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Thermal hydraulic and safety analysis for core conversion (HEU–LEU) of Syrian Miniature Neutron Source Reactor
Progress in Nuclear Energy 60 (2012) 140e145
Contents lists available at SciVerse ScienceDirect
Progress in Nuclear Energy journal homepage: www.elsevier.com/locate/pnucene
Thermal hydraulic and safety analysis for core conversion (HEUeLEU) of Syrian Miniature Neutron Source Reactor H. Omar*, N. Ghazi, A. Hainoun Nuclear Engineering Department, Atomic Energy Commission, P.O. Box 6091, Damascus, Syria
a r t i c l e i n f o
a b s t r a c t
Article history: Received 17 November 2011 Received in revised form 13 May 2012 Accepted 26 May 2012
The paper presents the behavior and properties analysis of the low enriched uranium fuel compared with the original high enriched uranium fuel. The MNSR reactor core was modeled with both fuel materials and the reactor behavior was studied during the steady state and abnormal conditions. The MERSAT code was used in the analysis. The steady state thermal hydraulic analysis results were compared with that obtained from the experimental results hold during commissioning the Syrian MNSR. Comparison with experimental data shows that the steady-state behavior of the HEU core was accurately predicted by the MERSAT code calculations. The validated model was then used to analyze LEU cores with two proposed UO2 fuel pin designs. With each LEU core, the steady state and 3.77 mk rod withdrawal transient were run and the results were compared with the available published data in the literatures for the low enriched uranium fuel core. The results reveal that the low enriched uranium fuel showed a good behavior and the peak clad temperatures remain well below the clad melting temperature during reactivity insertion accident. Ó 2012 Elsevier Ltd. All rights reserved.
Keywords: MNSR Safety analysis Core conversion MERSAT
1. Introduction HEU fuel is more economical, the fuel could be used longer in the reactor core, and it has a higher specific reactivity. In the 1970s, however, many people again became concerned about the possibility that some fuels and fuel cycles could provide an easy route to the acquisition of nuclear weapons. Since enrichment to less than 20% is internationally recognized as a fully adequate barrier to weapons usability, certain member states have moved to minimize the international trade in highly enriched uranium and have established programs to develop the technical means to help convert reactors to the use of lowenrichment fuels with minimum penalties. This could involve modifications in the design of the reactor and development of new fuels. As a result of these programs, it is expected that most reactors can be converted to the use of low enriched fuel. The use of new fuel elements in a research reactor would require that the reactor’s current safety analysis report be revised to assess the new balance of safety factors. Principal issues will include the effect of changes in enrichment and fuel technology on temperature and void coefficients of reactivity, thermal
* Corresponding author. Tel.: þ963 11 2132580; fax: þ963 6111926. E-mail addresses: pscientifi[email protected], scientifi[email protected] (H. Omar). 0149-1970/$ e see front matter Ó 2012 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.pnucene.2012.05.007
hydraulic safety criteria, fission product retention, and control system effectiveness. The Syrian Miniature Neutron Source Reactor (MNSR) is a 30 kW, highly enriched uranium, light-water moderated and cooled, beryllium reflected, tank in pool type research reactor that has been in operation since Mars 1996. It is a small, safe nuclear facility which employs high enriched uranium as fuel. MNSR is designed, manufactured and constructed by China Institute of Atomic Energy (CIAE), Beijing, China. The primary use of the reactor is to provide a neutron source with maximum flux of 1 1012 n/ cm2 s with 10 irradiation tubes; 5 inside and 5 outside the beryllium reflector. The core consists of fuel elements which form a fuel cage. The cage is inside an annular beryllium reflector and rests on a lower beryllium reflector plate. The volume of the vessel is 1.5 m3. The fuel elements are all highly enriched uraniumealuminum alloy extraction clad with aluminum. They are arranged in ten multiconcentric circle layers at a pitch distance of 10.95 mm. The element cage consists of two grid plates, four tie rods and a guide tube for the control rod. The two grid plates and four tie rods are connected by screws. The total number of lattice positions is 350 and the number of fuel elements is 347. The remaining positions are filled with dummy aluminum elements. The core region of the Syrian MNSR is located under 4.7 m of water close to the bottom of a watertight reactor vessel (Figs. 1 and 2).
H. Omar et al. / Progress in Nuclear Energy 60 (2012) 140e145
141
Fig. 1. Schematic diagram for MNSR according to MERSAT model; 1 e annular reflector, 2 e lower reflector, 3 e inlet, 4 e in-gap, 5 e ring, 6 e hole, 7 e out-gap, 8 e outlet, 9 e reactor vessel, 10 e reactor pool, 11 e inner irradiation site, 12 e control rode, 13 e outer irradiation site, 14 e shim tray, 15 e reactor base. Fig. 2. MERSAT nodalization.
The beryllium annulus and lower reflector are spaced to form the lower orifice which controls water flow through the core. The top plate of the core and annulus are spaced to form the upper orifice. There are no pumps or heat exchangers in the system. All coolant flow is due to natural circulation. Heat removal from the tank is by conduction through the tank wall to the pool water. Heat removal from the pool is by evaporation and conduction to the air and by conduction through the wall of the pool. An aluminum tray holds the upper reflector which contains semicircular beryllium shims which are added approximately once a year to compensate for fuel burn-up and samarium poisoning. In the current study, the thermal hydraulic analysis for core conversion of Syrian MNSR using the MERSAT code was conducted as a part of the on-going worldwide effort of converting from highly enriched in uranium (HEU) reactor cores into low enriched in uranium (LEU). Two proposed low-enriched uranium dioxide (UO2) cores were considered as shown in Table 1. 2. Computer code employed (MERSAT code) The code MERSAT (Model for Evaluation of Reactor Safety and Analysis of Thermal-Hydraulics) is one-dimensional, two-phase fluid dynamic. The code is under development and verification at the AECS (Hainoun and Ghazi, 2012). The mechanical nonequilibrium of the two phases is described in the momentum equation using the drift-flux model of Zuber and Findlay. The thermodynamic non-equilibrium of the two phases in the subcooled boiling regime is similar to that implemented in the code
ATHLET (Hainoun et al., 1996). The heat conduction and heat transfer module allows a flexible simulation of fuel rods and structures. The nuclear heat generation is calculated by a pointkinetics approach similar to that in PARET (Obenchain, 1973). The heat transfer package describing the wide range of two phase flow
Table 1 Design parameters of the MNSR.
