work, different MA compositions recycled
from the spent fuels of VVER-440 [4], PWR-
1000 [15] and VVER-1000 [19] with different
burnup levels and cooling time were therefore
employed in the analysis in order to estimate
the effects of various MA contents in the spent
fuel to the infinite multiplication factor (k-inf)
of the VVER-1000 LEU fuel assembly versus
burnup as well as the MA transmutation
efficiency. The SRAC code [20] was used for
modeling of the VVER-1000 LEU fuel
assembly based on the ENDF/B-VII.0 library.
The comparative analysis was conducted for
the two models of MA loading in the LEU fuel
assembly: homogeneous mixing in the UGD
(Uranium-Gadolinium) pins and coating a thin
layer to the UGD pins. The MA loading into
the LEU fuel assembly will be performed
without significant modification of the
assembly configuration to minimize the cost
for fuel fabrication process and respective
changes in reactor core design. The constraint
for these MA loadings is to ensure
In this investigation, we intend to
load the MAs in the UGD pins of the LEU
fuel assembly for their transmutation
without significant change in the fuel
assembly configuration. The purpose is to
investigate the transmutation capability of
the VVER-1000 LEU fuel assembly. To this
end, we consider two approaches to load the
MAs into the fuel assembly while tuning the
gadolinium
content
(1)
and
boron
MAs
concentration:
mixing
homogeneously with UO2 and Gd2O3 in the
UGD pins; and (2) coating a thin layer of
MAs around the UGD pellets. In these
cases, different vectors of MAs were
17
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