ABSTRACT. This paper presents recent advances in the modeling of two-phase boiling flow and critical heat flux that have been implemented in the Extended ...
Proceedings of the 2014 22nd International Conference on Nuclear Engineering ICONE22 July 7-11, 2014, Prague, Czech Republic
ICONE22-30844
COMPUTATIONAL FLUID DYNAMICS MODELING OF TWO-PHASE BOILING FLOW AND CRITICAL HEAT FLUX Adrian Tentner, Elia Merzari, Prasad Vegendla Argonne National Laboratory Argonne, Illinois, USA
ABSTRACT This paper presents recent advances in the modeling of two-phase boiling flow and critical heat flux that have been implemented in the Extended Boiling Framework (EBF) [1, 2, 3]. The EBF code was developed as a customized module built on the foundation of the commercial Computational Fluid Dynamics (CFD) code STAR-CD, which provides general twophase flow modeling capabilities, for the detailed analysis of the two-phase flow and heat transfer phenomena that occur in Boiling Water Reactor (BWR) fuel assemblies. These phenomena include coolant phase changes and multiple flow regimes that directly influence the coolant interaction with the fuel pins and, ultimately, the reactor performance. An effort to expand the EBF two-phase models and to explore their applicability to other CFD codes is currently underway. The paper presents results of recent CFD analyses of Critical Heat Flux (CHF) experiments that have measured the axial distribution of wall temperature in two-phase upward flow in a vertical channel with a heated wall. The experiments were designed to produce the onset of CHF in the upper half of the heated channel. The simulated axial distribution of wall temperature is compared with experimental data, illustrating the ability of the extended EBF model to capture the onset of CHF for a wide range of thermal-hydraulic conditions relevant for BWRs. The paper concludes with a discussion of results and plans for future work.
the CFD code STAR-CD [4]. An effort is underway to port the EBF to the high-fidelity CFD code NEK-5000 [5] which is being extended to provide general two-phase flow modeling capabilities. A first generation of models describing the interphase mass, momentum, and energy transfer phenomena specific for various flow regime topologies have been previously implemented in the CFD code STAR-CD [5]. The EBF boiling models, which describe the inter-phase mass, momentum, and energy transfer phenomena specific for various local flow topologies, allow the simulation of a wide spectrum of flow regimes expected in a BWR fuel assembly [1, 2, 3]. An overview of these models is included below in Section 1. This paper reviews the current status of key two-phase flow phenomenological models and focuses on the extension and validation of models that describe the cladding-to-coolant heat transfer and the onset of CHF. The EBF module uses a local inter-phase surface topology map in conjunction with models for the inter-phase mass, momentum, and energy exchanges for the bubbly, droplet, and transition flow topologies. It also calculates the conjugate heat transfer using a wall heat transfer model that describes the heat exchange between the heated wall and the two-phase or single-phase coolant. It is shown that the wall heat transfer model used in conjunction with various local flow topologies allows the prediction of the onset of CHF for a wide range of thermalhydraulic conditions relevant for BWRs without the use of empirical correlations traditionally used in sub-channel codes.