Fuel U-235 enrichment (%) Cladding material Coolant Moderator Fuel rod outer diameter (mm) Fuel meat outer diameter (mm) Clad thickness (mm) Height of core (mm) Total number of fuel rods U-235 content (g) in the core Reflector
HEU (existing core)
LEU (proposed core)
LEU (proposed core)
UAl4-Al 89.8
UO2 12.45
UO2 12
Aluminum H2O H2O 5.5
Zircoloy-4 H20 H20 5.5
Zircoloy-4 H20 H20 5.1
4.3
4.3
4.2
0.6 230 347
0.6 230 347
0.45 230 347
1009
1353
1264
Metallic beryllium
Metallic beryllium
Metallic beryllium
142
H. Omar et al. / Progress in Nuclear Energy 60 (2012) 140e145
22.5
Inlet temperature Tin , C
Input Temperature, C
22.0
Experiment RELAP MERSAT
36
Calculated RELAP Calculated MERSAT Experiment
21.5
21.0
20.5
32
28
24
20
0
20.0
2000
4000
6000
8000
10000
Time, s 0
100
200
300
400
500 Fig. 5. Comparison of experimental and calculated inlet coolant temperature during operation at nominal power of 30 kw.
Time, s Fig. 3. Comparison of experimental and simulated data for coolant inlet temperature during start-up till reaching the power 30 kw.
25
20
Tout-Tin , c
has been developed by consolidating and restructuring some modules in the codes COBRA-3CeRERTR (Woodruff, 1983), PARET (Obenchain, 1973), MINCS (Watanabe et al., 1992; Richards, 1988) and SIKADE (Meister, 1989; Langenbuch, 1979; Hainoun et al., 1996). The modeling of special hydrodynamic components, like fill, leak, time-dependent volume, is relying on the procedure applied in the code ATHLET (Lerchel and Austregesilo, 1996), and QUABOX-HYCA (Langenbuch, 1979).
15
10
3. The MERSAT model
Experiment RELAP MERSAT
5
The input model for the MERSAT code relies on the original design data of MNSR. Some simplifications and adjustments were applied to achieve the best representation of MNSR components according to the specific features and the one-dimensional approach in MERSAT code. The MNSR reactor was divided into primary and secondary systems. The primary system, hydrodynamics systems, representing the reactor vessel including reactor
0 0
2000
4000
6000
8000
10000
Time, sec Fig. 6. Comparison of experimental and calculated difference inlet and outlet coolant temperature during operation at nominal power of 30 kw.
40
1.0
35
0.8
30
25
Experiment Calculated RELAP Calculated MERSAT
20 0
100
200
300
400
500
Time, s Fig. 4. Comparison of experimental and simulated data for coolant outlet temperature during start-up till reaching the power 30 kw.
Relative Power
Output Temperatute, C
45
0.6
0.4
0.2
0.0 0.00
0.05
0.10
0.15
Distance from bottom, m Fig. 7. Axial power density distribution.
0.20
H. Omar et al. / Progress in Nuclear Energy 60 (2012) 140e145 Table 2 Results of steady-state thermal hydraulic analysis of MNSR.
45
a
Parameter
Temperature, C
40
30
25
HEU LEU-12% LEU-12.45%
20
0.05
0.10
0.15
0.20
Distance from bottom, m
Temperature, C
80
b
70
60
HEU LEU-12% LEU-12.45%
50
40 0.00
0.05
0.10
0.15
HEU LEU LEU (existing core) (proposed core) (proposed core)
Fuel pin diameter (mm) 5.5 U-235 enrichment (%) 89.8 Operating power (kW) 30 Pressure at core inlet (kPa) 1.5 111.8 Saturation temperature in core ( C) Peak clad surface 65 temperature ( C) Peak centerline fuel 69.5 temperature ( C)
35
15 0.00
143
0.20
Distance from bottom, m
5.5 12.45 30 1.5 111.8
5.1 12 30 1.5 111.8
67
65.8
78.8
77.5
so that the division took into account all the components of the reactor and if there is any interchange between the components the model should be able to simulate this interchange. Therefore, the sections: V-UP, P-UH, V-DN, P-DC, P-BW, TW-S, TW-M and CCWMS are parts of the reactor vessel immersed in water. The objects G1 to G10 represent coolant channels inside the core, whereas the fuel elements structure is not illustrated in the figure due to their small dimensions. Fig. 2 shows MERSAT nodalization of MNSR. The annular beryllium, bottom beryllium and reactor vessel considered as the heat structure outside the core. Coolant inlet temperature is 20 C and the inlet pressure is 1.5 bar. Values of feedback reactivity coefficients, prompt neutron lifetime and effective delayed neutron fraction in addition to the material properties for the HEU and for LEU fuel were taken from literatures (Dunn et al., 2007; Bokhari and Showket, 2010; Omar et al., 2010).
