Design analyses for full core conversion of the Dalat nuclear research reactor
Nuclear Science and Technology, Vol. 4, No. 1 (2014), pp. 10-25
Design Analyses for Full Core Conversion of
The Dalat Nuclear Research Reactor
Luong Ba Vien, Le Vinh Vinh, Huynh Ton Nghiem, Nguyen Kien Cuong
Reactor Center – Nuclear Research Institute – Vietnam Atomic Energy Institute
01 Nguyen Tu Luc, Dalat, Lamdong
(Received 5 March 2014, accepted 10 March 2014)
Abstract: The paper presents calculated results of neutronics, steady state thermal hydraulics and
transient/accidents analyses for full core conversion from High Enriched Uranium (HEU) to Low
Enriched Uranium (LEU) of the Dalat Nuclear Research Reactor (DNRR). In this work, the
characteristics of working core using 92 LEU fuel assemblies and 12 beryllium rods were investigated
by using many computer codes including MCNP, REBUS, VARI3D for neutronics, PLTEMP3.8 for
steady state thermal hydraulics, RELAP/MOD3.2 for transient analyses and ORIGEN, MACCS2 for
maximum hypothetical accident (MHA). Moreover, in neutronics calculation, neutron flux, power
distribution, peaking factor, burn up distribution, feedback reactivity coefficients and kinetics parameters
of the working core were calculated. In addition, cladding temperature, coolant temperature and ONB
margin were estimated in steady state thermal hydraulics investigation. The working core was also
analyzed under initiating events of uncontrolled withdrawal of a control rod, cooling pump failure,
earthquake and MHA. Obtained results show that DNRR loaded with LEU fuel has all safety features as
HEU and mixed HEU-LEU fuel cores and meets requirements in utilization as well.
Keywords: HEU, LEU, neutronics, thermal hydraulics, safety analyses
I. INTRODUCTION
kinetics parameters. Because the higher content
of 235U in a LEU FA compared to HEU FA, it
is needed to rearrange the fuel assemblies and
berrylium rods with the different way to the
first HEU core to meet the safety requirements.
In this full core conversion study,
neutronics, thermal hydraulics and safety
analysis were carried out to investigate
characteristics of LEU working core fully
loaded with LEU fuel. All computer codes
were validated with HEU and mixed cores.
Thermal hydraulics parameters at steady
state condition were obtained by using
PLTEMP3.8 code [11] introduced models and
correlations that suitable for the concentric
tube fuel type and natural convection regime of
the DNRR.
Using MCNP [6], REBUS-PC [5] and
VARI3D computer codes, a series of static
reactor physics calculation were performed to
obtain neutronics parameters of the working
core (see Fig. 1). Some parameters included in
the design of working core with shutdown
margin, excess reactivity taking into account of
irradiated Beryllium poisoning, control rod
worths, detailed power peaking factors,
neutron performance at the irradiation
positions, reactivity feedback coefficients, and
Based on the neutronics analysis
parameters of the LEU core, the postulated
transients and accidents selected for the DNRR
are analyzed. The RELAP5/MOD3.2 code [15]
was used for analysis of RIA (Reactivity
Initiated Accident), LOFA (Loss Of Flow
Accident) transients.
©2014 Vietnam Atomic Energy Society and Vietnam Atomic Energy Institute
LUONG BA VIEN, LE VINH VINH, HUYNH TON NGHIEM, NGUYEN KIEN CUONG
These study results showed that a LEU
core loaded with 92 fuel assemblies and 12
beryllium rods around the neutron trap satisfies
the safety requirements while maintaining the
utilization possibility similar to that of the
previous HEU and recent mixed fuel cores.
REBUS-MCNP Linkage [7] was used
to calculate burnup distribution using “two
way” linking option in which MCNP is used
for calculating neutron flux and cross section
in one group neutron energy and burn up
calculation is implemented by REBUS-PC.
The MCNP5 code using an ENDF-B/VI
cross section library was used to construct a
detailed geometrical model of each reactor
component and calculate control rod worths,
multiplication coefficient, power distribution,
neutron flux performance in irradiation
positions, reactivity feedback coefficients, and
kinetics parameters (prompt neutron life time
and delayed neutron fraction).
A detailed geometrical model of reactor
components including all fuel assemblies,
control rods, irradiation positions, beryllium
and graphite reflectors, horizontal beam tubes
and thermal column was made in the MCNP
model, except in the axial reflectors above and
below the fuel assembly where some materials
were homogenized. Fig. 2a provides the radial
and axial models of the reactor for Monte
Carlo Calculations.
Fig. 1. The new designed working core loaded with
92 LEU FA and 12 Beryllium rods.
II. CALCULATION MODELS AND
COMPUTER CODES
A. Neutronics and Thermal Hydraulics
Calculation
The diffusion code REBUS-PC with
finite difference flux solution method was used
to perform core calculation for reactor physics
characteristics and operation cycle calculations
with micro neutron cross sections according to
7 energy groups (collapsed from 69 energy
groups) that were generated by WIMS-ANL
The kinetics parameters were calculated
also by VARI3D code. The real and adjoint
fluxes which are required to compute these
parameters were provided by DIF3D-a main
module of REBUS-PC code.
In diffusion theory, the reactor was
modeled in hexagonal geometry with a
heterogeneous representation of the fuelled and
non-fueled portions (see Fig. 2a). Each
homogenized fuel assembly was modelled
using five equal volume axial depletion zones.