INTRODUCTION The Extended Boiling Framework (EBF) was developed for the fine-mesh, 3-dimensional simulation of the two-phase flow phenomena that occur in a Boiling Water Reactor (BWR) fuel assembly. These phenomena include coolant phase changes and multiple flow topologies that directly influence the reactor performance. The EBF was developed as a specialized module built on the foundation of
NOMENCLATURE Latin
G h
1
g
e
internal energy, enthalpy, J/kg gravitational acceleration, 9.806 m/s2 flow rate, kg/(m2s) heat transfer coefficient, W/(m2K)
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turbulence kinetic energy, m2/s2 mass transfer rate, kg/m3 sum of the inter-phase forces, N/m3 pressure, Pa heat flux, W/m2 inter-phase heat transfer, W/m3 time, s temperature, K saturation temperature, K velocity vector, m/s axial velocity scalar m/s
Mp
k m
u
q Q t T Tsat w
Greek
τ τ
,
t
determines the local flow configuration as a function of flow conditions and prescribes which models and properties are relevant for each computational cell. The sub-cooled bubbly flow topology, with vapor bubbles flowing in a continuous liquid has an established base of CFD modeling experience. The Extended Boiling Framework uses an inter-phase surface topology map that includes, in addition to the bubbly topology, a droplet or mist topology and a transition topology. The droplet topology consists of liquid droplets flowing in a continuous vapor stream. An additional wall-cell topology is used for cells adjacent to walls, as described below. The direct simulation of the transition from bubbly flow, through slug and churn flow, to annular flow was not pursued for practical reasons. To resolve large bubbles of size comparable to channel diameter would consume large computer resources on the necessarily fine grids and short timescales, and above all on the extraction of suitable time-averaged results from the simulation of a chaotic process. Instead, for transitiontopology cells we use a topology-based combination of the terms appropriate for the basic topologies, bubbly and mist. This can be interpreted as having a transition topology cell where a fraction of the cell volume presents the bubbly topology while the remaining volume presents the mist topology. An alternative interpretation is that the map is prescribing the probability of being in one topology or the other while solving equations for the time-averaged flow.
volume fraction void fraction difference characteristic mesh size dissipation rate of turbulence kinetic energy, m2/s3 thermal conductivity, W/(mK) area fraction density, kg/m3 laminar and turbulence shear stresses, N/m2
Subscripts bn bubble nucleation drop liquid droplets i phase interface k phase subscript l liquid g gas q quenching sat saturation state w wall
II. Transport Equations The STAR-CD Eulerian two-phase solver tracks the mass, momentum, and energy of the liquid and vapor phases in each computational cell. Full details of the Eulerian two-phase flow models in STAR-CD can be found in [4, 6]. The main equations solved are the conservation of mass, momentum and energy for each phase. The conservation of mass equation for phase k is:
u
1. TWO-PHASE FLOW MODEL OVERVIEW
τ
k
k
t k
k
. k
(2)
M
g
p
τ
u
k k k . k k t k k k
The CFD-LWR code, also referred to as the Extended Boiling Framework (EBF) was initially developed as a customized module built on the foundation of the commercial CFD-code STAR-CD which provides general two-phase flow modeling capabilities. It is now being expanded with the goal of implementing it in the open-source CFD code NEK-5000. The NEK-5000 code is a high-fidelity single-phase fluid flow code which is currently being expanded to model two-phase boiling flows. During the NEK-5000 transition from singlephase to two-phase modeling capabilities the extension and validation of the EBF is being pursued in the framework of the original STAR-CD implementation. The model development strategy adopted for the simulation of boiling flow phenomena in a BWR fuel bundle was described in [1]. A central concept is that of a local inter-phase surface topology map which
u u
(1) k k . k k k m ki m ik t The conservation of momentum equation for phase k is:
I. Approach
k k ek . k k t
u
The conservation of energy equation for phase k is: e . k k Tk Q
k k
(3)
An extended k - model containing extra source terms that arise from the inter-phase forces present in the momentum equations is used to model turbulence in the flow [6]. III. Inter-phase Surface Topology Map and Local Flow Configuration Inter-phase interactions in multiphase fluids depend on both the area and the topology of the phase interface. Sub-
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channel thermal-hydraulic codes rely on flow regime maps to evaluate the interface topology using cross-section-averaged flow parameters. CFD codes, which divide the flow space into much finer computational cells cannot rely on the traditional sub-channel flow regimes, but must evaluate instead the local inter-phase surface topology. The ensemble of many computational cells with relatively simple inter-phase surface topologies can provide complex global topologies that include all the traditional sub-channel flow regimes. Most of the advanced CFD codes currently allow the simulation of dispersed flows only (bubbly or mist flows) where the flow topology is originally defined and remains the same in both space and time. This approach is applicable only to flows without topology changes and without sharp interfaces. Flow topology changes are typical for boiling flows in BWR fuel assemblies as illustrated schematically in Figure 1a. Moreover, local sharp interfaces often exist in boiling flows, e.g., between Taylor bubbles and the near-wall film in slug flow, or between the gas core and the liquid film in annularmist flow in heated pipes or rod bundles. The changing topology conditions typical for boiling flows in BWR fuel assemblies, including sharp-interface topologies such as the annular-mist flow regime, cannot be modeled accurately with topology-independent correlations used for inter-phase interactions in such flows. Meanwhile, it is essential to adequately describe these regimes in BWR channel simulations, since they govern near-wall film formation and evolution, which, in turn, determines important fuel assembly characteristics such as CHF and wall dryout.
