Application of coupled code technique to a safety analysis of a standard MTR research reactor

Nuclear Engineering and Design 239 (2009) 2104–2118

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Application of coupled code technique to a safety analysis of a standard MTR research reactor Tewfik Hamidouche a,c,∗ , Anis Bousbia-Salah b,1 , El Khider Si-Ahmed c , Mohamed Yazid Mokeddem d , Franscesco D’Auria b a

Division de l’Environnement, de la Sûreté et des Déchets Radioactifs, Centre de Recherche Nucléaire d’Alger (CRNA), Alger, Algeria Dipartimento di Ingegneria Meccanica, Nucleari e della Produzione–Facoltà di Ingegneria, Università di Pisa, Pisa, Italy Laboratoire de Mécanique des Fluides Théorique et Appliquée, Faculté de Physique, Université Des Sciences et de la Technologie Houari Boumediene, (USTHB), Bab-Ezzouar, Alger, Algeria d Division de la Physique et des Applications Nucléaires, Centre de Recherche Nucléaire de Draria (CRND), Algeria b c

a r t i c l e

i n f o

Article history: Received 2 December 2008 Received in revised form 22 May 2009 Accepted 3 June 2009

a b s t r a c t Accident analyses in nuclear research reactors have been performed, up to now, using simple computational tools based on conservative physical models. These codes, developed to focus on specific phenomena in the reactor, were widely used for licensing purposes. Nowadays, the advances in computer technology make it possible to switch to a new generation of computational tools that provides more realistic description of the phenomena occurring in a nuclear research reactor. Recent International Atomic Energy Agency (IAEA) activities have emphasized the maturity in using Best Estimate (BE) Codes in the analysis of accidents in research reactors. Indeed, some assessments have already been performed using BE thermal–hydraulic system codes such as RELAP5/Mod3. The challenge today is oriented to the application of coupled code techniques for research reactors safety analyses. Within the framework of the current study, a Three-Dimensional Neutron Kinetics Thermal–Hydraulic Model (3D-NKTH) based on coupled PARCS and RELAP5/Mod3.3 codes has been developed for the IAEA High Enriched Uranium (HEU) benchmark core. The results of the steady state calculations are sketched by comparison to tabulated results issued from the IAEA TECDOC 643. These data were obtained using conventional diffusion codes as well as Monte Carlo codes. On the other hand, the transient analysis was assessed with conventional coupled point kinetics–thermal–hydraulic channel codes such as RELAP5 stand alone, RETRAC-PC, and PARET codes. Through this study, the applicability of the coupled code technique is emphasized with an outline of some remaining challenges. © 2009 Elsevier B.V. All rights reserved.

1. Introduction

Abbreviations: 3D-NKTH, 3-Dimensional Neutron Kinetics Thermal–Hydraulic Model; ANL, Argonne National Laboratory (USA); BE, Best Estimate Code; CFE, control fuel elements; CNEA, Comisiön Nacional de Energia Atömica, Argentina; COBE, Coupled Best Estimate Codes; CRNA, Nuclear Research Centre of Algiers; CSC, Cross Sections Code; DIMNP, Dipartimento di Ingegneria Meccanica, Nucleari e della Produzione – Facoltà di Ingegneria, Università di Pisa; EIR, Eidgenössisches Institut für Reaktorforschung, Switzerland; HEU, high enriched uranium; IAEA, International Atomic Energy Agency; INTERATOM, German Atomic Commission; LOFA, loss of flow accident; MTR, material test reactor; NK, neutron kinetic; NKC, Neutron Kinetic Code; NKTH, neutron kinetics and thermal hydraulic; PVM, parallel virtual machine; RCP, IAEA Research Coordinated Project; REA, rod ejection accident; RIA, reactivity insertion accident; RR(s), research reactor(s); SAR, safety analysis report; SFE, standard fuel elements; Tcmax , maximum clad temperature; TH, thermal hydraulic; THSC, Thermal–Hydraulic System Code; Twmax , maximum coolant temperature; SS, steady state; UNIPI, University of Pisa (Italy); XS, cross section. ∗ Corresponding author. E-mail addresses: [email protected], [email protected], [email protected] (T. Hamidouche), [email protected] (A. Bousbia-Salah), [email protected] (E.K. Si-Ahmed). 1 Currently moved to Bel V, Belgium. 0029-5493/$ – see front matter © 2009 Elsevier B.V. All rights reserved. doi:10.1016/j.nucengdes.2009.06.002

Nuclear research reactors (RR) have been of great support in the development of nuclear science and technology. Over the past fifty years, research reactors have progressed through a variety of tasks. These include materials research with neutron scattering and diffraction, materials characterization with activation analysis and radiography, isotope production, irradiation testing, training as well as services as centers of excellence in science and technology. According to the IAEA inventory (IAEA-RRDB, 2007), 248 research reactors (RR) are, for the time being, in operation with more than 50% having reached their first criticality more than thirty years ago. It is observed that a great number of these “aged” research reactors gave rise to new safety concerns, in particular in the areas of loss of corporate memory, personnel qualification, maintenance of components and systems, preparation and maintenance of documentation as well as updating the safety analysis report (SAR) in case of core reloading, planned power-up rating, component or system replacement, or in case of a critical occurred event.