4. Results and discussion 4.1. Steady state operation
75
c
Computed coolant inlet and outlet temperature during start-up till reaching the power 30 kW were shown in Figs. 3 and 4 respectively. The results agree reasonably well with measurements and the early reported data using the RELAP5 code (Omar et al., 2010). The error in measuring the inlet and outlet temperatures was 5% (SAR, 1993). As shown in Fig. 3 no change in the coolant inlet temperature is observed during the first 300 s, which
70
60
55
80
50
HEU LEU-12% LEU-12.45%
45 40 0.00
0.05
0.10
0.15
0.20
Distance from bottom, m Fig. 8. (a): Axial distribution of coolant temperature in HEU and in the two proposed LEU cores at steady-state power. (b): Axial distribution of fuel centerline temperature in HEU and in the two proposed LEU cores at steady-state power level. (c): Axial distribution of clad temperature in HEU and in the two proposed LEU cores at steadystate power level.
core and reactor primary components. The secondary system representing the reactor pool. Both primaries and secondary system are thermally interconnected through the wall of reactor vessel. The main components of primary and secondary loop are represented in Fig. 1. The reactor was divided into a number of sections
Reactor Power, kW
Temperature, C
65
HEU LEU-12.45% LEU-12%
60
40
20
0 0
1000
2000
3000
Time, sec Fig. 9. Power as a function of time for a 3.77 mk reactivity insertion with HEU and LEU Fuel using the MERSAT code.
144
H. Omar et al. / Progress in Nuclear Energy 60 (2012) 140e145 60
140
HEU 12.45% 12%
120
50
Temperature, C
Temperature, C
100
80
60
40
30
40
HEU LEU-12% LEU-12.45%
20
20 0
500
1000
1500
2000
2500
3000
3500
0
500
1000
1500
2000
2500
3000
3500
Time, sec
Time, s Fig. 10. Fuel temperature as a function of time for a 3.77 mk reactivity insertion with HEU and LEU fuel using the MERSAT code.
Fig. 12. Coolant temperature as a function of time for a 3.77 mk reactivity insertion with HEU and LEU fuel using the MERSAT code.
means that the initiated natural circulation has not yet induced the already exciting water from V-DN and P-DC (Fig. 1). After this period the first layer of heated water emerges in the inlet zone coming from P-UH, where the hot water coming from the core is mixed with the existing cold water, and thereafter the inlet temperature starts to increase. On the other hand, as shown in Fig. 4, the core outlet temperature is increasing continuously reaching a peak of 41.9 C. The continuous increasing in the outlet temperature due to the still low flow velocity of the water emerges from the hotter region. Fig. 5 shows the time evaluation of the coolant core inlet temperature during the operation at nominal power through a period of 2.5 h. The comparison illustrates a good agreement between calculations and measurements and early published data using the RELAP5 code. The continuous increase of inlet temperature indicates that the generated nuclear heat is only partially transferred to the pool water through the tank wall and therefore the primary water doesn’t get cooled sufficiently. Due to the small size of the core, the distance from the inlet orifice to the outlet orifice is short so that the coolant, after being heated in the core, goes up to the upper part of the core (V-UP in Fig. 1), mixes partially with the upper water layer (P-UH in Fig. 1) and then sinks in the downcomer (P-DC). Along this way it transfers only a part of
its enthalpy to the water pool and doesn’t get sufficient cooled. This behavior of insufficient natural circulation speeds up the rise of inlet temperature and consequently the average core temperature, which causing rapidly consumption of the available excess reactivity and shorten thereafter the reactor operation time. Fig. 6 represents the development of core temperature difference Tout Tin between outlet and inlet of the core. It is clear a reasonable agreement between calculated core temperature difference and measurements and the early published data using the RELAP code. This parameter is an indicator of the accuracy of the used form loss coefficients, since it reflects directly time behavior of _ inside the core through the energy balance coolant mass flow (m) _ p ðTout Tin Þ. Q_ ¼ mC Axial power profile taken from the safety analysis report was given in Fig. 7. The axial temperature distributions of fuel centerline, clad and coolant in hottest channel of the three cores were illustrated in Fig. 8(a)e(c), respectively and in Table 2. As shown in Table 2, the maximum clad temperatures were found to be 67 and 56.8 C, LEU (12%) and LEU (12.45%) cores respectively compared with 65 C for HEU. While the maximum centerline fuel temperatures were found to be 78.8 and 77.5 C for the two proposed cores which are higher compared with the original HEU core (69.5 C). 4.2. Control rod withdrawal transient
HEU 12.45% 12%
100
Temperature, C
80
60
40
20 0
500
1000
1500
2000
2500
3000
3500
Figs. 9e12 and Table 3 show the results of a 3.77 mk1 reactivity insertion for the HEU case and for the two LEU cases. The powers and peak clad temperatures with LEU fuel are lower than those with HEU fuel. The peak fuel temperatures are significantly higher for the LEU cases because of the lower thermal conductivity of the LEU oxide fuel. As shown in Table 3, the peak powers for the two proposed LEU cores are 72.3 and 73.5 kW compared with 70 and 73 kW obtained using RELAP-3D code (Dunn et al., 2007). The peak clad temperatures are 102.5 and 101.2 C for the two proposed LEU cores compared with 102 and 98 C obtained using RELAP-3D code. The peak fuel temperatures are 129.9 and 128.4 C for the two proposed LEU cores compared with about 140 C obtained using RELAP-3D code. The difference in result of the two codes could be related to the approximation made in modeling the MNSR reactor using the MERSAT code to go with the one dimensional capability of the code.
Time, s Fig. 11. Clad temperature as a function of time for a 3.77 mk reactivity insertion with HEU and LEU fuel using the MERSAT code.