fraction. This 1-dimensional topology map is illustrated in Fig. 2 together with the associated flow topologies.
Figure 2 Inter-phase Surface Topology Map used for the Extended Boiling Framework IV The wall-cell sharp-interface topology The topology map discussed above does not address directly the presence of cells that contain a sharp singleconnected interface, such as wall cells that contain a thin liquid film. Wall cells that satisfy specific conditions imposed on both the local void fraction and void fraction gradient are treated in the EBF boiling model as a special liquid film topology illustrated in Figures 3a and 3b. The wall cells can contain both a liquid film and liquid droplets. As the liquid volume fraction decreases the liquid film, which initially covers the entire wall surface (Figure 3a) is assumed to become unstable and to cover the wall surface only partially (Figure 3b).
Figure 3a Full liquid film topology Figure 1 Schematic view of upward flow in vertical channel with heated walls The EBF boiling model uses a locally calculated topology variable to allow the following topologies: a) a bubbly flow topology with spherical vapor bubbles in a continuous liquid, b) a droplet or mist topology with spherical liquid droplets flowing in a continuous vapor field, and c) a transition topology which combines the features of the two previous topologies in various proportions. The local topology is determined in this model using a local topology map based on the local void
Figure 3b Partial liquid film topology
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Because only one liquid velocity and one liquid temperature are defined for each cell, the droplet and film velocities and temperatures are the same. However, the partition of the liquid between the liquid film and droplets has important implications for the heat transfer between the heated wall and the two-phase coolant. The explicit modelling of the film entrainment and the partition of liquid between the film and droplets in the wall cells have been the focus of recent work aimed at improving the prediction of CHF. These models are described in Section 2.
focused on separate boiling effects [1, 2, 8]. The integral model validation was focused on the analysis of the OECD-NEA/USNRC Benchmark based on NUPEC BWR Full-size Fine-mesh Bundle Tests (BFBT) [9]. The initial validation of the claddingto-coolant heat transfer model discussed in Section 2 below was presented in [10]. 2. CLADDING-TO-COOLANT HEAT TRANSFER MODEL The calculation of wall heat flux is based on partitioning of heat flux between the following four heat transfer components: a) qg, the convective heat flux from wall to the gas or vapor phase, over the wall area not covered by the liquid film; a) ql, the convective heat flux from wall to the liquid phase, over the area covered by the liquid film but excluding the bubble nucleation area; c) qi, the evaporation heat flux from wall to the boiling interface, over the bubble nucleation area; and d) qq, the quenching heat flux from the wall to the liquid over the nucleation area. These heat flux components are specified per unit wall area and they decrease or become zero as the corresponding wall-contact area decreases. The wall heat flux is given by:
V. Inter-phase mass, momentum, and energy transfer models The inter-phase surface topology map is used to evaluate the interfacial area and inter-phase interactions. Three basic local flow configurations with specific interface topology are identified (bubbly flow, mist flow and sharp interface) and the interfacial area and inter-phase mass, momentum, and energy transfer models are defined for these configurations. In the domain identified in Fig. 2 as transitional topology it is assumed that a combination of basic flow configurations is present, and the quantities required for closure are found by determining the appropriate combination of mass, momentum, and energy exchange terms for the local flow topology. The most general transitional topology is illustrated in Fig. 4, and various other transitional topologies are obtained by retaining only a sub-set of the master-cell features.
