- Email: [email protected]
Design and analysis of a thermal core for a high performance light water reactor
Nuclear Engineering and Design 241 (2011) 4420–4426
Contents lists available at ScienceDirect
Nuclear Engineering and Design journal homepage: www.elsevier.com/locate/nucengdes
Design and analysis of a thermal core for a high performance light water reactor T. Schulenberg a,∗ , C. Maráczy b , J. Heinecke c , W. Bernnat d a
Karlsruhe Institute of Technology, Institute for Nuclear and Energy Technologies (IKET), D-76344 Eggenstein-Leopoldshafen, Germany Hungarian Academy of Sciences (KFKI), Atomic Energy Research Institute, H-1525 Budapest, Hungary AREVA NP GmbH, D-91058 Erlangen, Germany d University of Stuttgart, Institute for Nuclear and Energy Systems (IKE), Pfaffenwaldring 31, D-70569 Stuttgart, Germany b c
a r t i c l e
i n f o
Article history: Received 10 March 2010 Received in revised form 1 June 2010 Accepted 2 June 2010
a b s t r a c t The High Performance Light Water Reactor is a Generation IV light water reactor concept, operated at a supercritical pressure of 25 MPa with a core outlet temperature of 500 ◦ C. A thermal core design for this reactor has been worked out by a consortium of Euratom member states within the 6th European Framework Program. Aiming at peak cladding temperatures of less than 630 ◦ C, including uncertainties and allowances for operation, the coolant is heated up in three steps with intermediate coolant mixing to eliminate hot streaks. Different from conventional reactors, the radial power profile is intended to be non-uniform, with the highest power in the first heat-up step in the core center and the lowest power in the second superheater step to result in the same peak cladding temperatures in each region. The concept has been studied with neutronic, thermal-hydraulic and structural analyses to assess its feasibility. Coupled neutronic/thermal-hydraulic analyses are defining the initial distribution of enrichment, control rod positions and the use of burnable poisons. Sub-channel analyses predict the coolant mixing inside assemblies, and a porous media approach simulates the flow of moderator water between assembly boxes. Finally, structural analyses of the assembly boxes are needed to minimize deformations during operation. Even though the core design cannot yet considered to be final, this state of the art review shall summarize the progress achieved so far and outline the remaining challenges. © 2010 Elsevier B.V. All rights reserved.
1. Introduction In 2006, a consortium of 10 European partners from 8 European member states, supported by the European Commission in their 6th Framework Program, decided to work out a design concept of a supercritical water cooled reactor, which they called the High Performance Light Water Reactor (HPLWR). Starflinger et al. (2007) summarize the envisaged project plan and the time schedule for the 3.5 years duration of the second phase of this project. Compared with existing light water reactors, the advantages of this concept shall be lower plant erection costs, due to a once through steam cycle without the need of recirculation pumps, steam separators and dryers, high turbine enthalpies and thus smaller mass flow rates at a given turbine power, a higher thermal efficiency, and a compact containment design, all leading primarily to economic advantages. Basically, the core of such a HPLWR can have a thermal or fast neutron spectrum. During the course of the project, however, being restricted by manpower, the consortium decided to concentrate on a thermal core design, leaving the fast core studies to Ishiwatari et al. (2008) and thus to gain synergies in the Generation
∗ Corresponding author. Tel.: +49 7247 82 3450; fax: +49 7247 82 4837. E-mail address: [email protected] (T. Schulenberg). 0029-5493/$ – see front matter © 2010 Elsevier B.V. All rights reserved. doi:10.1016/j.nucengdes.2010.09.025
IV International Forum. Meanwhile, the project has passed successfully its midterm assessment (Starflinger et al., 2009), design details have been worked out and first analyses are available to get a first impression about the feasibility of the core concept. 1.1. Core design target Aiming at a net electric power of around 1000 MW and a net efficiency of almost 44%, the target thermal power of the reactor core needs to be 2300 MW, confirmed also by recent steam cycle analyses of Brandauer et al. (2009). Early cycle studies by Dobashi et al. (1998) indicated an optimum thermal efficiency at a feedwater temperature of 280 ◦ C which was kept also for the present study. The target core outlet temperature was chosen as 500 ◦ C which is still rather low for a once through steam cycle with single reheat, compared with latest fossil fired power plants, but appears to be challenging enough with regard to available fuel cladding materials. Their peak temperature limit was targeted at 630 ◦ C which is not only a challenge for oxidation and corrosion protection, but also for their creep strength and resistance to stress corrosion cracking. The fuel centerline temperature is a function of the linear power of the fuel rod. It has been limited to 39 kW/m under nominal conditions. To be competitive with respect to latest pressurized water reactors, the target burn up should be at least 45 GWd/tHM . Near future
T. Schulenberg et al. / Nuclear Engineering and Design 241 (2011) 4420–4426
4421
3. Mechanical design
Fig. 1. Core design concept with evaporator, superheater 1 and superheater 2 assembly clusters (shown exemplarily from right to left) (Koehly et al., 2009).