The beam tubes were modeled using a
homogenized mixture of air or concrete,
graphite and aluminum.
code [4].
The FA cross sections were
generated in a radial geometry with each fuel
element depleted based upon its unique neutron
spectrum in the WIMS-ANL model. The
REBUS-PC fuel depletion chains included
production of six Pu isotopes, Am-241, Np-
237, and lumped fission product. Isotopic
precursors of Xe-135 and Sm-149 were also
included in the depletion chains so that Xe and
Sm transients during periods of shutdown and
startup could be modelled.
The reactor models for diffusion and
Monte-Carlos computer codes were validated
by comparing with good agreement not only to
11
DESIGN ANALYSES FOR FULL CORE CONVERSION OF …
the fresh HEU configuration cores but also to
for PLTEMP code. A fuel assembly was
modelled as three concentric cylindrical tubes.
the HEU burnt cores. These models were then
applied for partial core conversion analyses of
DNRR [3]. The measured data collected during
the deployment of partial core conversion
project showed that the predicted calculation
results are quite acceptable [8,9].
Before using PLTEMP code to
calculate for DNRR with fully LEU fuel
assemblies, the code was validated by
comparing
analytical
results
with
experimental results of mixed-core.
The PLTEMP/ANL3.8 [15] thermal-
hydraulics code for plate and concentric-tube
geometries with capability of calculating
natural circulation flow was used for thermal-
hydraulics analyses. A chimney model as well
as Collier heat transfer correlation and CHF
Shah’s correlation have been recently
implemented make the code suitable DNRR
thermal-hydraulics calculation.
B. Transient/Accidents analyses
The DNRR has three barriers as other
research reactors that prevent or limit the
transport of fission products to the
environment, which are fuels and cladding,
reactor pool water and reactor confinement.
The safety system settings are showed in
Table I.
Fig. 2b shows the model of WWR-M2
fuel assembly, core and chimney of the DNRR
Table I. Safety system settings.
Parameters
Maximum thermal power (Pmax)
Minimum reactor period (Tmin)
Deficient level of pool water
Primary coolant flow rate
Safety system settings
550 kW (110% FP)
20s
60 cm
40 m3/h
Secondary coolant flow rate
70 m3/h
In the Safety Analysis Report (SAR) for
the DNRR [1], the possible initiating events
were classified by groups. The initiating events
in each group are then analyzed and justified in
order to identify the limiting event that will be
selected for further detail quantitative analysis.
The limiting event in each group has potential
consequences that exceed all others in that
group. Limiting events were selected for
detailed analyzed are as follows: (1)
Uncontrolled withdrawal of a control rod; (2)
Primary/Secondary
Pump
Failure;
(3)
Earthquake; (4) Fuel cladding failure. A
summary of the core parameters used for the
safety analysis is given in Table II.
Fig. 2a. Radial and Axial models for Monte
Carlo calculations (upper) and Radial model
for Diffusion Theory calculations (under).
12
LUONG BA VIEN, LE VINH VINH, HUYNH TON NGHIEM, NGUYEN KIEN CUONG
To ensure the fuel clad integrity in
operational condition and to protect the public
and the environment in case of accident, in the
SAR for the DNRR, the following acceptance
criteria were defined:
For anticipated
occurrences:
1)
-
operational
3)
(1) Minimum margin to departure from
nucleate boiling (DNB) shall be over 1.5;
(2) Maximum temperature of fuel cladding
shall not exceed 400oC;
2)
Fig. 2b. DNRR model for PLTEMP
(1-fuel assembly cross-section; 2-FA model
for PLTEMP; 3-reactor coolant system model).
(3) Fuel cladding integrity shall be
assured.
- For accident conditions:
core corresponding to a cooling channel with
maximum heat flux. The average channel
represents the rest of the cooling channels.
Each channel was modelled as three fuel
element plates and four coolant flow gaps. The
nodding diagram of the DNRR for
RELAP5/3.2 is presented in Fig. 2c.
(1) Core covering shall be maintained;
(2) Core shall not be remarkably
damaged;
(3) Release of fission products into the
environment shall not be remarkable.
The RELAP5 code was used for
analyzing the events of excess reactivity
insertion by uncontrolled withdrawal of a
control rod and earthquake. The piping of the
primary cooling system and pool volume were
divided into nodes with similar dynamic
characteristics. The reactor core was divided
into 2 channels with axial nodes. The hot
channel represents the hottest channel in the
The MACCS2 code [19] was used to
estimate the radiological impact of the
hypothetical accident on the surrounding
public. The core radiation inventories were
calculated by ORIGEN2 code [20] using
neutron cross-sections of the actinides obtained
from MCNP5 code.
Fig. 2c. Nodding diagram of DNRR for RELAP5/3.2.
13
DESIGN ANALYSES FOR FULL CORE CONVERSION OF …
Table II. Core parameters used for safety analysis.
Parameters
Values
500
Power, kW
Coolant inlet temperature, oC
Peaking factor (shim rods at 300 mm)
- Axial peaking factor
- Radial peaking factor
- Local peaking factor
Reactor kinetics
32
1.363
1.376
1.411
8.92510-5
7.55110-3
- Prompt neutron life, s
- Delayed neutron fraction (1$)
Temperature reactivity coefficients
- Moderator, %/K; (293-400oK)
- Fuel, %/oC;
- 1.26410-2
(293-400oK)
- 1.8610-3
- 1.9210-3
- 1.5610-3
(400-500oK)
(500-600oK)
- Void, %/% of void
(0-5%)
(5-10%)
(10-20%)
-0.2432
-0.2731
-0.3097
Reactivity control
- Shutdown worth, % (2 safety rods)
- Maximum withdrawal speed of one shim
rod, mm/s
3.7
3.4
and of the regulating rod, mm/s
Reactor protection characteristics
- Response time to overpower scram, s
- Response time to fast period scram, s
Start-up range
20
0.16
9.1
6.7
Working range
- Drop time of control rods, s
0.67
results in large negative reactivities which alter
flux and power distributions.