q w = qg + q l + q i + qq
(4)
The heat flux components are calculated using the appropriate heat transfer coefficients and heat transfer area based on the local wall-cell topology: 1 1 1
1
(5) (6) (7) (8)
Where is the ratio of the wall-area covered by vapor to the total wall area, and is the ratio of the area covered by bubble nucleation sites to the area covered by liquid. Figure 4 Master-cell topology
The evaluation of the heat transfer coefficients and area used in equations (5) through (8) was ratios and presented in [1] and [10]. The heat transfer model enhancement presented in this paper has focused on improving the evaluation the wall area ratio which determines the heat flux transferred to the liquid film and the bubble nucleation sites, and thus plays a central role in the simulation of CHF conditions. The previous model for the evaluation of in wall-cells with a thin liquid film used in [10] was:
As demonstrated in [7], the use of the local inter-phase surface topology map allows the modeling of complex subchannel-scale topologies that emerge from combinations of many computational cells with one of the local topologies shown in Figures 2 and 3. E.g., the typical sub-channel annular flow regime could be resolved into a distinct core flow region in which the gas phase is continuous and the local mist topology is used, separated by transition topology cells from a liquid film on the wall where the local bubbly topology and the wall-cell topology are used.
DRY
VI, Two-Phase Flow Model validation
g 1DRY max 0, min 1, DRY DRY 2 1
(9)
Rigorous validation efforts during the EBF model development phase have analyzed numerous two-phase flow experiments
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Where the recommended breakpoints were 0.9 and 0.95. This model assumes that the wall area covered by the liquid film decreases rapidly as the vapor volume fraction and . It does not account increases between explicitly for the presence of both liquid droplets and film in the wall-cells. Although the model using equation (9) was able to predict the occurrence of CHF reasonably well in two experiments analyzed in [10], subsequent analyses of CHF experiments with higher flow rates and lower power-to-flow ratios showed the model using equation (9) cannot predict accurately the CHF location over a wider range of flow conditions. To improve the EBF capability to predict the occurrence of dryout conditions, the wall-to fluid heat transfer model was modified to account for the simultaneous presence of liquid droplets and a liquid film in the wall-cells, based on the wall-cell topologies shown in Figures 3a and 3b. The liquid in a wall-cell divided between the film and droplets:
described in this section lead to a significantly improved prediction of CHF conditions over a wide range of flow parameters relevant for BWRs, as described below in Section 4. Future work will explore the replacement of the droplet entrainment correlation (13) with mechanistic models of the droplet entrainment and deposition processes. 3. CRITICAL HEAT FLUX EXPERIMENTS ANALYZED To validate the wall heat transfer models included in the EBF we analyzed several static dryout experiments conducted by Becker, et al., [11]. These experiments were designed to study CHF and post-dryout heat transfer in vertical circular pipes. The loop consisted of a 7 m long test section, a condenser, feed water and main recirculation pumps, flow measuring devices and a preheater. Subcooled water was fed at the bottom of the test section. The wall was heated uniformly and all typical BWR flow regimes were produced in the upward water/steam flow. In the experiments the outer wall temperature was measured, and the inner wall temperature was calculated assuming an adiabatic boundary condition. The experimental data are presented as axial distributions of the inner wall temperature. A schematic of the experimental test section is presented in Figure 5. Six experiments in a vertical channel 0.01 m in diameter and 7 m in length with a uniformly heated wall were used as verification test-cases. Pressure was approximately 7 MPa and the inlet subcooling was approximately 5 - 10 K in all the experiments analyzed. The experiments differed in inlet mass flux G and wall heat flux qw as shown in Table 1, which also includes the ratio qw/G for each experiment.