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In the field of safety analysis, the needs for reassessment of relevant chapter of the SAR are of great importance. Unfortunately, there are no standardized computer tools and/or procedures for nuclear research reactors. This lack, without being exhaustive, is due to the large differences in research reactor power densities, varieties in fuels, geometries, various modes of operations, and of course purposes of utilization. It should be observed that in general, each of these research reactors (RR) has its own tools and codes (Wenxi Tian et al., 2007; Bousbia Salah et al., 2006a; Bousbia Salah and Hamidouche, 2005; Housiadas, 2002; Mariy et al., 2004; Akaho and Maakuu, 2002; Baggoura et al., 1994; Sofu and Dodds, 1994; IAEA-TECDOC, 1992; Woodruff, 1984; Obenchain, 1969). An attempt to perform standardized safety analysis for RR was initiated, by the IAEA, in the framework of core conversion from highly enriched uranium fuel to low enriched uranium fuel (IAEA TECDOC-233, 1980). For that purpose, a safety-related benchmark problem for an idealized generic 10 MW Material Test Reactor (MTR) light-water pool-type reactor was specified. The objective is to compare calculation methods and tools used in various research centers and institutions. However, one could notice that there are as many tools as participating laboratories. Nowadays, with the sustained development in computer technology, the possibilities of code capabilities have been enlarged substantially. Consequently, advanced safety evaluations and design optimizations, that were not possible a few years ago, can now be performed using new generation of computational tools such as Best Estimate (BE) Codes and Coupled Best Estimate (COBE) Codes. Accordingly, the previous IAEA Research Coordinated Project (IAEA-RCP J7-10.10, 2002) has shown the maturity of using Thermal–Hydraulic System Codes (THSC) BE codes in the analysis of accidents in research reactors (Hamidouche et al., 2008). In addition, during this project (IAEA-RCP J7-10.10, 2002) the importance of transferring NPP safety technology tools and methods to RR safety technology was pointed out. Within this objective, a new RCP is launched in 2009 by the IAEA in order to investigate the application of innovative methods in the safety analysis of research reactors (IAEA CRP J7.10.13, 2009). These new analysis methods have to be thoroughly evaluated, with specific emphasis on the capabilities of producing results that should, in general terms, be beneficial to the RR operations and at least reflect acceptable margins compared to conservative tools. In previous works, the assessment and validation of thermal–hydraulic system code have been performed in order to investigate the applicability of RELAP5/Mod3.2 to evaluate safety margins of an MTR RR (Hamidouche and Bousbia-Salah, 2006). In addition some features related to the use of such BE tool have been outlined (Hamidouche et al., 2004). The COBE method was widely assessed for various NPP safety analyses (Bousbia Salah and D’Auria, 2009; Bousbia Salah and D’Auria, 2007; Ivanov and Avramova, 2007; Jeong et al., 2006; D’Auria et al., 2004; IAEA-TECDOC-1539, 2007; Ivanov et al., 2002; Solis et al., 2001). In case of RR accident analysis, several attempts have been performed to apply, partially or entirely, BE tools. In some cases BE Thermal–Hydraulic codes coupled with point kinetic model were used (Woodruff et al., 1996; Hamidouche et al., 2004; Bousbia Salah et al., 2006b; Hedayat et al., 2007; Bokhari et al., 2007) and in other cases BE Neutron Kinetics tools coupled with simple thermal–hydraulic models were applied (Meftah et al., 2006; Khattab et al., 2006; Ahmed et al., 2007; Varvayanni et al., 2007; Waqar et al., 2008). However, applications of BE Neutron Kinetics Code (NKC) and BE thermal hydraulic system code to research reactor safety, are very limited up to now. In a pioneering work, Pautz and Birkhofer (2003) achieved an internal coupling by incorporating a 3D NK transport code (DORT-TD) into a THSC code (ATHLET). The resulting code was, then, applied to the analysis of RIA transients in the German FRM-II RR. The authors pointed

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out the feasibility of such technique for better analysis of fast and slow Reactivity Insertion Accidents (RIAs) within a limited computational time. The objective of the present paper is concerned by the application of the coupled codes technique to research reactor safety analysis and its assessment. This goes through the development of all required steps ranging from cross section generation through unit cells modeling, and full core simulation under steady state and transient conditions. For this purpose the cross section WIMSD5, the 3D Neutron Kinetic (NK) diffusion code PARCS and the Thermal–Hydraulic (TH) System Code (THSC) RELAP5/Mod3.3 code are used. The model is successfully applied to generate the steady state and the dynamic behavior under RIA of the IAEA standard MTR benchmark HEU core. 2. Overview of computational tools and methods One of the objectives of safety analysis is to have all safety requirements satisfied. In other words sufficient margins exist between real values of important safety parameters and their threshold values at which damage of the barriers against the release of radioactivity would occur. To reach such goals established deterministic methods are used based on the conservative and the BE analysis approaches (Kyrki-Rajamaki, 1998). 2.1. Conservative approach The conservatism methodology is generally introduced to take into account the uncertainties induced by the limited modeling capabilities, the computational tools, and the knowledge of physical phenomena. Conservatism is usually used through biasing the plant conditions and physical models to their extreme values in relation to the specific acceptance criteria. These biased conditions allow getting margins that, in any case, cannot be over passed. The approach consists of defining the hot channel factors: concept generally used to account for the uncertainties related to the fuel design, the water gap thickness and for imprecise neutron flux representation in the core. The uncertainties are generally given as the combination of three separate components corresponding to uncertainties (Woodruff, 1997): (1) that influence the heat flux, (2) in the temperature rise or enthalpy change in the channel, (3) in the heat transfer coefficient. These parameters are combined either through a product, statistically or by a combination of both. Examples of engineering factor development methods could be found in the work of Woodruff (1997) and Hu and Bernard (1998). 2.2. Best Estimate (BE) approach Currently, there is a strong tendency to perform accident analysis using BE codes (D’Auria et al., 2001). It allows to reduce conservatism in the calculation models and to get more realistic simulation under normal and abnormal transient conditions and to establish a more effectiveness productivity balance and a wider range of acceptance criteria (IAEA-Safety Report Series 52, 2008). However, conservative analyses are still widely used to overcome the needs of developing realistic models based on experimental data, or simply to avoid the tedious task in changing the approved code and/or the approaches or procedures to get licensing. Although the acceptability of a methodology to use for accidents analysis needs to be defined by the regulatory body, applying only conservative approaches (models, input data

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Fig. 1. Concept of safety margins.

and plant conditions) may lead to unwarranted results (Bishop, 1999). 2.3. Safety margins The gap between conservative and BE methods for defining the safety margins is illustrated in Fig. 1. There are two ways of defining safety margins: either in absolute conditions, in relation to expected damage of safety barriers or in relation to acceptance criterion, typically set up by the regulatory body. In the conservative approach, the results are expressed in terms of a set of calculated conservative values, while in Best Estimate approach the results are expressed as uncertainty ranges for each of the calculated parameters (IAEAN-SG-1.2, 2002). 3. Coupled code method In order to take into account for realistic phenomenon that occurs in a nuclear power plant, COBE codes, which incorporate both BE neutron kinetic and BE thermal hydraulic modules are considered. As shown in Fig. 2, the coupled code method needs: • a code for deriving neutron kinetics Cross Sections (CSC or Cross Section Code). • a Thermal–Hydraulic System Code (THSC). • a Neutron Kinetics Code (NKC). The CSC can be used out of line since a global cross section library is only needed; the latter is obtained as output from CSC codes and then used by coupled THSC and NKC which interact at each time step during the phenomenon simulation. However a qualification process should be followed to fulfill a correct application of such complex codes to the prediction of tran-