III. RESULTS AND DISCUSSIONS
A. Neutronics and Thermal Hydraulics
Program Beryl [10] has been modified
to calculate the 3He, 6Li and 3H concentrations.
The MCNP5 was then used to determine the
poisoning effect of 3He, 6Li and 3H
concentrations on reactor core reactivity. The
comparison of reactivities between calculation
results and measured data of some beryllium
blocks irradiated in DNRR (Table III) shows
that the negative reactivity of irradiated
beryllium determined by above-mentioned
3He and 6Li Poisoning of Irradiated
Beryllium [10]
Since 1984, the DNRR has been put into
operation with a considerable amount of
Beryllium used for neutron trap at the core
center and periphery for improving neutron
reflection around. Because Beryllium has large
thermal neutron absorption cross sections, the
3
6
3
buildup of He, Li and H concentrations
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LUONG BA VIEN, LE VINH VINH, HUYNH TON NGHIEM, NGUYEN KIEN CUONG
method is reliable. Six beryllium rods were
used for measurement, two fresh beryllium
rods and four irradiated beryllium rods (two
beryllium rods at the end 1994 and two at the
end 2002). 9-6 and 5-6 positions were chosen
to measure reactivity of couple beryllium rods
through changing position of control rod
(Regulating Rod). The error of control rod
position is estimated about 0.4 cent.
Following calculation scheme for beryllium
poisoning above, reactivity of the poisoning
process in new configuration cores about -1$.
All calculation for design LEU cores,
beryllium poisoning is included in the model
for MCNP code.
Table III. Comparison of calculated and measured of reactivities of irradiated beryllium rods in DNRR.
Measured reactivity
(Cent)
Calculated reactivity
(Cent)
Error
(%)
2 Beryllium Rods at
the end 1994
16.34
-3.89 0.4
-4.65 0.0038
2 Beryllium Rods at
the end 2002
12.66
-6.28 0.4
-7.19 0.0039
The working core characteristics
the Table IV. The shutdown margins of the
core is met the safety requirement of -1.0%.
Calculated neutron flux at the neutron trap of
the core is nearly the same as that of mixed
core (92HEU+12LEU). Table V shows the
control rod worths. Detailed neutron flux
performance at the main irradiation positions
are presented in Table VI.
From the calculation results of shutdown
margins, excess reactivities, power peaking
factors, and neutron performance at the
irradiation positions of 4 candidates cores, the
working core with the better features from the
safety and utilization point of view was chosen
for detailed analysis. The main calculated
characteristics of working core is showed in
Table IV. Calculation results of working core compared with current mixed core.
Current Mixed
Core
Parameters
LEU Core
Excess Reactivity (%) – Fresh
Excess Reactivity (%) – After 600FPDs
Shutdown Margin (%) – Fresh
Shutdown Margin (%) – After 600 FPDs
Radial Power Peaking Factor
Control Rods Out
6.63
3.79
-2.92
-6.62
-4.56
1.431
1.398
1.434
Control Rods In
Thermal Neutron Flux at Neutron Trap Center (n/cm2)
Control Rods Out
2.22E+13
2.14E+13
2.22E+13
3.15E+12
Control Rods In
Fast Neutron Flux at Neutron Trap Center (n/cm2)
Control Rods Out
1.95E+12
1.92E+12
Control Rods In
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DESIGN ANALYSES FOR FULL CORE CONVERSION OF …
Table V. Control Rods worths (%k/k).
Core1
Fresh
MCNP
error
Core1
Burnt
MCNP
error
Control Rods
Shim rod 1
2.5896
2.6100
2.7784
2.4687
0.4363
2.1955
2.2356
0.000091
0.000111
0.000118
0.000122
0.000126
0.000106
0.000119
2.3539
2.4033
2.5381
2.2604
0.3629
2.3084
2.3579
0.000091
0.000124
0.000122
0.000117
0.000119
0.000115
0.000105
Shim rod 2
Shim rod 3
Shim rod 4
Regulating rod
Safety rod 1
Safety rod 2
Table VI. Neutron flux performance.
Thermal, <0.625eV
(n/cm2.s)
Epithermal,
<0.821MeV
(n/cm2.s)
Fast, <10MeV
(n/cm2.s)
Irradiation positions
Fresh
Maximum 2.07E+13 2.20E+13 6.79E+12 7.12E+12 1.83E+12 1.92E+12
Average 1.45E+13 1.49E+13 6.00E+12 6.04E+12 1.62E+12 1.63E+12
Maximum 9.45E+12 9.86E+12 8.19E+12 8.42E+12 2.98E+12 3.02E+12
Average 7.00E+12 7.12E+12 6.53E+12 6.51E+12 2.46E+12 2.44E+12
Maximum 5.41E+12 5.66E+12 9.63E+12 9.76E+12 4.22E+12 4.26E+12
Average 4.11E+12 4.18E+12 7.23E+12 7.15E+12 3.19E+12 3.15E+12
Maximum 9.24E+12 9.71E+12 8.02E+12 8.22E+12 2.92E+12 2.99E+12
Burnt
Fresh
Burnt
Fresh
Burnt
Neutron
Trap
Channel
13-2
Channel
7-1
Channel
1-4
Average
6.85E+12 7.01E+12 6.41E+12 6.40E+12 2.42E+12 2.40E+12
Rotary
Specimen
Average
3.55E+12 3.56E+12 7.58E+11 7.56E+11 1.93E+11 1.93E+11
Power Distribution and Power Peaking
Factors
of control rods at 250 mm. Detailed axial
power distribution according to control rod
position was also calculated. Radial power
distributions at different control rod position
are showed in Fig. 3.