(10)
The model used to evaluate is modified to account for the presence of liquid droplets in the vapor field:
DRY
g drop 1DRY max 0, min 1, 2DRY 1DRY
(11)
When using equation (11) the fraction of wall-area covered by the film is determined by the liquid associated with the film, excluding the liquid droplets. To use this new modelling capability we need to describe the partition of the liquid in the wall-cells between the liquid film and liquid droplets. The model implemented partitions the liquid using the entrainment defined as the difference between the vapor velocity velocity in the bulk flow and the vapor velocity in the wall-cell:
(12)
The bulk-flow vapor velocity is obtained from the cell next to the wall-cell in the radial direction. The liquid volume fraction assigned to the liquid droplets is calculated as follows:
w E w1E E E w w 1 2
drop l max 0, min 1,
(13)
2.5 / and Where the recommended breakpoints are 5.5 / . Equation (13) reflects the assumption that an increasing fraction of the liquid film is entrained by the vapor stream in the form of liquid droplets as the entrainment velocity increases. It is noted that the liquid film and the liquid droplets in a wall-cell retain the same velocity and temperature as the two-fluid model implemented in the STAR-CD code allows only one liquid and one vapor component in each cell. However, the changes in the wall-to-coolant heat transfer model
Figure 5 Schematic of the experimental test section
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There are no wavy film regions in the simulation of experiments D through F (see Figs.6) which were performed at higher heat flux. The actual topology of the liquid film cannot be determined from the pipe wall temperatures measured in the Becker experiments. These experiments did not allow visual observation of the flow conditions.
Table 1 Parameters of the CHF experiments analyzed Experiment A B C D E F
G [kg/(sm2] 497.0 1009.6 1008.9 1495.0 1994.9 2482.9
qw [W/m2] 35.0e4 40.1e4 49.9e4 79.7e4 79.6e4 80.0e4
qw/G 704 397 495 533 399 322
For the experiments D through F, which were conducted with a significantly higher heat flux the location of the sharp rise in the wall temperature agrees well with the location of the location of the measured wall temperature rise, as shown in Figures 7d through 7f. The dryout location in these experiments moves further away from the pipe inlet as the ratio qw/G shown in Table 1 - decreases. The wall temperature rise and general shape of the wall temperature curve in the post-dryout region agree remarkably well with the corresponding measured results for these experiments. For experiment D the location of the sharp increase in the wall temperature shown in Figure 7d coincides with the disappearance of the liquid film/droplet region near the wall as observed in Figure 6d. This indicates that the volume fraction of the liquid droplets in the wall cells at the dryout location is small or zero. For experiments E and F the liquid film/droplet region near the wall continues well beyond the dryout location. In these cases the film has disappeared due to entrainment and evaporation, but the higher liquid fraction region persists due to the presence of the entrained droplets. The extended cladding-to-coolant heat transfer model described in Section 2 plays a key role in the correct prediction of the wall dryout conditions and wall temperature changes in these cases. In the experiments E and F the calculated wall temperature peaks near the dryout location and decreases afterwards, in good agreement with the experimental temperature (Figures 7e and 7f). Acording to Hoyer [13], the wall temperature decrease in the post-dryout region is caused by evaporation of water droplets in superheated steam above the dryout elevation. This evaporation produces steam and therefore the vapor velocity increases rapidly. This in turn increases the wall-steam heat transfer coefficient and decreases wall temperature. This is consistent with the presence of droplets generated by film entrainment in the wall-cells at and above the dryout location, as shown in Figures 6e and 6f. For experiment D, where the calculated amount of droplets in the wall-cells at the dryout location is low, the post-dryout wall temperature decrease is much smaller (Figure 7d). A slight increase in the calculated wall temperature is observed near the end of the pipe, indicating that the amount of liquid droplets at the dryout location may be slightly under-predicted in this case. However, the overall agreement between the calculated and measured wall temperatures in the six experiments analyzed is considered satisfactory. Planned analyses of additional CHF experiments will provide additional about the EBF performance in predicting the CHF conditions under a wide range of conditions of interest for LWR analyses.