1 In our previous work (Omar et al., 2010) the RIA was for 3.8 mk and by typing mistake was mentioned for 3.6 mk.
H. Omar et al. / Progress in Nuclear Energy 60 (2012) 140e145
145
Table 3 Comparison of RELAP-3D code and MERSAT code results for steady and transient analysis of MNSR.
U-235 enrichment (%) Steady state
280 240
Temperature, C
LEU (proposed core)
LEU (proposed core)
RELAP-3D
RELAP-3D
200 160
89.8
NA NA 80 102 103
65 69.5 77.6 102.9 113.9
Saturation Temp.
120 80
Coolant Fuel Clad
40 0 0
50
100
150
200
250
300
350
MERSAT
89.8
Onset of Nucleat Boiling
RIA, 3.77 mk
Peak clad surface temperature ( C) Peak centerline fuel temperature ( C) Maximum power (kW) Peak clad surface temperature ( C) Peak centerline fuel temperature ( C)
HEU (existing core) RELAP-3D
400
450
Power, kW
12.45 NA NA 70 102 140
MERSAT 12.45 67 78.8 72.3 102.5 129.9
MERSAT
12
12
NA NA 73 98 140
65.8 77.5 73.5 101.2 128.4
For two proposed LEU cores, maximum clad temperature at the steady-state power level of 30 kW were 67.0 C and 65.8 C, respectively, which are below water saturation temperature (111.8 C). Both of the two proposed LEU cores were also safe (like the existing HEU core) against 3.77 mK positive reactivity insertion transient. Results for LEU cores show similar trends as with HEU core. The clad surface temperatures remain below (by 10 C) the coolant saturation temperature. To check the starting point of onset of nucleate boiling using MERSAT code, the power was increased up to 400 kW for the core with 12.5% enriched fuel. The peak fuel temperature rised linearly with increasing power to the onset of sub-cooled boiling at a power level of 120 kW. It was concluded that any of the proposed LEU cores will also be ‘inherently safe’ and would have safety margins that are higher than for HEU core, owing to very high melting points of oxide fuel (2800 C) and Zr-4 cladding material (1850 C) compared with the melting points of both fuel and clad used in HEU core (about 650 C).
Fig. 13. Coolant, Clad and Fuel temperatures as a function of power with 12.45% LEU fuel using the MERSAT code.
Acknowledgments
4.3. LEU steady-state thermal hydraulic safety margins
The authors thank Professor I. Othman, Director General of the Syrian Atomic Energy Commission, for his encouragement and support of this work.
Steady-state thermal hydraulic safety margins were calculated for the 12.45% LEU case. The results are shown in Fig. 13. The peak clad surface and fuel temperatures rise almost linearly with power to a reactor power of about 120 kW. At this point sub-cooled boiling begins. As shown in Fig.13, at this point, the slope of the curve changes due to the increase in the coolant heat transfer coefficient after the start of sub-cooled boiling. The difference between the saturation and clad temperature was about 9 C. The difference between the saturation and clad temperature, that is necessary for onset of nucleate boiling, is in the range 8.4 C, and 12.2 C (Omar et al., 2010).
5. Conclusion The aim of the present work was to study the thermal hydraulic analysis for core conversion of Syrian MNSR using the MERSAT code as a part of the on-going worldwide effort of converting from highly enriched in uranium (HEU) reactor cores to low enriched in uranium (LEU). It was found that the MERSAT code results of HEU core were in a good agreement with the manufacturer’s results and early published data using the RELAP5 code. Calculations were also carried out for the proposed LEU core with two suggested fuel pin sizes (5.5 mm and 5.1 mm diameter with 12.45% and 12% enrichment, respectively). Comparison of the LEU results with the existing HEU fuel and with the published results using RELAP5-3D code was made and discussed.
References Bokhari, I.H., Showket, P., 2010. Safety analysis for core conversion (from HEU to LEU) of Pakistan research reactor-2 (PARR-2). Nuclear Engineering and Design 240, 123e128. Dunn, F.E., Thomas, J., Liaw, J., Matos, J.E., 2007. MNSR Transient Analyses and Thermal Hydraulic Safety Margins for HEU and LEU Cores Using the RELAP5-3D Code. RERTR Meeting. Hainoun, A., Ghazi, N., 2012. Development of the Thermal Hydraulic Code MERSAT. Final Report under preparation. Hainoun, A., Wolters, J., Hicken, E., 1996. Modeling of void formation in the subcooled boiling regime in the ATHLET code to simulate flow instability for research reactors. Nuclear Engineering and Design, 175e191. Langenbuch, S., 1979. Die Modellgleichungen des 3D-Kernmodells QUABOX-HYCA fuer Leichtwasserreaktoren. GRS-A-256. Lerchel, G., Austregesilo, H., 1996. ATHLET Mode 1.2 Cycle A, User’s Manual. Meister, G., 1989. SIKADE-2, Ein Computer programm zur Simulation der Dynamik von gasbeheizten Dampferzeugern. Juel-Spez-544. Obenchain, C.F., 1973. A Users Guide for the Current ANL Version of the PARET Code, NESC. Omar, H., Ghazi, N., Alhabit, F., Hainoun, A., 2010. Thermo-hydraulic analysis of Syrian MNSR research reactor using RELAP5/Mod3.2 code. Annals of Nuclear Energy 37, 572e581. Richards, B.J., 1988. Validation of the CATHENA two-fluid code, 3rd intern. Topical Meeting on Reactor Thermohydraulic and Operations, Soeul, Korea, 14e17. SAR, 1993. Safety Analysis Report (SAR) for the Syrian Miniature Neutron Source Reactor. CIAE, China. Watanabe, T., Hirano, M., Akimoto, M., Tanabe, F., Kohsaka, A., 1992. MINCS, TwoPhase Flow Characteristics Analysis Code. JAERI 1326. Woodruff, W.L., 1983. Supplemental Information for COBRA-3C/RERT. ANL, Applied Physics Division.