Power peaking factors of the core with
different position of control rods were
calculated and presented in Table VII. The
maximum power peaking factor is in position
Table VII. Power peaking factor according to control rod positions
Peaking Factor
F.A. Radial Core Radial Axial
Position
(mm)
Total
2.498
2.589
2.615
2.648
2.646
2.605
2.537
0
1.378
1.378
1.375
1.377
1.376
1.378
1.378
1.398
1.399
1.403
1.409
1.411
1.415
1.434
1.296
1.343
1.356
1.365
1.363
1.336
1.284
150
200
250
300
350
600
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LUONG BA VIEN, LE VINH VINH, HUYNH TON NGHIEM, NGUYEN KIEN CUONG
0.980
0.957
1.165
1.126
1.164
1.129
0.973
0.947
0.849 0.858 0.979 1.107 1.006 0.885
0.851 0.858 0.964 1.085 1.001 0.895
0.913
0.901
0.963
0.958
1.004
1.008
0.835
0.889
0.877 0.762
0.875 0.817
1.021
1.023
0.775 0.908
0.833 0.913
0.825
0.876
SR
1.082
1.124
1.406
1.368
ShR
ShR 1.090 1.410
0.790
0.841
0.787 0.921
0.843 0.915
0.917
0.906
1.126 1.370
0.962 0.872 0.906 1.281
0.939 0.873 0.953 1.296
1.259 0.882 0.842
1.283 0.937 0.850
0.919
0.903
1.138
1.090
0.983
0.975
0.959
0.933
0.998
0.974
1.313
1.289
1.353
1.321
0.865
0.855
0.930
0.929
1.139
1.122
0.911
0.900
RgR
1.056 1.220
1.027 1.192
1.307
1.284
1.167 1.018 0.996 1.364
1.117 0.991 0.987 1.332
0.918 0.843 0.911
0.919 0.837 0.887
0.868
0.919
1.252
1.273
0.845
0.835
1.020 0.903
0.992 0.904
0.929 1.296
0.975 1.312
0.808
0.818
0.996 0.830
0.977 0.881
1.079
1.114
ShR
ShR
1.421 1.408
1.381 1.368
1.106
1.145
0.755 0.841
0.808 0.836
0.968 0.810 0.860 1.038
0.959 0.863 0.910 1.038
1.005 0.816 0.745 0.831
1.006 0.866 0.803 0.836
SR
1.016 0.918 1.031 1.124 0.986 0.858 0.843 0.895
0.994 0.917 1.019 1.097 0.970 0.857 0.846 0.889
1.180 0.985
1.137 0.960
1.198
1.156
0.973
0.949
Fig. 3. Radial power distribution (Upper values: Fresh Core; Under values: Burnt Core)
Reactivity Feedback Coefficients and
Kinetics Parameters
kinetics parameters of the LEU cores
calculated using the VARI3D and MCNP5
codes. The calculated results from the two
computer codes are in good agreement.
These data will be used in transient
calculation for safety analysis of fully LEU
core of DNRR.
Reactivity
feedback
coefficients
calculated with the MCNP5 are depicted in
Table VIII. The negative results of reactivity
feedback coefficients show the inherent
safety of the LEU core. Table IX shows the
Table VIII. Feedback reactivity coefficients.
Parameter
DATA
±σ
Moderator Temperature Reactivity Coefficient (%/oC)
293 oK to 400 oK
-0.01317
0.00005
Fuel Temperature (Doppler) Reactivity Coefficient
(%/oC)
293 oK to 400 oK
400 oK to 500 oK
500 oK to 600 oK
-0.00192
-0.00182
-0.00154
0.00005
0.00003
0.00002
Moderator Density (Void) Reactivity Coefficient (%/%
of void)
0 to 5 %
-0.2514
-0.2784
-0.3255
0.0011
0.0012
0.0006
5% to 10 %
10 % to 20 %
17
DESIGN ANALYSES FOR FULL CORE CONVERSION OF …
Table IX. Calculated results of kinetics parameters for LEU core.
Decay Const.
Relative Yield
ai
Fraction
λi (s-1)
βi
Family, i
1
2
3
4
5
6
1.334E-02
3.273E-02
1.208E-01
3.030E-01
8.503E-01
2.856E+00
3.507E-02
1.804E-01
1.742E-01
3.843E-01
1.594E-01
6.666E-02
2.648E-04
1.363E-03
1.315E-03
2.902E-03
1.204E-03
5.033E-04
Total delayed neutron fraction, β
VARI3D
7.551E-03
7.761E-03
7.762E-03
MCNP5 – Fresh
MCNP5 – Burnt
Prompt neutron life time, ℓ
8.925E-05
Burn up calculation
extended about 11 years (calculated with 1300
hours per year) or 600 full power days (FPDs).