4. RESULTS AND DISCUSSION The calculated void fraction distributions are shown for the six experiments analyzed in Figures 6a through 6f. The characteristic sub-channel flow regimes in a pipe with heated walls are simulated. Since the inlet temperature was only slightly below saturation the bubbly flow regime was limited to the near-inlet region of the pipe, but the slug, annular-mist and mist flow regimes are clearly observed. The corresponding calculated wall temperature is shown in Figures 7a through 7f. We discuss first the experiments A through C, which had lower heat flux values than the experiments D through F. For the experiments A and C the location of the calculated sharp rise in the wall temperature agrees reasonably well with the location of the measured wall temperature rise as shown in Figures 7a and 7c. In both cases the calculated sharp rise in the wall temperature coincides with the disappearance of the calculated liquid film (Figures 6a and 6c), indicating that there are few droplets at the dryout location and the liquid droplet entrainment does not play a significant role in these cases. The slope of the calculated wall temperature in the post-dryout region is similar to the corresponding measured temperature slope, but the calculated temperatures are lower than the measured values by approximately 50 - 70 K. The dryout location in experiments A and C moves further away from the pipe inlet as the ratio qw/G - shown in Table 1 - decreases. For experiment B, which has a lower qw/G ratio than experiment C (heat flux lower than C but a similar flow rate) the calculated wall temperature does not exhibit a sharp rise, in agreement with the measured temperatures indicating that film dryout did not occur in this case. In the simulation of the experiments A through C wavy film regions can be observed in Figures 6a through 6c. These waves are unlikely to be a numerical effect, since: a) the peakto-peak distance is about 7-10 times the longitudinal cell size, b) the waves do not disappear with mesh refinement, and c) the waves do not change noticeably with the number of iterations. These wavy film regions could be the effect of physical liquid film instability (“vapor chimney” mentioned by Sugawara [12]) and appear to be associated with the lower heat flux conditions.
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Figure 6a Experiment A - Vapor volume fraction
Figure 7a Experiment A - Twall versus z
Figure 6b Experiment B - Vapor volume fraction
Figure 7b Experiment B - Twall versus z
Figure 6c Experiment C - Vapor volume fraction
Figure 7c Experiment C - Twall versus z
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Figure 6d Experiment D - Vapor volume fraction
Figure 7d Experiment D - Twall versus z
Figure 6e Experiment E - Vapor volume fraction
Figure 7e Experiment E - Twall versus z
Figure 6f Experiment F - Vapor volume fraction
Figure 7f Experiment F - Twall versus z
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5. CONCLUSIONS
5. P. F. Fischer, J. Lottes, W.D. Pointer, A. Siegel. “Petascale algorithms for reactor hydrodynamics”, J. Phys. Conf. Series (2008) 6. S.Lo, “Modeling multiphase flow with an Eulerian approach”, von Karman Institute Lecture Series, Industrial Two-Phase Flow CFD, May 23-27, von Karman Institute, Belgium (2005) 7. A. Tentner, S. Lo, A. Splawski, A. Ioilev, V. Melnikov, M. Samigulin, V. Ustinenko,. "Computational Fluid Dynamics Modeling of Two-Phase Flow and Inter-Phase Surface Topologies in a BWR Fuel Assembly," Proceedings of ICONE16, the 16th International Conference on Nuclear Engineering, Orlando, FL, USA, May 11-15, 2008. 