sient scenarios (D’Auria and Galassi, 1998). Accordingly, one can mention that the application of RELAP5/mod3 to RR was already assessed (Hamidouche et al., 2004; Woodruff et al., 1996; Hedayat et al., 2007). Furthermore, the latest’s version of RELAP5/Mod3.3 was made suitable for RR applications (Allison et al., 2005). The code RELAP5/Mod3 assessment studies have shown good agreements with some limitations on the subcooled model (Hamidouche et al., 2004; Woodruff et al., 1996). However, this limitation will not have great impact in the range of transients that are considered in the following study. Nevertheless, a validation and assessment effort should be sustained in order to ensure a better capture of the main phenomena during normal and abnormal RR conditions. The PARCS code, standalone, could not be applied to reproduce the Steady State (SS) of the current benchmark core since the embedded Thermal–hydraulic routines are not suitable for low pressure and temperature operating conditions. Nevertheless, coupling PARCS with the RELAP5 will disregard this issue. Hence, the second condition will be demonstrated partly in the present paper since the objective is to compare the 3D NKC/THSC model to other tabulated results from the IAEA TECDOC-643 (1992) for both steady state and transient conditions. 4. Benchmark problem The IAEA-10 MW HEU core consist of 5 × 6 grid core containing 21 Standard Fuel Elements (SFEs) and 4 Control Fuel Elements (CFEs) arranged in the configuration shown in Fig. 3a. The core is reflected by graphite on two opposite sides and surrounded by light water. The standard fuel elements contain 23 plates whereas the control fuel elements contain 17 standard plates and a special region to receive the four-fork-type absorber blades; each absorber blade is composed of Ag (80%), In 14.6%) and Cd (4.9%) contained in a 4 mm layer of nickel on each side. For steady state conditions, the boundary conditions fixed for this configuration are zero flux at a distance of 3 SFE width in X and Y direction as shown in Fig. 3b. Cross views of the SFE and CFE are shown in Fig. 4, and detailed specifications of this benchmark are described in Table 1. The benchmark problem consists in the determination of the steady state core conditions and its behavior under postulated initiating events like RIA and LOFA. 5. Problem modeling 5.1. PARCS model

Fig. 2. Coupled 3D Neutron Kinetics Thermal Hydraulic System Codes.

PARCS/2.7 (Downar et al., 2004) is a three-dimensional (3D) reactor core simulator which solves the steady state and time-

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Fig. 3. MTR benchmark core cross section and specified boundary conditions.

dependent, multi-group neutron diffusion and SP3 transport equations in orthogonal and non-orthogonal geometries. The feedback effects are evaluated through a thermal hydraulic calculation modulus via a few group cross sections. A simple thermal hydraulic model is incorporated in the code for core temperature computation. For better thermal hydraulic parameters evaluation, the code PARCS could be coupled to external system codes like TRAC, TRACE or RELAP5 (Downar et al., 2004). The few-group cross sections library is generated for different conditions of coolant density and/or temperature, fuel temperature and boron concentration. The macroscopic nodal cross sections are function of boron concentra-

tion (B, in ppm), square root of the effective fuel temperature, the moderator temperature, the density, the void fraction and the effective “rodded” fractions. In terms of symbols, the cross sections are evaluated as follows (Downar et al., 2004):



(B, Tf , Tm , Dm , ˛, )



= ˙0 + a1 (B − B0 ) + a2 (

Tf −



T0 ) + a3 (Tm − Tm0 )

+ a4 (Dm − Dm0 ) + a5 (Dm − Dm0 )2 + a6 ˛ + a7 ˛2 +

Fig. 4. Cross section view of Standard (SFE) and Control (CFE) Fuel Elements.

˙CR

(1)

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Table 1 Benchmark core specifications. Active core height Space at the grid plate per fuel element Fuel element

600 mm 77 mm × 81 mm

Meat dimensions

63 mm × 0.51 mm × 600 mm

UAlx -Al fuel

Enrichment 93 w/o (weight%) U-235 280 g U-235 per fuel element 21 w/o of uranium in the UAlx -Al

76 mm × 80.5 mm (including support plate) 76 mm × 80.0 mm (without support plate)

Density of aluminum-canning

2.7 g/cm3

Support plate - Thickness - Density

4.75 mm 2.7 g/cm3

Fuel plate - Thickness - Number per fuel element - Number per control element

1.27 mm 23 17

Identification of the remaining plate positions of the control element: 4 plates of pure aluminum (Al = 1.7 g/cm3 ), each 1.27 mm thick in the position of the first (1st), the third (3rd), the twenty-first (21st), and the twenty-third (23rd) standard plate position; water gaps between the two sets of aluminum plates Graphite element - Dimension 77 mm × 81 mm - Density 1.7 g/cm3 Total power

10 MWth

where ˙ 0 = reference cross section; B: boron concentration; T: temperature; D: density; ˛: void fraction; : control rod inserted fraction; ai : the partial derivatives of cross sections with respect to the control rod fraction and other feedback variables (fuel temperature, coolant temperature, density, boron concentration). The fuel temperature is the effective Doppler fuel temperature and can be computed using either the volume average fuel temperature or a linear combination of the centerline and surface temperatures as described in the PARCS theory manual (Downar et al., 2004). The effective “rodded” fraction is defined as the product of the volume of the “rodded” fraction and the flux depression fac-

tor that is computed by solving a three-node problem using a finite difference method (FDM) (Downar et al., 2004). The main modeling problem, in the PARCS code, is to represent the physical system with a numerical model. This is achieved by the development of various issues important in the reactor kinetics calculation such as the core geometry, the cross section representation as well as the associated thermal–hydraulic parameters. These topics are addressed hereafter. 5.1.1. Geometric and kinetic representation Taking into account the benchmark problem specifications (see Fig. 3 and Table 1), the PARCS core model is defined in XYZ geometry. In this model, each fuel assembly is represented by a homogenized region of four (04) nodes in the XY-plan (two nodes in each direction X and Y) and 21 nodes in Z direction. This subdivision was adopted in order to isolate the central flux trap which contains only water. It should be mentioned that there was no indication about the number of fuel plates in the central fuel elements besides the flux trap. Therefore, in the present study, in order to maintain the core symmetry, each of these central fuel elements is considered as a half SFE (i.e. containing equivalent volume of 11.5 fuel plates). The planar nodalization of the active part of the benchmark core for PARCS is shown in Fig. 5. Each fuel element is depicted by a set of corresponding cross section for each identified region (1–7) including burn-up and geometry as specified in the benchmark problem. The end boxes of the fuel elements are described by a special reflector region consisting of a homogenized region of 15.0 cm Al–H2 O reflectors at the top and bottom (20% Al–80% H2 O volume fractions). This region is represented by a set of cross section and is ascribed a configuration number 6 (not shown in Fig. 5). The water surrounding the core is modeled as region 2. The number of water elements in the X and Y directions is chosen in order to comply with the boundary conditions fixed by the benchmark problem (see Fig. 3b). 5.1.2. Cross section development To solve the neutron kinetics equations, the macroscopic cross section library for various materials in the core is set-up. For this purpose the WIMS-D5 lattice code is used (Askew et al., 1966). In practice, the cells which may correspond to any region of the core (fueled and non-fueled) are identified. When defining the unit cell

Fig. 5. X-Y PARCS model.