The burn up of U-235 reached average value of
8.2% and maximum value of 11.4%. In the
next cycle, about 8 fuel assemblies will be
inserted so the reactor core will operate with
100 fuel assemblies. The Fig. 4 shows burn up
distribution after 600 FPD operation.
The first cycle length was estimated by
REBUS-MCNP Linkage system code. Burn up
calculations were performed by assuming that
shim rods and regulating rod were in critical
position following each burn-up step. The
value of reactivity for Xe-135 poisoning was
estimated about 1.2% k/k. The result of
depletion shows that operating time may be
Fig. 4. Burn up distribution using REBUS-MCNP Linkage system after 600 FPD.
18
LUONG BA VIEN, LE VINH VINH, HUYNH TON NGHIEM, NGUYEN KIEN CUONG
working core meets the requirements of
The PLTEMP code was used for
thermal hydraulics safety. At the power of
calculating cladding temperature, coolant
500kW with systematic errors, maximum
temperature and safety margins for the
cladding temperatures are below the
candidate cores. The calculated results are
permissible value of 103oC [2] and far below
presented in Table X and Fig. 5. At nominal
the ONB temperature (estimated about 116oC
power without uncertainties and maximum
using Forster-Greif correlation). The maximum
permissible inlet temperature (32oC), the
outlet coolant temperature is calculated about
maximum cladding temperature is 90.50oC.
60oC, much lower than saturated temperature
(108oC).
Calculation was carried out for nominal power
with systematic errors (equivalent to 70kW
power) and the maximum cladding temperature
is 95.69oC. In this case, by using Shah’s
correlation, the obtained minimum DNBR is
9.9. The minimum flow instability power ratio
(MFIPR) is 2.04. From above-mentioned
calculated results, it may conclude that the
Fig. 6 shows the comparison of cladding
temperature of 92FA LEU cores and 89FA
fresh HEU core. Compared to the 89FA fresh
HEU core established in 1984, cladding
temperature of working core is about 2oC
lower.
Table 10. Cladding temperature and ONB margin by PLTEMP Code.
500kW
550kW
with sys. error
T-
600kW
with sys. error
T-
Distance
(cm)
without sys. error
with sys. error
T-
T-
Tc(oC)
63.91
70.56
78.46
84.83
88.77
90.50
89.86
87.10
83.98
79.67
74.91
71.21
ONB(oC)
51.89
45.59
38.07
31.90
27.95
26.05
26.34
28.58
31.14
34.76
38.73
41.70
Tc(oC)
ONB(oC)
Tc(oC)
ONB(oC)
Tc(oC)
ONB(oC)
2.5
7.5
66.89
74.13
82.71
89.61
93.85
95.69
94.95
91.91
88.24
82.92
77.42
73.32
49.24
42.36
34.18
27.50
23.26
21.25
21.63
24.13
27.24
31.91
36.64
40.02
68.95
76.58
85.63
92.89
97.33
99.23
98.43
94.41
89.94
84.43
78.79
74.64
47.39
40.14
31.51
24.48
20.06
17.97
18.41
21.94
25.90
30.74
35.57
38.98
70.96
78.97
88.46
96.05
100.68
102.65
100.76
96.22
91.57
85.89
80.13
75.94
45.59
37.97
28.91
21.57
16.95
14.80
16.40
20.48
24.60
29.59
34.52
37.93
12.5
17.5
22.5
27.5
32.5
37.5
42.5
47.5
52.5
57.5
100.0
90.0
80.0
70.0
60.0
50.0
40.0
30.0
20.0
10.0
100
92 LEU FA Core2
95
90
85
80
75
70
65
60
89 HEU FA Core
92 LEU FA Core1
T-clad
T-coolant
DT-ONB
30
0.0
0
10
20
40
50
60
0
10
20
30
40
50
60
Distance from core bottom (cm)
Distance from core bottom (cm)
Fig. 6. Comparison of calculated cladding temperature
between 92FA LEU cores and HEU core.
Fig. 5. T/H parameters at 500kW without
uncertainties.
19
DESIGN ANALYSES FOR FULL CORE CONVERSION OF …
The event of one shim rod inadvertently
2. Transient/Accidents analyses
withdrawal with speed of 3.4 mm/s from stable
operation of 100%FP (500 kW) are showed in
Fig. 7 and Table XI. In this case, the reactor
power increases and reaches to the over-power
setting value (110%FP) within 3.39 seconds
generating a scram signal. After a delay time of
0.16 seconds the reactor power is rapidly
suppressed because of the control rods
insertion. The peak power of the reactor is only
attained 0.553 MW with a slight increase of the
maximum fuel cladding temperature. With the
assumption of no overpower scram signal
appeared, a fast period scram signal is
generated after 8.33 seconds from the initiation
of transient event. The reactor will be
shutdown after 6.7 second delay with a peak
power of 0.957 MW. The maximum fuel
cladding temperature is predicted to be 113.0oC
without any nucleate boiling occurrences. The
minimum DNBR (Departure from Nucleate
Boiling Ratio) estimated about 6.5 is much
higher than the acceptance criterion of 1.5.
Uncontrolled withdrawal of one shim
rod or the regulating rod
In this event, it is assumed that one of
the shim rods or the regulating rod is
withdrawn in the most effective part from 200
mm to 400 mm at the speed for 3.4 mm/s of
shim rod and 20 mm/s for regulating rod. The
initial conditions are as follows:
a) Start-up case:
(1) -1% k/k sub-critical; Power level: 10-
5%FP; Coolant inlet temperature: 32oC.