Contents lists available at SciVerse ScienceDirect
Progress in Nuclear Energy journal homepage: www.elsevier.com/locate/pnucene
Thermal hydraulic and safety analysis for core conversion (HEUeLEU) of Syrian Miniature Neutron Source Reactor H. Omar*, N. Ghazi, A. Hainoun Nuclear Engineering Department, Atomic Energy Commission, P.O. Box 6091, Damascus, Syria
a r t i c l e i n f o
a b s t r a c t
Article history: Received 17 November 2011 Received in revised form 13 May 2012 Accepted 26 May 2012
The paper presents the behavior and properties analysis of the low enriched uranium fuel compared with the original high enriched uranium fuel. The MNSR reactor core was modeled with both fuel materials and the reactor behavior was studied during the steady state and abnormal conditions. The MERSAT code was used in the analysis. The steady state thermal hydraulic analysis results were compared with that obtained from the experimental results hold during commissioning the Syrian MNSR. Comparison with experimental data shows that the steady-state behavior of the HEU core was accurately predicted by the MERSAT code calculations. The validated model was then used to analyze LEU cores with two proposed UO2 fuel pin designs. With each LEU core, the steady state and 3.77 mk rod withdrawal transient were run and the results were compared with the available published data in the literatures for the low enriched uranium fuel core. The results reveal that the low enriched uranium fuel showed a good behavior and the peak clad temperatures remain well below the clad melting temperature during reactivity insertion accident. Ó 2012 Elsevier Ltd. All rights reserved.
Keywords: MNSR Safety analysis Core conversion MERSAT
1. Introduction HEU fuel is more economical, the fuel could be used longer in the reactor core, and it has a higher specific reactivity. In the 1970s, however, many people again became concerned about the possibility that some fuels and fuel cycles could provide an easy route to the acquisition of nuclear weapons. Since enrichment to less than 20% is internationally recognized as a fully adequate barrier to weapons usability, certain member states have moved to minimize the international trade in highly enriched uranium and have established programs to develop the technical means to help convert reactors to the use of lowenrichment fuels with minimum penalties. This could involve modifications in the design of the reactor and development of new fuels. As a result of these programs, it is expected that most reactors can be converted to the use of low enriched fuel. The use of new fuel elements in a research reactor would require that the reactor’s current safety analysis report be revised to assess the new balance of safety factors. Principal issues will include the effect of changes in enrichment and fuel technology on temperature and void coefficients of reactivity, thermal
* Corresponding author. Tel.: þ963 11 2132580; fax: þ963 6111926. E-mail addresses: pscientifi[email protected], scientifi[email protected] (H. Omar). 0149-1970/$ e see front matter Ó 2012 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.pnucene.2012.05.007
hydraulic safety criteria, fission product retention, and control system effectiveness. The Syrian Miniature Neutron Source Reactor (MNSR) is a 30 kW, highly enriched uranium, light-water moderated and cooled, beryllium reflected, tank in pool type research reactor that has been in operation since Mars 1996. It is a small, safe nuclear facility which employs high enriched uranium as fuel. MNSR is designed, manufactured and constructed by China Institute of Atomic Energy (CIAE), Beijing, China. The primary use of the reactor is to provide a neutron source with maximum flux of 1 1012 n/ cm2 s with 10 irradiation tubes; 5 inside and 5 outside the beryllium reflector. The core consists of fuel elements which form a fuel cage. The cage is inside an annular beryllium reflector and rests on a lower beryllium reflector plate. The volume of the vessel is 1.5 m3. The fuel elements are all highly enriched uraniumealuminum alloy extraction clad with aluminum. They are arranged in ten multiconcentric circle layers at a pitch distance of 10.95 mm. The element cage consists of two grid plates, four tie rods and a guide tube for the control rod. The two grid plates and four tie rods are connected by screws. The total number of lattice positions is 350 and the number of fuel elements is 347. The remaining positions are filled with dummy aluminum elements. The core region of the Syrian MNSR is located under 4.7 m of water close to the bottom of a watertight reactor vessel (Figs. 1 and 2).
H. Omar et al. / Progress in Nuclear Energy 60 (2012) 140e145
141
Fig. 1. Schematic diagram for MNSR according to MERSAT model; 1 e annular reflector, 2 e lower reflector, 3 e inlet, 4 e in-gap, 5 e ring, 6 e hole, 7 e out-gap, 8 e outlet, 9 e reactor vessel, 10 e reactor pool, 11 e inner irradiation site, 12 e control rode, 13 e outer irradiation site, 14 e shim tray, 15 e reactor base. Fig. 2. MERSAT nodalization.
The beryllium annulus and lower reflector are spaced to form the lower orifice which controls water flow through the core. The top plate of the core and annulus are spaced to form the upper orifice. There are no pumps or heat exchangers in the system. All coolant flow is due to natural circulation. Heat removal from the tank is by conduction through the tank wall to the pool water. Heat removal from the pool is by evaporation and conduction to the air and by conduction through the wall of the pool. An aluminum tray holds the upper reflector which contains semicircular beryllium shims which are added approximately once a year to compensate for fuel burn-up and samarium poisoning. In the current study, the thermal hydraulic analysis for core conversion of Syrian MNSR using the MERSAT code was conducted as a part of the on-going worldwide effort of converting from highly enriched in uranium (HEU) reactor cores into low enriched in uranium (LEU). Two proposed low-enriched uranium dioxide (UO2) cores were considered as shown in Table 1. 2. Computer code employed (MERSAT code) The code MERSAT (Model for Evaluation of Reactor Safety and Analysis of Thermal-Hydraulics) is one-dimensional, two-phase fluid dynamic. The code is under development and verification at the AECS (Hainoun and Ghazi, 2012). The mechanical nonequilibrium of the two phases is described in the momentum equation using the drift-flux model of Zuber and Findlay. The thermodynamic non-equilibrium of the two phases in the subcooled boiling regime is similar to that implemented in the code
ATHLET (Hainoun et al., 1996). The heat conduction and heat transfer module allows a flexible simulation of fuel rods and structures. The nuclear heat generation is calculated by a pointkinetics approach similar to that in PARET (Obenchain, 1973). The heat transfer package describing the wide range of two phase flow
Table 1 Design parameters of the MNSR.