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Table 2 Fresh core atom number densities for 93%. N (×1024 )

Fig. 6. SFE WIMS/D5 unit cell.

dimensions, the principle of conservation of volume ratio of the different materials in the fuel assembly is considered. The fuel cell dimensions are calculated taking into account the fuel meat conservation criteria. The unit cell for SFE is shown in Fig. 6. An extra region is added accounting for the remaining water and aluminum in the same proportions as in the physical fuel element; this region includes the aluminum in the plates beyond the width of the meat and the aluminum side plates, the water beyond the width of the meat, and the water channels surrounding the fuel element. In the particular case of CFE, the super cell option of WIMS-D5 is used. The representative cell is modeled with 15 regions as shown in Fig. 7. The same cell is used for absorber-in or absorber-out. In case of absorber-out, the regions 10, 11 and 12 are filled by water. The same procedure has been adopted to define the super cell for the endboxes and the central trap cross-sections calculations. The atom densities used in WIMS-D5 are summarized in Table 2 for the fuel cells, the graphite cell, the water cell, the SFE Endboxes and central trap cells. The macroscopic cross-section data were generated by the WIMS-D5 code as function of burn-up, fuel temperature and coolant density. Different burn-up values, ranging from 0% to 50% by steps of 5%, were considered in order to generate all the conditions fixed in the benchmark specification problem for Fresh, Beginning of Life (BOL) and End of Life (EOL) cores. Six (06) Fuel temperatures and seven (07) water densities were chosen in order to cover a large set of core conditions for normal and transient conditions (see Table 3). The WIMS-D5 with 69-group microscopic cross section (XS) library was run using the SLAB geometry option and a buckling of 7.8367 × 10−3 cm−2 . The cell XS was generated for 69-Group and then collapsed to the desired number (5 and 2 energy groups) of

Fuel U235 U238 Al

0.0016179 0.0001202 0.0570110

Al Graphite

0.060260 0.085234

H2 O O H

0.033428 0.066856

Axial end boxes (20% Al–80% H2 O) Al O H

0.0120520 0.0267424 0.0534848

groups using the MTR-PC package (NEA, 2003). The macroscopic cross section table has been developed according to the PARCS lookup table format for “rodded” and “unrodded” fuel assemblies, and for other regions such as graphite, water, the central trap and the endboxes. The organization of the lookup table is shown in Table 3. The developed set of cross sections was verified qualitatively using the code-to-code assessment. Two sets of XS were developed and assessed: the first one for five energy groups was compared to the XS generated for five quite similar energy groups developed by (ANL-USA) using EPRI-CELL and the second set for two energy groups by (EIR-Switzerland) using WIMS-D1, a native WIMS-D5 code. The XS were developed at a specified coolant density (0.998) and fuel temperature (20 ◦ C). For illustration, only the comparison of the two energy groups XS sets are shown in Table 4. In general, some discrepancies were observed which could be explained by the differences in the original microscopic XS libraries used in each case (in case of WIMS-D1 code (IAEA-TECDOC-233), the 69-group microscopic cross section used data of 1980 whereas, in the current study, an up-to-date 69-XS library is used (WLUP, 2001)). Indeed, according to the work of Siraj-ul-Islam and Nasir (2005), the differences in the microscopic libraries could affect the criticality results by almost 2–3%. As shown in Fig. 5, the CFE are modeled as region 7 in the PARCS model. According to the PARCS XS requirements, two libraries of XS are provided: 1st one for “absorber-out” and named as “Unrodded library” and the 2nd for “absorber-in” and named as

Fig. 7. CFE WIMS/D5 super cell (absorber in).

Table 3 Macroscopic cross-section lookup tables format. Burnup (%) Tf (fuel temperature, K) m (moderator density, kg/m3 )

Macroscopic cross-sections (cm−1 )

293 0.6

311 400

323 800

348 960

373 980

473 990

x (m1 , Tf1 ) . . . . . . x (m7 , Tf1 )

x (m1 , Tf2 ) . . . . . . x (m7 , Tf2 )

x (m1 , Tf3 ) . . . . . . x (m7 , Tf3 )

x (m1 , Tf4 ) . . . . . . x (m7 , Tf4 )

x (m1 , Tf5 ) . . . . . . x (m7 , Tf5 )

x (m1 , Tf6 ) . . . . . . x (m7 , Tf6 )

999

2110

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Table 4 Comparison of cross sections sets. Group

XS

Fresh-fuel (0%)

BOL (5%)

Wims-D1 (EIR)

WIMS-D5 (DESDR)

Wims-D1 (EIR)

WIMS-D5 (DESDR)

Wims-D1 (EIR)

WIMS-D5 (DESDR)

Fast

D1 Abs1 Fiss 1 Nu Fiss1

1,4971E+00 3,2921E−03 1,8654E−03 4,5515E−03

1,5180E+00 3,1018E−03 1,7523E−03 4,2904E−03

1,49710E+00 3,29210E−03 1,86540E−03 4,55150E−03

1,51750E+00 3,05350E−03 1,67520E−03 4,08750E−03

1,51800E+00 3,10180E−03 1,75230E−03 4,29040E−03

1,51710E+00 3,00140E−03 1,59307E−03 3,88710E−03

Thermal

D1 Abs1 Fiss 1 Nu Fiss1 Upscat

2,9508E−01 9,3341E−02 6,8800E−02 1,6787E−01 2,6838E−02

2,9580E−01 9,3128E−02 6,8697E−02 1,6748E−01 2,4577E−02

2,95080E−01 9,33410E−02 6,88000E−02 1,67870E−01 2,68380E−02

2,96080E−01 9,40790E−02 6,51393E−02 1,58940E−01 2,46100E−02

2,95800E−01 9,31280E−02 6,86970E−02 1,67480E−01 2,45770E−02

1,59307E−03 9,11330E−02 6,23934E−02 1,52240E−01 2,46640E−02

Table 5 Control rod worth from different codes and labs.

EOL (10%)

5.2. Thermal–hydraulic modeling

Code/model

Worth ($)

3D NKTH (PARCS/RELAP5) DIFF 2D(ANL) INTERATOM EXTERMINATOR (OESGAE)

22.10 22.39 22.20 23.50

“Rodded library”. One can use the suggested approach by (Olmos and Hansen, 1975) to model the rodded composition by a simple change in thermal absorption cross section. This is done by changing only the thermal capture cross-section of the region of concern by an amount such that the reactivity of an inserted control rod would correspond to the Benchmark case. The control rod worth is estimated by calculating keff with all four absorber blades inserted in the core. The results are shown on Table 5. The axial power shape as function of the control rod position in the core is shown in Fig. 8 for different positions of the four CFE ranging from fully inserted to fully withdrawn positions. An axial peaking factor of 1.5 is obtained bounded by the limits fixed in the TECDOC-643 (1992). The effect of the insertion rate on the flux is illustrated in Fig. 8 whereas Fig. 9 shows the worth of the fully inserted four control fuel elements.

Fig. 8. Control rod insertion effect on axial neutron fluxes along the Z-axis BOL.