(2) Critical state; Power level: 10-3%FP;
Coolant inlet temperature: 32oC.
b) Steady-state operation:
Power level: 100%FP; Coolant inlet
temperature: 32oC.
In sub-critical status, when one shim rod
is inadvertently withdrawn with the speed of
3.4 mm/s, from the core, the reactor power
only increases to the maximum value of
2.7810-7 MW while the fuel cladding
temperature is unchanged. With initial
conditions of criticality at the power level of
10-3%FP (510-6 MW) if there is no fast period
signal and the overpower trip setting is
110%FP, the fuel clad temperature reaches to
97.8oC, but still far below ONB (Onset of a
Nucleate Boiling) temperature.
With the same initial conditions, the
calculated results for the event of withdrawal
of the regulating rod are slightly different from
those of above-mentioned event, when one
shim rod is withdrawn. This can be explained
by the similar insertion rate of reactivity in the
two cases (about 0.02$/s). The regulating rod
has lower reactivity worth but higher
withdrawal velocity compared to those of a
shim rod.
Table XI. Transient results of one shim rod withdrawal from 100%FP.
Values
110%FP Scram
Parameters
Period Scram
15.1
Time to Peak Power, s
Peak Power, MW
3.6
0.553
3.7
0.957
15.2
Time to Peak Clad Temperature, s
Peak Clad Temperature, oC
Minimum DNBR
91.9
113.0
6.5
20
LUONG BA VIEN, LE VINH VINH, HUYNH TON NGHIEM, NGUYEN KIEN CUONG
120
110
100
90
1.0
0.9
0.8
0.7
0.6
0.5
0.4
0.3
0.2
0.1
0.0
20
18
16
14
12
10
8
Period scram
Period scram
80
DNBR
70
6
60
4
Overpower scram with set point 110% FP
50
Overpower scram with set point 110% FP
2
40
0
0
2
4
6
8
10
12
14
16
18
Time (s)
20
0
2
4
6
8
10
12
14
16
18
Time (s)
20
Fig. 7. Reactor power and cladding temperature transient of one shim rod withdrawal
from a stable operation of 100%FP.
Cooling pump failure
measures
undertaken
in
design
and
construction, the removal of all control rods
would not exceed 10 mm and insert a step
positive reactivity estimated of 0.3$. With this
reactivity insertion, the scram set-point of
reactor overpower is attained almost
instantaneously. If the reactor scram is
initiated by overpower signal with a delay of
0.16 sec, the fuel surface temperature
increases slightly before decreases with the
power, the residual heat after shutdown is
sufficiently removed from the fuel by natural
convection of pool water without considerable
increase of the temperature.
In the event of in-service primary or
secondary cooling pumps stopped working, the
reactor is automatically shutdown by an
abnormal technological signal on low flow rate
(the setpoint is 40 m3/h for the primary flow,
and 70 m3/h for the secondary flow). The
residual heat after shutdown is about 6% FP
(30 kW) in maximum and the natural
convection process can itself assure the good
cooling of the core.
If the reactor is purposely maintained at
full power operation, failure of cooling pumps
leads to loss of heat removal from the pool
water, and thus gradually increases of the pool
water temperature. The results show that the
clad temperature reaches the maximum
Fig. 8 shows the analyzing results of
the earthquake event assuming the
protection system fails to shutdown the
reactor, and Because of the loss of offsite
power due to the earthquake, the primary and
secondary pumps stop operating. In this case,
the reactor power increases to the max value of
1.525 MW after 200 seconds from the
initiation of this event. The reactor power then
rapidly decreases because the significant
increasing of core water temperature so that the
positive reactivity insertion is overtaken by the
negative reactivity feedback (about -0.44$).
The reactor is then kept at subcritical state. The
cladding temperature reaches a maximum
value of 118.2oC, then decreases with no
o
allowable operating clad temperature of 103 C
at about 55 min; i.e. the reactor could continue
its operation for 55 minutes within the envelope
of the limiting conditions of operation. The
results also show that even at the end of the
simulation (7000 s) the clad temperature has
been well below that of the acceptance criterion
for anticipated operational occurrences.
Earthquake
The postulated event of an earthquake of
intensity grade VI is assumed to occur while
the reactor is at full power. Owing to the
21
DESIGN ANALYSES FOR FULL CORE CONVERSION OF …
significant overheating of the fuel. The
value of about 1.12 MW. The cladding
temperature reaches to a maximum value of
118.38oC then gradually decreases to a stable
value of 115oC without nucleate boiling. The
maximum temperature of outlet water is 89oC
at the peak power then decreases and
stabilizes at about 82oC, well below the
saturation point. The minimum DNBR in this
case estimated about 4.74 is still far from the
acceptance criterion.
maximum outlet water reaches 89oC and
gradually decreases to a value at about 60oC,
which is still far below the saturation
temperature. The minimum DNBR of 4.79 is
much higher than the acceptance value.
In case the cooling pumps remain
working after the earthquake event (very
unlikely); the peak power reaches 1.57 MW
within 300 seconds and decreases due to
negative temperature feedback to a stable
1.6
120
16
Max. Cladding Temperature
110
14
1.4
100
12
DNBR
1.2
90
1.0
0.8
0.6
0.4
0.2
10
80
8
70
Max. Water Temperature at Outlet
6
4
2
0
60
50
40
30
Water Temperature at Inlet
Power
3000
0.0
0
500
1000
1500
2000
2500
3000
3500
4000
4500
5000
0
500
1000
1500
2000
2500
3500
4000
4500
5000
Time (s)
Time (s)
Fig. 9. Power and Temperature responses to earthquake event while cooling pumps are stopped functioning.