Fuel U-235 enrichment (%) Cladding material Coolant Moderator Fuel rod outer diameter (mm) Fuel meat outer diameter (mm) Clad thickness (mm) Height of core (mm) Total number of fuel rods U-235 content (g) in the core Reflector
HEU (existing core)
LEU (proposed core)
LEU (proposed core)
UAl4-Al 89.8
UO2 12.45
UO2 12
Aluminum H2O H2O 5.5
Zircoloy-4 H20 H20 5.5
Zircoloy-4 H20 H20 5.1
4.3
4.3
4.2
0.6 230 347
0.6 230 347
0.45 230 347
1009
1353
1264
Metallic beryllium
Metallic beryllium
Metallic beryllium
142
H. Omar et al. / Progress in Nuclear Energy 60 (2012) 140e145
22.5
Inlet temperature Tin , C
Input Temperature, C
22.0
Experiment RELAP MERSAT
36
Calculated RELAP Calculated MERSAT Experiment
21.5
21.0
20.5
32
28
24
20
0
20.0
2000
4000
6000
8000
10000
Time, s 0
100
200
300
400
500 Fig. 5. Comparison of experimental and calculated inlet coolant temperature during operation at nominal power of 30 kw.
Time, s Fig. 3. Comparison of experimental and simulated data for coolant inlet temperature during start-up till reaching the power 30 kw.
25
20
Tout-Tin , c
has been developed by consolidating and restructuring some modules in the codes COBRA-3CeRERTR (Woodruff, 1983), PARET (Obenchain, 1973), MINCS (Watanabe et al., 1992; Richards, 1988) and SIKADE (Meister, 1989; Langenbuch, 1979; Hainoun et al., 1996). The modeling of special hydrodynamic components, like fill, leak, time-dependent volume, is relying on the procedure applied in the code ATHLET (Lerchel and Austregesilo, 1996), and QUABOX-HYCA (Langenbuch, 1979).
15
10
3. The MERSAT model
Experiment RELAP MERSAT
5
The input model for the MERSAT code relies on the original design data of MNSR. Some simplifications and adjustments were applied to achieve the best representation of MNSR components according to the specific features and the one-dimensional approach in MERSAT code. The MNSR reactor was divided into primary and secondary systems. The primary system, hydrodynamics systems, representing the reactor vessel including reactor
0 0
2000
4000
6000
8000
10000
Time, sec Fig. 6. Comparison of experimental and calculated difference inlet and outlet coolant temperature during operation at nominal power of 30 kw.
40
1.0
35
0.8
30
25
Experiment Calculated RELAP Calculated MERSAT
20 0
100
200
300
400
500
Time, s Fig. 4. Comparison of experimental and simulated data for coolant outlet temperature during start-up till reaching the power 30 kw.
Relative Power
Output Temperatute, C
45
0.6
0.4
0.2
0.0 0.00
0.05
0.10
0.15
Distance from bottom, m Fig. 7. Axial power density distribution.
0.20
H. Omar et al. / Progress in Nuclear Energy 60 (2012) 140e145 Table 2 Results of steady-state thermal hydraulic analysis of MNSR.
45
a
Parameter
Temperature, C
40
30
25
HEU LEU-12% LEU-12.45%
20
0.05
0.10
0.15
0.20
Distance from bottom, m
Temperature, C
80
b
70
60
HEU LEU-12% LEU-12.45%
50
40 0.00
0.05
0.10
0.15
HEU LEU LEU (existing core) (proposed core) (proposed core)
Fuel pin diameter (mm) 5.5 U-235 enrichment (%) 89.8 Operating power (kW) 30 Pressure at core inlet (kPa) 1.5 111.8 Saturation temperature in core ( C) Peak clad surface 65 temperature ( C) Peak centerline fuel 69.5 temperature ( C)
35
15 0.00
143
0.20
Distance from bottom, m
5.5 12.45 30 1.5 111.8
5.1 12 30 1.5 111.8
67
65.8
78.8
77.5
so that the division took into account all the components of the reactor and if there is any interchange between the components the model should be able to simulate this interchange. Therefore, the sections: V-UP, P-UH, V-DN, P-DC, P-BW, TW-S, TW-M and CCWMS are parts of the reactor vessel immersed in water. The objects G1 to G10 represent coolant channels inside the core, whereas the fuel elements structure is not illustrated in the figure due to their small dimensions. Fig. 2 shows MERSAT nodalization of MNSR. The annular beryllium, bottom beryllium and reactor vessel considered as the heat structure outside the core. Coolant inlet temperature is 20 C and the inlet pressure is 1.5 bar. Values of feedback reactivity coefficients, prompt neutron lifetime and effective delayed neutron fraction in addition to the material properties for the HEU and for LEU fuel were taken from literatures (Dunn et al., 2007; Bokhari and Showket, 2010; Omar et al., 2010).