8. V. Ustinenko, M. Samigulin, A. Ioilev, S. Lo, A. Tentner, A. Lychagin, A. Razin, V. Girin, Y. Vaniukov, “Validation Of CFD-BWR, A New Two-Phase Computational Fluid Dynamics Model For Boiling Water Reactor Analysis,” Proceedings of CFD4NRS: OECD/NEA International & International Atomic Energy Agency (IAEA) Workshop on Benchmarking of CFD Codes for Application to Nuclear Reactor Safety, Garching, Munich, Germany, September 5-7, 2006. 9. A. Tentner, W. D. Pointer, S. Lo, A. Splawski, "Integral Validation of a CFD Model for the Simulation of Two-Phase Flow Phenomena in a Boiling Water Reactor: Analyses of the BFBT Full Bundle Tests", Nuclear Engineering and Design, June 2009 10. A. Ioilev, M. Samigulin, V. Ustinenko, P. Kucherova, A. Tentner, S. Lo, A. Splawski “Advances in the modeling of cladding heat transfer and critical heat flux in boiling water reactor fuel assemblies”, Proc. 12th International Topical Meeting on Nuclear Reactor Thermal Hydraulics (NURETH12), Pittsburgh, Pennsylvania, USA, September 30-October 4, 2007 11. K.M.Becker, C.H.Ling, S.Hedberg, G.Strand, An experimental investigation of post dryout heat transfer. Department of Nuclear Reactor Engineering, Royal Institute of Technology, KTH-NEL-33, Sweden, 1983. 12. S.Sugawara. Analytical prediction of CHF by FIDAS code based on three-fluid and film-dryout model. J. Nucl. Sci. Technol., 1990, Vol.27, No.1, p.12-29. 13. N.Hoyer, Calculation of dryout and post-dryout heat transfer for tube geometry. Int. J. Multiphase Flow, 1998, Vol.24. No.2. p.316-334.
We presented recent advances in the modeling of claddingto-coolant heat transfer and critical heat flux that have been implemented in the Extended Boiling Framework. The validation and extension of the EBF CHF models is performed in the context of the CFD code STAR-CD. The cladding-tocoolant heat transfer model is used in conjunction with the local wall-cell topology, which has been extended to allow the simultaneous presence of a liquid film and liquid droplets. The extended cladding-to-coolant heat transfer model accounts for the entrainment of liquid droplets and corresponding reduction of the liquid film. Six experiments involving upward boiling water flow and dryout in a heated circular channel were analysed. Comparisons of the calculated wall temperatures with the corresponding measured values show that the EBF including the extended cladding-to-coolant heat transfer model provides a reasonably accurate prediction of the onset of Critical Heat Flux (CHF) for a wide range of flow rates and wall heat fluxes relevant for BWRs, without the use of the traditional CHF correlations used in sub-channel codes. Future work will focus on the analysis of additional CHF experiments relevant for LWR conditions and further extension of the EBF CHF models.
ACKNOWLEDGMENTS The authors gratefully acknowledge the contributions of Drs. S. Lo and A. Splawski from CD-adapco and Drs. A. Ioilev, M. Samigulin, and V. Ustinenko from Sarov Labs to the initial development and validation of the Extended Boiling Framework (EBF). Support for the initial development of the EBF was provided by the U.S. DOE GIPP Program.
REFERENCES 1. A.Tentner, S.Lo, A.Ioilev, M.Samigulin, V.Ustinenko, Computational Fluid Dynamics Modeling of Two-phase Flow in a Boiling Water Reactor Fuel Assembly, Proc. Int. Conf. Mathematics and Computations, American Nuclear Society, Avignon, France, Sept. 2005. 2. A.Tentner, S.Lo, A.Ioilev, M.Samigulin, V.Ustinenko, V.Melnikov, V.Kozlov, Advances in computational fluid dynamics modeling of two phase flow in a boiling water reactor fuel assembly. Proc. Int. Conf. Nuclear Engineering ICONE-14, Miami, Florida, USA, July 17-20, 2006. 3. A.Tentner, W.D.Pointer, T.Sofu, D.Weber, Development and Validation of an Extended Two-Phase CFD Model for the analysis of Boiling Flow in Reactor Fuel Assemblies, Proc. Int. Conf. Advances in Nuclear Power Plants, Nice, France, May 13-18, 2007. 4. STAR-CD Version 3.20 Methodology Manual, Chapter 13, CD-adapco, UK. 2004.
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