Fuel cladding failure (MHA)
95% efficiency is not available, and there are
no decay and deposition of radionuclides
within the reactor building.
For the derivation of source term of this
event, it is assumed that no core melting occurs
but cladding rupture of one fuel assembly is
involved. It is also assumed that the damaged
fuel assembly is irradiated at the maximum
neutron flux position in the core and the fuel
damage occurs immediately at the end of
operating cycle of 100 hrs with no decay.
The evaluation of dose to a member of
the public is calculated by code MACCS2
version
1.13.1,
using
the
following
assumptions: (1) The radionuclides are
released to the environment through the 40 m
stack; (2) The Gaussian plume model is used to
calculate air concentration of radioactivity; (3)
Tadmor and Gur parameterization is used for
this analysis; (4) No building in the vicinity (an
open area release), plume rise mechanics only
due to momentum rise (non-buoyant plume)
and no wet deposition are assumed; (5) The dry
deposition velocity is assumed to be 0.01 m/s,
which corresponds to a particle with an
aerodynamic equivalent diameter of 2 m to 4
m (for unfiltered particulate releases) [15]; (6)
From the damaged fuel assembly, 100%
of noble gases (Xe, Kr), 25% halogens (I), and
1% of other radionuclides (Cs, Te) [21] are
released directly to the reactor building with
the assumption of no retention of volatile
fission products in the pool water. During the
accident evolution, the emergency ventilation
system is not in place, the normal ventilation
system V1 is in operation but HEPA filter with
22
LUONG BA VIEN, LE VINH VINH, HUYNH TON NGHIEM, NGUYEN KIEN CUONG
Surface roughness length is specified as 50 cm;
(7) Mixing layer height is assumed to be 500 m
(see Table 36 in Appendix VII of Ref. 21); (8)
The breathing rate is 3.3x10-4 m3/s; (9) No
shielding and sheltering are assumed; (10)
Doses at each downwind distance are
calculated for one year after the arrival of the
plume (11). The environmental release is
assumed to begin at the start of the weather
conditions: Pasquill class D2.0 (most frequent
stability class and most frequent wind speed).
The effective equivalent doses, including
cloudshine dose, inhalation dose and
groundshine dose, as a function of the distance
from the source are shown in Table XII and
Fig. 10. It is seen that radiation exposure to the
general public with the maximum effective
dose of 0.64 mSv/year at distance from 400 m
to 500 m from the stack. This value is lower
than the annual dose limit of 1.0 mSv specified
for the public [22].
Table XII. The annual effective dose to the public vs distance for the MHA.
Distance
(m)
Effective Dose
(mSv)
Distance
(m)
Effective Dose
(mSv)
50
4.80E-02
1.43E-01
4.95E-01
6.42E-01
6.44E-01
5.94E-01
5.33E-01
4.74E-01
4.21E-01
3.75E-01
1100
1300
1500
1700
1900
2250
2750
3250
3750
4250
3.18E-01
2.59E-01
2.16E-01
1.83E-01
1.57E-01
1.23E-01
9.14E-02
7.08E-02
5.66E-02
4.64E-02
150
250
350
450
550
650
750
850
950
7.00E-01
6.00E-01
5.00E-01
4.00E-01
3.00E-01
2.00E-01
1.00E-01
0.00E+00
0
500
1000
1500
2000
2500
3000
3500
4000
4500
5000
Downwind Distance, m
Fig. 10. The annual effective dose to the public in MHA event within 5 km.
23
DESIGN ANALYSES FOR FULL CORE CONVERSION OF …
- If one of the cooling pumps stopped
IV. CONCLUSIONS
Neutronics, steady-state
working, the reactor is automatically shutdown
by a scram signal on low flow rate. The decay
heat is removed from the fuel by natural
convection of pool water. In this event, if the
reactor was purposely maintained at full power,
it could be safely operated for 55 minutes when
maximum cladding temperature is still lower
than the permissible value of 103oC.
thermal-
hydraulic and transient/accidents analyses for
Dalat Nuclear Research Reactor show that with
a slight change in arrangement of Be rods, the
main features of 92 LEU WWR-M2 FA cores
are equivalent to those of HEU and current
mixed fuel cores.
- The postulated earthquake event of
MSK intensity grade VI would cause a step
reactivity insertion of 0.3$. Even if the reactor
fails to be scrammed, this positive reactivity
can be covered by negative temperature
feedback if the cooling pumps are stopped
simultaneously, keeping the reactor sub-critical.
In case the cooling pumps continue operating
after earthquake event, the negative temperature
feedbacks act to bring the reactor power to a
stable level of about 1.12 MW without nucleate
boiling. The minimum DNBR is much higher
than the acceptance criterion of 1.5.
The negative values of reactivity feedback
coefficients show the inherent safety feature and
shutdown margin of both candidate cores meets
the safety required value of -1% k/k. The
working core with 92 fresh LEU fuel assemblies
can be operated for 600FPDs or about 11 years
based on the current operating schedule without
shuffling. The neutron fluxes at the irradiation
positions are not much different from those of the
current mixed fuel core.
In thermal hydraulics aspect, the
requirement of thermal-hydraulic safety margin
for two candidate cores in normal operational
condition is satisfied. The calculated maximum
cladding temperature in operational condition
is below the permissible value of 103oC.