fuel enrichment plants for light water reactors are expected to be limited by 6% enrichment, which was chosen as the enrichment target accordingly. Like with boiling water reactors, boron acid cannot be used to compensate the excess reactivity at the beginning of a burn-up cycle, so that burnable absorbers must be used instead. The target power and temperatures result in a coolant mass flow rate of 1179 kg/s. Schlagenhaufer et al. (2009) suggest a feedwater pressure of 25 MPa for all load conditions which keeps some margin from the critical pressure of 22.1 MPa.
2. Core layout These target data differ from conventional light water reactors not only by the higher pressure and core outlet temperature, but also by a significantly higher enthalpy rise in the core. Indeed, the difference between life steam enthalpy and feedwater enthalpy of 1936 kJ/kg exceeds the one of pressurized water reactors by around a factor of 8. Assuming an overall hot channel factor of 2 between the peak and the average coolant heat-up, this enthalpy rise would result in peak coolant temperatures of 1200 ◦ C which is far beyond the target temperature limit. A strategy to overcome this issue can be learned from fossil fired boiler design. These boilers are characterized by multiple heat-up steps with intensive coolant mixing between them to eliminate hot streaks. Schulenberg et al. (2008) applied such a strategy for a core layout with a first heat-up of the coolant as moderator water, comparable with the economizer of a fossil fired boiler. The second heat-up should be in the evaporator assemblies in the center of the core, followed by coolant mixing in a plenum above the core. From there, the coolant is directed downwards in assemblies of the first superheater, surrounding the evaporator, to be mixed again in an annular chamber underneath the core. Final heat-up to the envisaged core outlet temperature of 500 ◦ C was proposed to happen in a second superheater stage with upward flow again in assemblies at the core periphery. The core design concept is shown in Fig. 1. Assuming a hot channel factor of 2 for each heat-up step, as an initial guess, the power ratio of evaporator to superheater 1 to superheater 2 should be around 4:2:1 to reach the same peak coolant temperature in each region. The proposed core layout is trying to reach this power ratio by placing the second superheater at the core periphery where the neutron leakage is reducing the neutron flux anyway.