The coupling of the RELAP5/Mod3.3 thermal–hydraulic meshes to the PARCS neutronic nodes is done through the assignment of mapping weights between the two meshes. These mapping weights, with values between 0 and 1, inclusive, determine the distribution of neutronic power in the thermal–hydraulic and heat structure components, as well as the calculation of thermal– hydraulic feedback in the neutronic nodes. According to the adopted nodalization scheme, each RELAP5/ Mod3.2 T/H node in X or Y direction is represented by four PARCS nodes. In other words, each four NK nodes of the same fuel element corresponds to one heat structure in RELAP5 and consequently to one pipe. A standard nodalization of a typical MTR research reactor used in a previous representation of the benchmark problem by RELAP5/ Mod3.2 (Hamidouche et al., 2004) was modified (see Fig. 10) to comply with the coupling to PARCS. This adapted new nodalization includes the main reactor components such as the core zone, the reactor pool, the holdup tank, the main coolant pump, and the heat exchanger. The core is represented by 25 heat structures representing individually each fuel element present in the core. The fuel elements are connected to the same upper and lower plenum components using two multiple junctions (mutlipljun) components to specify the inner and upper boundary conditions. Each active part of the fuel elements was divided into 21 axial nodes and 10

Fig. 9. Control rod worth.

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Fig. 10. IAEA benchmark RELAP5/Mod3.3 nodalization.

Table 6 Comparison of criticality calculations. Code/model

Fresh

BOL

EOL

3D NKTH (PARCS/RELAP5) MCNP5 (DIMNP) DIFF 2D(ANL) VIM (MCNP-ANL) CODIFF (EIR) EXTERMINATOR (OESGAE) NEPTUNE (CEA) ADC (JAERI) EXTERMINATOR II (CNEA)

1.194939 1.18962 ± 0.00034 1.1914 1.20165 ± 0.00315 1.19394 1.1966 1.2020 1.18104 1.20018

1.030429 1.05768 ± 0.00032 1.0351

1.008631 1.03959 ± 0.00031 1.0004

1.0368 1.0320 1.0404 1.0420 1.0377

1.0138 1.0404 1.05337 1.02195 1.01425

radial nodes. The coolant region is also divided into 21 axial nodes. The water level above the core was set to 8 m. The assessment of this nodalization has been successfully performed by comparison to results of tabulated boundary conditions specified in the benchmark problem using RELAP5/Mod3.3 and by code-to-code comparison with channels codes (Hamidouche et al., 2004). 5.3. Coupling process The coupling between PARCS and the RELAP5 code is achieved by the inter-process communication protocol based on the Parallel Virtual Machine or PVM. Both processes are loaded in parallel in the PVM machine allowing PARCS to transfer the nodal power data to the RELAP5/Mod3.3 and inversely the latter will send back the temperature (fuel and coolant) and density data to PARCS. This mapping induced the definition of 416 NK planar (X–Y) nodes coupled to 25 planar T–H nodes both combined with 23 axial regions. 6. Results and discussions 6.1. Steady state calculations The 3D NKTH (PARCS/RELAP5) model was run for the criticality calculation of the configuration shown in Fig. 3. The 3D-NKTH

calculated eigenvalues are compared to the results obtained from other diffusion and MCNP codes results provided in (IAEA TECDOC 233, 1980) and (Bousbia Salah et al., 2008). It can be seen in Table 6, that the keff calculated by the 3D-NKTH is within the range of the results obtained from previous codes; these stand from 1.18104 to 1.2020 for fresh core. In general, the deviation is less than 1% for fresh and BOL cores. However, larger discrepancies of keff values are observed for the EOL core. Indeed, at EOL, excess reactivity may be very small and may tend to zero as it is predicted by both ANL and the present calculations. Additional kinetics parameters obtained by 3D NKTH and others codes (IAEA TECDOC-233, 1980; IAEA TECDOC-643, 1992) are outlined in Table 7. As expected and according to design issues, the maximum thermal flux occurs at the central flux trap, as shown in Fig. 11, and drops Table 7 Basic kinetic parameters from different codes and labs. Code/model

BETA (%)

LAMBDA (␮s)

3D NKTH (PARCS/RELAP5) DIFF 2D(ANL) CITATION (JEN) CODIFF (EIR) EXTERMINATOR (JAERI) IAMADY (INTERATOM)

0.760 0.7607 0.736 0.7772 0.7444 0.762

57.50 57.55 51.10 58.80 57.60 54.50

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Fig. 11. 3D-core neutron flux distribution along the X–Y–Z axes (BOL).

exponentially as function of the distance from the core. The thermal flux distribution in X, Y and Z directions obtained by 3D-NKTH (PARCS/RELAP5) are compared, in Figs. 12 and 13, with the results of the ANL Diff-2D diffusion code (IAEA-TECDOC-233, 1980). Similar trends of the flux behavior are observed with some discrepancies mainly due to the code used and the differences in the XS sets of each laboratory. In the current calculations the neutron fluxes are obtained using a XS lookup table with different coolant densities, fuel temperature and burn-up (see Table 2). Hence, they are strongly dependent on local fuel temperature and coolant density distributions. This inherent dependency was not taken into account in the previous calculations where Hot Zero Power (HZP) conditions (fixed fuel temperature (20 ◦ C) and coolant density (0.998)) were considered. Table 8 gathers the fluxes at the central water flux trap obtained by different laboratories and the current model. The observed differences are due to the way the central trap is modeled. In the present case (see Section 4) the central flux trap is divided into four

Fig. 13. Axial neutron fluxes along the Z-axis (BOL).

Table 8 Thermal fluxes (n/cm2 s) in the flux trap from different codes and labs. Code/model 3D NKTH (PARCS/RELAP5) MCNP5 (DIMNP) DIFF 2D(ANL) CODIFF (EIR) EXTERMINATOR (JAERI) EXTERMINATOR II (CNEA)

Fig. 12. Planar neutron fluxes along the X–Y axis (BOL).

Fresh

BOL

EOL

3.1724E+14 2.7803E+14

3.1535E+14 2.8533E+14 2.7518E+14 2.220E+14 – 2.5734E+14

3.2343E+14 2.7833E+14 2.8132E+14 2.285E+14 2.38E+14 2.6409E+14

–

2.220E+14 2.56E+14 1.9813E+14

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Fig. 14. 3D (X–Y–Z) power density distribution (BOL).

nodes and dissociated from the adjacent fuel elements. However, comparative study cannot be conducted since there is no information available on the models adopted by other laboratories. Fig. 14 exhibits a 3D view of the power distribution in the core; the maximum occurs in the peripheral fuel elements beside the graphite reflector and in the proximity of the central flux trap as depicted by the planar representation in Fig. 15. One can notice that the power distribution does not exhibit a 1/4 symmetry as stated in

the calculation hypothesis of IAEA TECDOC-233 (1980). The asymmetrical distribution is confirmed by the Monte-Carlo simulation as shown in Fig. 15 (Bousbia Salah et al., 2008). 6.2. Transient modeling and analysis As specified in the benchmark problem, the exercise concerns the analysis of a reactivity insertion accident (RIA). This category of

Fig. 15. Distribution of power fraction (%)–BOL.