4. Results and discussion 4.1. Steady state operation
75
c
Computed coolant inlet and outlet temperature during start-up till reaching the power 30 kW were shown in Figs. 3 and 4 respectively. The results agree reasonably well with measurements and the early reported data using the RELAP5 code (Omar et al., 2010). The error in measuring the inlet and outlet temperatures was 5% (SAR, 1993). As shown in Fig. 3 no change in the coolant inlet temperature is observed during the first 300 s, which
70
60
55
80
50
HEU LEU-12% LEU-12.45%
45 40 0.00
0.05
0.10
0.15
0.20
Distance from bottom, m Fig. 8. (a): Axial distribution of coolant temperature in HEU and in the two proposed LEU cores at steady-state power. (b): Axial distribution of fuel centerline temperature in HEU and in the two proposed LEU cores at steady-state power level. (c): Axial distribution of clad temperature in HEU and in the two proposed LEU cores at steadystate power level.
core and reactor primary components. The secondary system representing the reactor pool. Both primaries and secondary system are thermally interconnected through the wall of reactor vessel. The main components of primary and secondary loop are represented in Fig. 1. The reactor was divided into a number of sections
Reactor Power, kW
Temperature, C
65
HEU LEU-12.45% LEU-12%
60
40
20
0 0
1000
2000
3000
Time, sec Fig. 9. Power as a function of time for a 3.77 mk reactivity insertion with HEU and LEU Fuel using the MERSAT code.
144
H. Omar et al. / Progress in Nuclear Energy 60 (2012) 140e145 60
140
HEU 12.45% 12%
120
50
Temperature, C
Temperature, C
100
80
60
40
30
40
HEU LEU-12% LEU-12.45%
20
20 0
500
1000
1500
2000
2500
3000
3500
0
500
1000
1500
2000
2500
3000
3500
Time, sec
Time, s Fig. 10. Fuel temperature as a function of time for a 3.77 mk reactivity insertion with HEU and LEU fuel using the MERSAT code.
Fig. 12. Coolant temperature as a function of time for a 3.77 mk reactivity insertion with HEU and LEU fuel using the MERSAT code.
means that the initiated natural circulation has not yet induced the already exciting water from V-DN and P-DC (Fig. 1). After this period the first layer of heated water emerges in the inlet zone coming from P-UH, where the hot water coming from the core is mixed with the existing cold water, and thereafter the inlet temperature starts to increase. On the other hand, as shown in Fig. 4, the core outlet temperature is increasing continuously reaching a peak of 41.9 C. The continuous increasing in the outlet temperature due to the still low flow velocity of the water emerges from the hotter region. Fig. 5 shows the time evaluation of the coolant core inlet temperature during the operation at nominal power through a period of 2.5 h. The comparison illustrates a good agreement between calculations and measurements and early published data using the RELAP5 code. The continuous increase of inlet temperature indicates that the generated nuclear heat is only partially transferred to the pool water through the tank wall and therefore the primary water doesn’t get cooled sufficiently. Due to the small size of the core, the distance from the inlet orifice to the outlet orifice is short so that the coolant, after being heated in the core, goes up to the upper part of the core (V-UP in Fig. 1), mixes partially with the upper water layer (P-UH in Fig. 1) and then sinks in the downcomer (P-DC). Along this way it transfers only a part of
its enthalpy to the water pool and doesn’t get sufficient cooled. This behavior of insufficient natural circulation speeds up the rise of inlet temperature and consequently the average core temperature, which causing rapidly consumption of the available excess reactivity and shorten thereafter the reactor operation time. Fig. 6 represents the development of core temperature difference Tout Tin between outlet and inlet of the core. It is clear a reasonable agreement between calculated core temperature difference and measurements and the early published data using the RELAP code. This parameter is an indicator of the accuracy of the used form loss coefficients, since it reflects directly time behavior of _ inside the core through the energy balance coolant mass flow (m) _ p ðTout Tin Þ. Q_ ¼ mC Axial power profile taken from the safety analysis report was given in Fig. 7. The axial temperature distributions of fuel centerline, clad and coolant in hottest channel of the three cores were illustrated in Fig. 8(a)e(c), respectively and in Table 2. As shown in Table 2, the maximum clad temperatures were found to be 67 and 56.8 C, LEU (12%) and LEU (12.45%) cores respectively compared with 65 C for HEU. While the maximum centerline fuel temperatures were found to be 78.8 and 77.5 C for the two proposed cores which are higher compared with the original HEU core (69.5 C). 4.2. Control rod withdrawal transient
HEU 12.45% 12%
100
Temperature, C
80
60
40
20 0
500
1000
1500
2000
2500
3000
3500
Figs. 9e12 and Table 3 show the results of a 3.77 mk1 reactivity insertion for the HEU case and for the two LEU cases. The powers and peak clad temperatures with LEU fuel are lower than those with HEU fuel. The peak fuel temperatures are significantly higher for the LEU cases because of the lower thermal conductivity of the LEU oxide fuel. As shown in Table 3, the peak powers for the two proposed LEU cores are 72.3 and 73.5 kW compared with 70 and 73 kW obtained using RELAP-3D code (Dunn et al., 2007). The peak clad temperatures are 102.5 and 101.2 C for the two proposed LEU cores compared with 102 and 98 C obtained using RELAP-3D code. The peak fuel temperatures are 129.9 and 128.4 C for the two proposed LEU cores compared with about 140 C obtained using RELAP-3D code. The difference in result of the two codes could be related to the approximation made in modeling the MNSR reactor using the MERSAT code to go with the one dimensional capability of the code.
Time, s Fig. 11. Clad temperature as a function of time for a 3.77 mk reactivity insertion with HEU and LEU fuel using the MERSAT code.
1 In our previous work (Omar et al., 2010) the RIA was for 3.8 mk and by typing mistake was mentioned for 3.6 mk.
H. Omar et al. / Progress in Nuclear Energy 60 (2012) 140e145
145
Table 3 Comparison of RELAP-3D code and MERSAT code results for steady and transient analysis of MNSR.