Beyond 390 ◦ C, the coolant density is less than 200 kg/m3 , hardly enough to produce a thermal neutron spectrum. Therefore, colder feedwater is foreseen as moderator water to run inside moderator boxes in the fuel assemblies and in gaps between assembly boxes. As Schulenberg et al. (2008) predicted a pressure drop of about 0.5 MPa from core inlet to its outlet, but the mass of structural material in the core should be minimized to limit the neutron absorption, Hofmeister et al. (2007) concluded that the fuel assemblies should be small, preferably with 40 fuel pins each and a single moderator box in their center to enable a small wall thickness of moderator and assembly boxes. To ease handling during maintenance, they recommended to group 9 assemblies to a cluster with common head and foot piece. Fischer et al. (2007) adopted this idea to design a fuel assembly cluster for the three pass core as described above, preferably such that clusters can be exchanged between evaporator and superheater positions. Wire wraps were proposed as grid spacers to improve coolant mixing in both flow directions. The clusters can be disassembled at their foot piece to exchange single fuel rods for repair. Control rods shall be inserted from the top of the core. They run inside 5 of the 9 moderator boxes. Different control rod designs, e.g. cruciform rods and square tubes, and their shut down reactivities have been studied by Schlagenhaufer et al. (2007). A thermal insulation of assembly and moderator boxes has been proposed by Herbell et al. (2008) using a sandwich design with internal honeycomb structures to reach the envisaged stiffness of less than 0.5 mm deflection towards the fuel rods. With a pitch to diameter ratio of only 1.18, the fuel rod bundle is rather tight, which results in a mass flux of more than 1600 kg/m2 s for a nominal fuel assembly. A lower mass flux might cause concern with respect to the deterioration of heat transfer in the evaporator. The envisaged fuel rod diameter of 8 mm leads to an average core power density of about 60 MW/m3 . Mixing chambers above and underneath the core have been designed by Fischer et al. (2007) and updated by Koehly et al. (2009) to enable a downward flow of moderator water inside the moderator boxes, but an upward flow in the gaps between the assembly boxes. The latest design uses 50% of the feedwater to run at first downwards through the moderator boxes, then upwards between the assembly boxes, and then downwards again to cool the reflector around the core. The remaining fraction of feedwater is supplied through the downcomer to be mixed with the moderator water, released from the reflector, in a mixing volume underneath the core at the evaporator inlet. The upper mixing chamber is aligned near the steam outlet flanges of the reactor pressure vessel. It is separated internally by walls to serve also as the steam plenum above the core outlet. C-ring sealings avoid feedwater ingress into the plenum. Free thermal expansion of the assembly clusters is allowed at their upper ends, where the head pieces are sealed with c-rings at their penetrations through the upper chamber, whereas the foot pieces are standing on the core support plate. This core design concept is certainly quite unusual. It cannot be compared with any other core design built and tested in the past. Detailed analyses are being performed, therefore, to assess if such a three pass core will be feasible. The following chapters will summarize the current status of neutronic, thermal-hydraulic and structural analyses.
4. Analyses of the three pass core concept 4.1. Power distribution The thermal core is characterized by a large change of coolant density, ranging from more than 700 kg/m3 at the core inlet to
4422
T. Schulenberg et al. / Nuclear Engineering and Design 241 (2011) 4420–4426
Fig. 2. Core power distribution in case of fresh fuel with 5% uniform enrichment (Monti et al., 2009a).
Fig. 3. Core power distribution after 200 days burn up (Monti et al., 2009a).
less than 90 kg/m3 at its outlet. Even though additional moderator water has been foreseen to compensate missing coolant density, the local core power is changing significantly with coolant heat-up. As a consequence, neutronic codes need to be coupled with thermalhydraulic codes to account for this effect. Waata et al. (2005) were
the first who coupled the Monte Carlo code MCNP with the subchannel code STAFAS by running them iteratively to obtain the axial power profile of a single assembly of the HPLWR. Meanwhile, the method has been extended to predict the power distribution of the entire core.
Fig. 4. Radial core power distribution, relative numbers, axially averaged (Maráczy et al., 2008).
T. Schulenberg et al. / Nuclear Engineering and Design 241 (2011) 4420–4426
4423
Fig. 5. Peaking factors of the core power in each core region (Maráczy et al., 2008).