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Fig. 17. Power course during Realistic REA. Fig. 16. Compensated reactivity during Realistic REA.

transients is characterized by positive reactivity addition into the core leading to a prompt core power excursion, followed by strong coupled feedback effects related to Doppler, coolant temperature, and void (if present). The RIA proposed in the benchmark problem is initiated by a super prompt ramp addition of $1.5/0.5 s in the cold core. The safety system trip point is set at 120% of the nominal power; the shutdown system provides a negative reactivity of −$10 in 0.5 s with a response delay time of 0.025 s. In our case, it was not obvious to simulate a linear reactivity insertion as it could be done straightforwardly with point kinetics codes. In case of 3D NKTH, the reactivity inserted is only governed by the control rods movement (i.e. by corresponding cross section). Therefore, the RIA transient is modeled as a rod ejection. In order to overcome the bias of the specified Benchmark core initial conditions where −$10 of shutdown reactivity is available even though the core power is 1 W, new realistic initial conditions are considered, i.e. cold core at low power and low temperature are achieved by all the control rods fully inserted in the core. Under these conditions, the $1.5/0.5 s accident is initiated by the ejection of a single CFE’s control blades. With these realistic start up accident conditions, it is not possible to ensure a SCRAM reserve of −$10/0.5 s. Furthermore, owing to the form of the control rod worth (see Fig. 8), the rate of inserted reactivity during the rod ejection phase is not linear as shown in Fig. 16. For comparison purposes, RELAP5/Mod3.3 standalone utilizing the point kinetics module is used to simulate the same transient with the reactivity calculated by the 3D-NKTH code as an input. In response to the rod ejection, the power behavior exhibits an exponential rise with the decreasing period (see Fig. 17). One can notice that the predicted 3D NKTH power experiences faster excursion than the point kinetics model. This behavior is induced by the fundamental difference that exists between diffusion-models and conventional point kinetic model (Duderstadt and Hamilton, 1976) and more particularly on the prompt neutron generation time which is space–time dependant under super prompt RIA (Ronen, 1986). It should be noticed here that the feedback effect did not play any crucial role during this phase since the temperature rise becomes effective just after the SCRAM time, as shown in Fig. 18. However, even though the transient was faster and the SCRAM

reactivity bank limited, the excursion has been stopped by the reinsertion of the ejected CFE. During the transient the power density peak value changes in time (Fig. 19) and space (Fig. 20) according to the CFE position and also to the thermal–hydraulic feedback (fuel and coolant temperature) and to the flow distribution in the core. The temperatures responses in space and time could be followed during the transient since all the core channels are considered individually (one by one). As shown in Fig. 21, the maximum fuel temperature (hot spot) is observed in the trap peripheral fuel elements (SFE-45% at position D3 and D4 in Fig. 3a); the maximum coolant temperature is observed, at later time scale, in another region of the core different from the region of maximum fuel temperature (see Fig. 22). One can notice that the peak values are not necessarily located at high power density and smaller water gap

Fig. 18. Clad surface temperature behavior during Realistic REA.

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Table 9 Realistic REA simulation results.

Fig. 19. Peak power peaking factor evolution during Realistic REA.

Parameter

3D NKTH

RELAP5/Mod3.3

Trip (s) Power (MW) Tcmax (◦ C) Twmax (◦ C)

0.573 481.9 (0.614) 359.61 (0.650) 96.88 (0.765)

0.652 180.4 (0.707) 286.6 (0.838) 88.89 (0.779)

zone as it is, usually, the case in the conservative approach. It results that the position of the hottest region in the core during the transient is easily identified using the COBE model, while this is not possible using conventional channel codes or standalone BE codes such as RELAP5/Mod3 where the hottest channel is fixed in advance in the code input. It has been also found that the maximum power, cladding and coolant temperatures are higher than the one predicted by the RELAP5/Mod3.3 standalone (see Table 9). This means that realistic simulation of transient considering real physical core conditions and parameters may reverse the conservative approach. In fact, even if the biased conditions of the Benchmark problem are used, the same conclusions could be drawn. Table 10 shows the results of calculations in which the initial conditions consider a core with all the control rods are out except one. The latter is positioned

Fig. 20. 2D and 3D power distribution at specific time.

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Fig. 21. 2D and 3D core temperature distribution at time of maximum fuel temperature.

around the middle of the core in order to get a quasi-linear insertion of $1.5/0.5 s. Here again one can notice that the excursion is faster and the peak power and temperatures values obtained by the COBE code are greater than the conventional codes. 6.2.1. Concluding remarks The model developed for PARCS and RELAP5/Mod3.3 codes has been successfully applied to the IAEA HEU Benchmark cores. The results of steady state analysis emphasized an asymmetrical power distribution. However good agreement is observed compared to previous results obtained by Best Estimate standalone neutron diffusion and channel or thermal–hydraulics codes using point kinetics modules. Under transient conditions, the 3D-NKTH model

emphasized the fact that results obtained via conservative tools are not necessarily, as expected, conservative. But, this is not a general statement since it concerns only fast RIA in HEU core where feedbacks (in particular Doppler effect) do not play an important role during the excursion phase. In general, through the current study a demonstration of the applicability of qualified best-estimate system codes to RR accident analysis was carried out. This technique emphasized a more realistic view and a good description of spatiotemporal physical phenomena occurring during the transient. As a rule, with the objective of achieving and maintaining a high level and quality of safety analysis of research reactors, one can base on the benefits of the experience available from NPP safety

Table 10 REA simulation results using Benchmark Hypothesis. Parameter

3D NKTH*

RELAP5 (UNIPI)

PARET (ANL)

RETRAC (CRNA)

Trip (s) Power (Mw) Tcmax (◦ C) Twmax (◦ C)

0.584 248.82 (0.628) 186.46 (0.630) 86.05 (0.790)

0.609 131.17 (0.655) 163.31 (0.673) 80.30 (0.770)

0.609 129.01 (0.655) 155.25 (0.672) 84.32 (0.76)

0.608 128.44 (0.655) 162.04 (0.668) 82.97 (0.745)

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Fig. 22. 2D and 3D core temperature distribution at time of maximum coolant temperature.