U-235 enrichment (%) Steady state
280 240
Temperature, C
LEU (proposed core)
LEU (proposed core)
RELAP-3D
RELAP-3D
200 160
89.8
NA NA 80 102 103
65 69.5 77.6 102.9 113.9
Saturation Temp.
120 80
Coolant Fuel Clad
40 0 0
50
100
150
200
250
300
350
MERSAT
89.8
Onset of Nucleat Boiling
RIA, 3.77 mk
Peak clad surface temperature ( C) Peak centerline fuel temperature ( C) Maximum power (kW) Peak clad surface temperature ( C) Peak centerline fuel temperature ( C)
HEU (existing core) RELAP-3D
400
450
Power, kW
12.45 NA NA 70 102 140
MERSAT 12.45 67 78.8 72.3 102.5 129.9
MERSAT
12
12
NA NA 73 98 140
65.8 77.5 73.5 101.2 128.4
For two proposed LEU cores, maximum clad temperature at the steady-state power level of 30 kW were 67.0 C and 65.8 C, respectively, which are below water saturation temperature (111.8 C). Both of the two proposed LEU cores were also safe (like the existing HEU core) against 3.77 mK positive reactivity insertion transient. Results for LEU cores show similar trends as with HEU core. The clad surface temperatures remain below (by 10 C) the coolant saturation temperature. To check the starting point of onset of nucleate boiling using MERSAT code, the power was increased up to 400 kW for the core with 12.5% enriched fuel. The peak fuel temperature rised linearly with increasing power to the onset of sub-cooled boiling at a power level of 120 kW. It was concluded that any of the proposed LEU cores will also be ‘inherently safe’ and would have safety margins that are higher than for HEU core, owing to very high melting points of oxide fuel (2800 C) and Zr-4 cladding material (1850 C) compared with the melting points of both fuel and clad used in HEU core (about 650 C).
Fig. 13. Coolant, Clad and Fuel temperatures as a function of power with 12.45% LEU fuel using the MERSAT code.
Acknowledgments
4.3. LEU steady-state thermal hydraulic safety margins
The authors thank Professor I. Othman, Director General of the Syrian Atomic Energy Commission, for his encouragement and support of this work.
Steady-state thermal hydraulic safety margins were calculated for the 12.45% LEU case. The results are shown in Fig. 13. The peak clad surface and fuel temperatures rise almost linearly with power to a reactor power of about 120 kW. At this point sub-cooled boiling begins. As shown in Fig.13, at this point, the slope of the curve changes due to the increase in the coolant heat transfer coefficient after the start of sub-cooled boiling. The difference between the saturation and clad temperature was about 9 C. The difference between the saturation and clad temperature, that is necessary for onset of nucleate boiling, is in the range 8.4 C, and 12.2 C (Omar et al., 2010).
5. Conclusion The aim of the present work was to study the thermal hydraulic analysis for core conversion of Syrian MNSR using the MERSAT code as a part of the on-going worldwide effort of converting from highly enriched in uranium (HEU) reactor cores to low enriched in uranium (LEU). It was found that the MERSAT code results of HEU core were in a good agreement with the manufacturer’s results and early published data using the RELAP5 code. Calculations were also carried out for the proposed LEU core with two suggested fuel pin sizes (5.5 mm and 5.1 mm diameter with 12.45% and 12% enrichment, respectively). Comparison of the LEU results with the existing HEU fuel and with the published results using RELAP5-3D code was made and discussed.
References Bokhari, I.H., Showket, P., 2010. Safety analysis for core conversion (from HEU to LEU) of Pakistan research reactor-2 (PARR-2). Nuclear Engineering and Design 240, 123e128. Dunn, F.E., Thomas, J., Liaw, J., Matos, J.E., 2007. MNSR Transient Analyses and Thermal Hydraulic Safety Margins for HEU and LEU Cores Using the RELAP5-3D Code. RERTR Meeting. Hainoun, A., Ghazi, N., 2012. Development of the Thermal Hydraulic Code MERSAT. Final Report under preparation. Hainoun, A., Wolters, J., Hicken, E., 1996. Modeling of void formation in the subcooled boiling regime in the ATHLET code to simulate flow instability for research reactors. Nuclear Engineering and Design, 175e191. Langenbuch, S., 1979. Die Modellgleichungen des 3D-Kernmodells QUABOX-HYCA fuer Leichtwasserreaktoren. GRS-A-256. Lerchel, G., Austregesilo, H., 1996. ATHLET Mode 1.2 Cycle A, User’s Manual. Meister, G., 1989. SIKADE-2, Ein Computer programm zur Simulation der Dynamik von gasbeheizten Dampferzeugern. Juel-Spez-544. Obenchain, C.F., 1973. A Users Guide for the Current ANL Version of the PARET Code, NESC. Omar, H., Ghazi, N., Alhabit, F., Hainoun, A., 2010. Thermo-hydraulic analysis of Syrian MNSR research reactor using RELAP5/Mod3.2 code. Annals of Nuclear Energy 37, 572e581. Richards, B.J., 1988. Validation of the CATHENA two-fluid code, 3rd intern. Topical Meeting on Reactor Thermohydraulic and Operations, Soeul, Korea, 14e17. SAR, 1993. Safety Analysis Report (SAR) for the Syrian Miniature Neutron Source Reactor. CIAE, China. Watanabe, T., Hirano, M., Akimoto, M., Tanabe, F., Kohsaka, A., 1992. MINCS, TwoPhase Flow Characteristics Analysis Code. JAERI 1326. Woodruff, W.L., 1983. Supplemental Information for COBRA-3C/RERT. ANL, Applied Physics Division.