- The maximum hypothetical accident
assumes 100% of noble gases (Xe, Kr), 25%
halogens (I), and 1% of other radio-nuclides (Cs,
Te) in a most power fuel assembly after a long
run are released into the environment through
40m high stack. This event is considered to be
very unlikely to occur for the DNRR. Even so, it
would not cause undue radiological risk to the
environment or the public.
In transient/accidents aspect, some
postulated initiating events and accident related
to the conversion of the DNRR to full LEU
core were selected and analyzed. Based on the
calculated results, conclusions might be
withdrawn as following:
ACKNOWLEADGMENTS
- The excess reactivity insertions when
inadvertent withdrawals of control rod from
start-up or nominal power operation are
prevented by safety settings to initiate the
reactor scram at overpower and fast period.
None of these initiators would lead to the ONB
and DNB, ensuring the integrity of the fuel
cladding. The residual heat after shutdown is
sufficiently removed from the fuel by natural
convection of pool water.
The authors would like to express their
gratitude to experts from the Reduced
Enrichment for Research and Test Reactors
(RERTR) program of Argonne National
Laboratory for financial support as well as very
useful discussions during design calculation of
full core conversion for the Dalat Nuclear
Research Reactor.
24
LUONG BA VIEN, LE VINH VINH, HUYNH TON NGHIEM, NGUYEN KIEN CUONG
[13] Technical Design for Reconstruction and
Enlargement of the Dalat Nuclear Research
Reactor - Volume 3. State Design Institute,
USSR State Committee for the Utilization of
Atomic Energy, Moscow, (in Russian) (1979).
REFERENCES
[1] Safety Analysis Report for the Dalat Nuclear
Research Reactor, (2003).
[2] VVR-M2 and VVR-M5 Fuel Assemblies-
Operation Manual, 0001.04.00.000 PЭ, (2006).
[14] Additional Physics and Thermal-Hydraulic
Data for Reactor IVV-9, State Design Institute,
USSR State Committee for the Utilization of
Atomic Energy, Moscow, (in Russian) (1980).
[3] Report on the 7th Core Loading for the Dalat
Nuclear Research Reactor, NRI, Dalat, (in
Vietnamese), (2009).
[15] RELAP5/MOD3 Code Manual, SCIENTECH,
Inc. Rockville, Maryland, (1999).
[4] J.R. Deen, W.L. Woodruff, C.I. Costescu, and
L. S. Leopando, “WIMS-ANL User Manual
Rev. 5”, ANL/RERTR/TM-99-07, Argonne
National Laboratory, (February 2003).
[16] W. L. Woodruff, et al., A comparison of the
PARET/ANL and RELAP5/MOD3 Codes for
the Analysis of IAEA Benchmark Transients
and the SPERT Experiments, RERTR
Program, ANL.
[5] P. Olson, “A Users Guide for the REBUS-PC
Code, Version 1.4,” ANL/RERTR/TM02-32,
(December 21, 2001).
[17] V. V. Le and T. N. Huynh, Application of
RELAP5/MOD3.2 for the DNRR, Proceedings
of JAEA Conf. 2006-001.
[6] J. F. Briesmeister, Ed., “MCNP – A General
Monte Carlo N-Particle Transport Code,
Version 4C”, LA-13709-M (April 2000).
[18] M. M. Shah, Improved General Correlation for
Critical Heat Flux during Upflow in Uniformly
Heated Vertical Tubes, International Journal of
Heat and Fluid Flow, Vol. 8, No. 4, pp. 326-
335 (1987).
[7] John G. Stevens, “The REBUS-MCNP Linkage”,
Argonne National Laboratory, (2007).
[8] N.A. Hanan, J.R. Deen, J.E. Matos, “Analyses
for Inserting Fresh LEU Fuel Assemblies
Instead of Fresh Fuel Assemblies in the DNRR
in Vietnam, 2004 International Meeting on
RERTR, Vienna, (2004).
[19] MACCS2 Computer Code Application
Guidance for Documented Safety Analysis,
U.S. Department of Energy, (June 2004).
[9] V. V. Le, T. N. Huynh, B. V. Luong, V. L.
Pham, J. R. Liaw, J. Matos, “Comparative
Analyses for Loading LEU Instead of HEU
Fuel Assemblies in the DNRR”, RERTR Int’l
Meeting, Boston, (2005).
[20] A.G. Croff, A User Manual for the ORIGEN2
Computer Code, Oak Ridge National
Laboratory, (1980).
[21] INTERNATIONAL
ATOMIC
ENEGY
AGENCY, Derivation of the Source Term and
Analysis of Radiological Consequences for
Research Reactor Accidents, SAFETY
REPORTS SERIES No. 53, VIENNA (2008).
[22] Governmental Decree for the Implementation
of the Ordinance on Radiation Protection and
Control, Government of the Socialist Republic
of Vietnam, No. 50/1998/ND-CP, Hanoi, (in
Vietnamese) (1998).
[10] Teresa Kulikowska et al., Raport IAE-40/A,
(1999).
[11] Arne P. Olson, M. Kalimullah, “A users guide
to
the
PLTEMP/ANL
V3.8
Code”,
ANL/RERTR, Argonne National Laboratory,
(June, 2009).
[12] Le Vinh Vinh, Huynh Ton Nghiem and
Nguyen Kien Cuong, “Preliminary results of
full core conversion from HEU to LEU fuel of
the Dalat Nuclear Research Reactor”, RERTR
Int’l Meeting, Beijing, (2009).
25
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