technology; in particular, it is recommended to provide sufficient effort for assessment and validation (using experimental data from Integral Test Facilities) of codes and techniques before any effective application. Accordingly, the next step will be oriented towards the validation of this model through the analysis of some Integral experiments such as the RIA accidents experimented during the SPERT project (Neal and Toole, 1961). Acknowledgment The authors wish to express special thanks to Dr. H. Mazrou for his fruitfull discussions. References Ahmed, R., Nasir Ahmad, A., Javed Khan, M., Tanvir, M., 2007. Utilization of highdensity fuel and beryllium elements for the neutron flux enhancement in typical MTR type research reactors. Progress in Nuclear Energy 49, 246–252. Akaho, E.H.K., Maakuu, B.T., 2002. Simulation of reactivity transients in a miniature neutron source reactor core. Nuclear Engineering and Design 213, 31–42. Allison, C., Hohorst, J.K., Antariksawan, A., Huda, M.Q., Liu, T., Zmitkova, 2005. Validation of RELAP/SCDAPSIM/MOD3.4 for Research Reactor Applications. In: Proceeding of the 13th International Conference on Nuclear Engineering (ICONE13/2005). Askew, J.R., Fayers, F.J., Kemshell, P.B., 1966. A general description of the code WIMS. Journal of British Nuclear Energy Society.

Bishop, G.E., 1999. Removing Unreasonable Conservatisms. DOE Safety Analyses, Safety Analysis Working Group, Portland, OR, June 13–18, 1999. U.S. Department of Energy. Bokhari, I.H., Mahmood, T., Chaudri, K.S., 2007. Steady-state thermal hydraulic analysis of the equilibrium core of Pakistan research reactor-1. Annals of Nuclear Energy 34, 836–838. Baggoura, B., Hamidouche, T., Bousbia Salah, A., 1994. RETRAC: A computer code for the analysis of materials test reactors. Nuclear Science and Engineering 118. Bousbia Salah, A., D’Auria, F., 2009. Application of coupled code method for the analysis of VVER-1000 coolant transient Benchmark. Progress in Nuclear Energy 51, 146–154. Bousbia Salah, A., Benkharfia, H., Kriangchaiporn, N., Petruzzi, A., D’Auria, F., Ghazi, N., 2008. MTR benchmark static calculations with MCNP5 code. Annals of Nuclear Energy 35, 845–855. Bousbia Salah, A., D’Auria, F., 2007. Use of coupled code technique for Best Estimate safety analysis of nuclear power plants. Progress in Nuclear Energy 49, 1–13. Bousbia Salah, A., Meftah, B., Hamidouche, T., Si-Ahmed, E.K., 2006a. Model for the analysis of loss of decay heat removal during loss of coolant in MTR pool type research reactors. Annals of Nuclear Energy 33, 405–414. Bousbia Salah, A., Jirapongmed, A., Hamidouche, T., D’Auria, F., White, J.R., 2006b. Assessment of RELAP5 model for the University of Massachusetts Lowell Research Reactor. Nuclear Technology and Radiation Protection Journal XXI, 1–12. Bousbia Salah, A., Hamidouche, T., 2005. Analysis of the IAEA 10 MW research reactor exercises by RETRAC-PC code. Nuclear Engineering and Design 235, 661–674. D’Auria, F. (Project Coordinator), Bousbia Salah, A., Galassi, G.M., Vedovi, J., Reventos, F., Cuadra, A., Gago, J.L., Sjoberg, A., Yitbarek, M., Sandervag, O., Garis, N., Anhert, C., Aragones, J.M., Verdù, G., Mirò, R., Hadek, J., Macek, J., Ivanov, K., Rizwan-Uddin, Sartori, E., Rindelhardt, U., Rohde, U., Frid, V., Panayotov, D., 2004, Neutronics/Thermal-Hydraulics Coupling in LWR Technology—CRISSUE-S WP2: State-of-the-Art Report. OECD/NEA-No. 5436, Vol. 2.

2118

T. Hamidouche et al. / Nuclear Engineering and Design 239 (2009) 2104–2118

D’Auria, F., Fischer, K., Mavko, B., Sartmandjiev, A., 2001. Validation of accident and safety analysis methodology. IAEA Report CRP/J.020.03. D’Auria, F., Galassi, G.M., 1998. Code validation and uncertainties in system thermalhydraulics. Journal of Progress in Nuclear Energy 33, 175–216. Downar, T., Lee, D., Xu, Y., Kozlowski, T., 2004. PARCS v2 7 U.S. NRC core neutronics simulator. Code Manual. https://engineering.purdue.edu/PARCS. Duderstadt, J.J., Hamilton, L.J., 1976. Nuclear Reactor Analysis. Jonh Wiley & sons, Inc. Hamidouche, T., Bousbia Salah, A., Si-Ahmed, E.K., D’Auria, F., 2008. Overview of accidents analysis in research reactors. Progress in Nuclear Energy 50, 7–14. Hamidouche, T., Bousbia-Salah, A., 2006. Relap5/3.2 assessment against low pressure onset of flow instability in parallel heated channels. Annals of Nuclear Energy 33, 510–520. Hamidouche, T., Bousbia Salah, A., Adorni, M., D’Auria, F., 2004. Dynamic calculations for the IAEA Safety MTR research reactor Benchmark problem using RELAP5/3.2 Code. Annals of Nuclear Energy 31, 1385–1402. Hedayat, A., Davilu, H., Jafari, J., 2007. Loss of coolant accident analyses on Tehran research reactor by RELAP5/MOD3 2 code. Progress in Nuclear Energy 49 (7), 511–528. Housiadas, C., 2002. Lumped parameters analysis of coupled kinetics and thermalhydraulics for small reactors. Annals of Nuclear Energy 29, 1315–1325. Hu, L.W., Bernard, J.A., 1998. Thermal-hydraulic analysis for the upgraded MIT nuclear research reactor. IEEE Transactions on Nuclear Science 45. IAEA CRP J7.10.13, 2009. Innovative Methods in Research Reactor Analysis: Benchmark Against Experimental Data on Neutronics and Thermal-Hydraulic Computational Methods and Tools for Operation and Safety Analysis of Research Reactors. IAEA Research Reactor Safety Division. IAEA Safety Reports Series No. 52, 2008. Best Estimate Safety Analysis for Nuclear Power Plants: Uncertainty Evaluation. IAEA-RRDB, 2007. http://www.iaea.org. IAEA-TECDOC-1539, 2007. Use and Development of Coupled Computer Codes for the Analysis of Accidents at Nuclear Power Plants Proceedings of a Technical Meeting. Vienna, 26–28 November 2003. IAEA-N-SG-1.2, 2002. Safety Assessment and Verification for Nuclear Power Plants Safety Guide. Safety Standards Series No. NS-G-1.2. IAEA-RCP J7-10.10. 2002. Safety Significance of Postulated Initiating Events for Different Research Reactors and Assessment of Analytical Tools. IAEA Research Reactor Safety Division. IAEA-TECDOC, 1992. Research Reactor Core Conversion Guidebook, Vol. 3. Analytical Verification. IAEA-TECDOC-643. IAEA-TECDOC, 1980. Research Reactor Core Conversion from the Use of Highly Enriched Uranium to the Use of Enriched Uranium Fuels Guidebook. IAEATECDOC-233. Ivanov, K., Avramova, M., 2007. Challenges in coupled thermal–hydraulics and neutronics simulations for LWR safety analysis. Annals of Nuclear Energy 34, 501–513. Ivanov, B., Ivanov, K., Groudev, P., Pavlova, M., Hadjiev, V., 2002. VVER-1000 Coolant Transient Benchmark—Phase 1 (V1000CT-1), Vol. I. Final Specifications. US Department of Energy and OECD Nuclear Energy Agency. NEA/NSC/DOC. Jeong, J.J., Lee, W.L., Chung, B.D., 2006. Simulation of a main steam line break accident using a coupled system thermal-hydraulics, three-dimensional reactor kinetics, and hot channel analysis code. Annals of Nuclear Energy 33, 820–828. Khattab, K., Omar, H., Ghazi, N., 2006. The effect of temperature and control rod position on the spatial neutron flux distribution in the Syrian miniature. Nuclear Engineering and Design 236 (23), 2419–2423. Kyrki-Rajamaki, R., 1998. The need of coupled 3D neutronics in DBA and BDBA analyses using conservative or best-estimate approach. In: OECD/CSNI Seminar on Best Estimate Methods in Thermal Hydraulic Analyses, Ankara, June 29–July 1, 1998. Mariy, A.H., El Morshdy, S.E., Khater, H.A., Abdel-Raouf, A.A., 2004. Transient thermal hydraulic modeling of ETRR-2 for off site power loss. In: International Conference on Research Reactor Utilization, Safety, Decommissioning, Fuel and Waste Management, 10–14 November 2004, Santiago, Chili. Meftah, B., Zidi, T., Bousbia-Salah, A., 2006. Neutron flux optimization in irradiation channels at NUR research reactor. Annals of Nuclear Energy 33, 1164–1175. NEA-2003, 2003. MTR-PC, V2.6, Neutronic. In: thermal hydraulic and shielding calculations on personal computers. INVAP, Nuclear Engineering Division, NEA DATA BANK.

Neal, J., Toole, C.R., 1961. Ramp induced power excursion tests in SPERT III at room temperature. Transactions of American Nuclear Society 4 (1), 72. Obenchain C., 1969. PARET: A program for the analysis of reactor transients. AEC Research and Development Report, IDO-17282. Olmos, J., Hansen, K.F., 1975. Light-water Reactor Physics Parameters for Transient Analysis. Massachusetts Institute of Technology—Energy Laboratory Report No. MIT-EL 75-022. June 1975. Pautz, A., Birkhofer, A., 2003. Coupling of time-dependent neutron transport theory with the thermal hydraulics code ATHLET and application to the research reactor FRM-II. Nuclear Science and Engineering 145, 320–341. Ronen, Y., 1986. Handbook of Nuclear Reactors Calculations, vol. II. CRC Press. Siraj-ul-Islam, A., Nasir, A., 2005. Effect of updated WIMSD libraries on neutron energy spectrum at irradiation site of Pakistan Research Reactor-1 using 3D modeling. Annals of Nuclear Energy 32, 521–548. Sofu, T., Dodds, H.L., 1994. A Computer Model for the Transient Analysis of Compact Research Reactors with Plate Fuel. ANL/RA/19957. Solis, J., Ivanov, K., Sarikaya, B., Olson, A., Hunt, K.W., 2001. Boiling Water Reactor Turbine Trip (TT) Benchmark, Vol. 1. Final Specifications. NEA/NSC/DOC. Varvayanni, M., Grigoriadis, D., Catsaros, N., Stakakis, E., 2007. Neutronic calculations for the conversion to LEU of a research reactor core. In: RERTR-2007 International Meeting on Reduced Enrichment for Research and Test Reactors, Prague, Czech Republic, September 23–27, 2007. Waqar, S., Mirza, S.M., Mirza, N.M., 2008. Two-group, three dimensional model based study of reactivity induced transients in upgraded LEU material test reactors. Annals of Nuclear Energy 35 (4), 647–655. Wenxi Tian, W., Qiu, S., Su, G., Jia, D., Liu, X., Zhang, J., 2007. Thermal-hydraulic and safety analysis on China advanced research reactor under station blackout accident. Annals of Nuclear Energy 34, 288–296. WLUP, 2001. IAEA Project for the Update of WIMS-D Libraries. http://wwwrcp.ijs.si/ wlup. Woodruff, W.L., 1997. Evaluation and Selection of Hot Channel (Peaking) Factors for Research Reactor Applications. RERTR Program Argonne National Laboratory. ANL/RERTR/TM-28. Woodruff, W.L., Hanan, N.A., Smith, R.S., Matos, J.E., 1996. A comparison of the PARET/ANL and RELAP5/MOD3.3 codes for the analysis of IAEA benchmark transients. In: International Meeting on Reduced Enrichment for Research and Test Reactors, Seoul, South Korea, October 7–10, 1996. Woodruff, W.L., 1984. A kinetics and thermal-hydraulic capability for the analysis of research reactors. Nuclear Technology 64, 196–206. Tewfik Hamidouche 1983–1988, bachelor degree in physics (solid state physics) from University of Sciences and Technology of Algiers (ALGERIA), 1988–1992, master degree in nuclear engineering from the Nuclear research Center of Algiers (ALGERIA), 2004–2008, PhD degree in physics (fluid mechanics) from University of Sciences and Technology of Algiers (Algeria). His main activities concern Nuclear Safety Analysis of research reactor inkling development and utilization of advanced computational tools. He is the author and co-author of papers in International Journals and Conference Proceedings. Anis Bousbia Salah 1984–1988, bachelor degree in physics (thermodynamics and solar energy) from University of Constantine (Algeria), 1988–1992, master degree in nuclear engineering from the Nuclear Research Center of Algiers (Algeria), 2001–2004, PhD degree in nuclear engineering and industrial safety from University of Pisa (Italy). His main activities concern Nuclear Safety Analysis using advanced computational tools. He is the author of 30 papers in International Journals. El Khider Si-Ahmed bachelor degree in physics (electronic) from University of Algiers (Algeria); MS in nuclear engineering from M.I.T. (USA); DEng in nuclear engineering and science from M.I.T. (USA); PhD in nuclear engineering and engineering physics (fluid mechanics and thermal sciences) from R.P.I (USA). Presently professor of energetic and fluid mechanics and Head of laboratory of theoretical and applied fluid mechanics. His main activities concern experimental and theoretical fluid mechanic, multiphase flow, heat and mass transfer and nuclear safety. He is the author and co-author of more than 90 papers in International Journals and conference proceedings and supervised many PhDs.