Monti and Schulenberg (2009) coupled the neutron transport code ERANOS with the system code TRACE to predict the power distribution of a quarter of the core made from fresh fuel with uniform enrichment. Maráczy et al. (2008) coupled the diffusion code KARATE with the one-dimensional thermal-hydraulic code SPROD to optimize the core power distribution with the help of control rods and burnable poisons. Bernnat and Conti (2009) coupled MCNP5 and STAFAS with ORIGEN to predict burn-up effects of a single assembly and even of a sector of the core. The physical mechanisms causing the power distribution of this strange core design can be understood when we look at the hypothetical case of fresh fuel with 5% uniform enrichment. Monti et al. (2009a) show under this assumption that the power ratio of the individual core regions (evaporator:superheater 1:superheater 2) is rather like 6:2.5:1, meaning that the evaporator is producing even more power than envisaged. This power ratio decreases quickly within 200 days of burn-up to about 3.3:2:1. The comparison is shown in Figs. 2 and 3 for 1/4 of the core. Here, the finest resolution of the power distribution is a fuel assembly cluster. In these figures, 13 of such clusters form the evaporator, the first and the second superheater each. The general strategy to reach the envisaged power ratio of 4:2:1 is therefore to add Gd absorber rods, controls rod or older fuel assemblies to the evaporator. Maráczy et al. (2009), Fig. 4, show a first example of a core power distribution in which the central power peak is suppressed by Gd rods and partially inserted control rods. Here, the finest resolution of the power distribution is a single assembly. 3 × 3 assemblies form a cluster. The evaporator
region and both superheater regions are separated by solid lines. The position of inserted control rods is visible in the evaporator region as local, cross shaped low power regions. These are at cluster positions (1;5), (2;2), (2;4), (3;4), (4,2), (4;3) and (5;1) counted from the core symmetry lines in x- and y-direction, respectively. Once the envisaged power ratio has been reached, the power profile needs to be homogenized in each region to minimize the peak power. We see in particular in the superheater 2 region of Fig. 4 that the radial flux gradient causes hotter assemblies towards the first superheater and colder ones towards the outer reflector. Therefore, a further optimization of the power profile by Maráczy et al. (2009) includes a water layer in the reflector and uses different enrichments to simulate different burn-up stages which are substituted by fresh fuel for this analysis. Fig. 5 shows the peaking factors, i.e. the assembly power normalized with the average power of each core region. Now, the hottest power peak in the second superheater is reduced to 1.35× the average power there. Further design studies are aiming at an optimization of the burnup cycle. Monti et al. (2009b) discuss the importance of having thermally insulated assembly and moderator box walls to keep a high density of the moderator water. Comparing solid steel boxes with insulated boxes, they got an increase of the initial excess reactivity of 8600 pcm by the thermal insulation proposed by Herbell et al. (2008). Bernnat and Conti (2009) give first indications how the burn-up can be increased by shuffling assembly clusters from the second superheater, where some Pu is produced by the harder neutron spectrum, to evaporator cluster positions. A detailed look into the power distribution inside a fuel assembly is indicating that Gd
4424
T. Schulenberg et al. / Nuclear Engineering and Design 241 (2011) 4420–4426
Fig. 7. Coolant temperature profile of the hottest superheater assembly with 20% assumed radial power gradient (Himmel et al., 2008).
Fig. 6. Flow reversal in the gap volume between the assembly boxes if moderator water is supplied from the top (Kunik et al., 2009).
rods should preferably be positioned in the corners of the assemblies. Moreover, their sensitivity study of the influence of assembly box bending on the power distribution inside assemblies will help to quantify hot channel factors based on uncertainties later on. 4.2. Flow structures and coolant mixing While the interaction of the coolant density inside assemblies with the power profile can be modeled with one-dimensional thermal-hydraulic codes, the flow structure in the gaps between assembly boxes and its density distribution need to be modeled at least in two dimensions. Kunik et al. (2009) approximated this flow structure with a porous media approach, assuming that the moderator water in the gaps was supplied from the top. Despite the thermal insulation of the assembly boxes, they conclude that a change of the flow pattern from purely downward to local recirculation may occur, caused by buoyancy of hot zones around the second superheater as shown in Fig. 6. Such change in the flow pattern, however, can effect the power profile which, in turn, can influence the flow structure again. As a consequence, to avoid instabilities, the flow path of the moderator water was changed to a supply through the foot pieces of the assembly clusters and thus to obtain a stable upward flow from there. Excellent coolant mixing inside assemblies and in the mixing chambers above and underneath the core is of crucial importance for this reactor concept. It has been the intent of the wire wrapped fuel rods to serve as mixing devices inside assemblies. The resultant coolant temperature distribution has been checked with the sub-channel code C3CLM, an AREVA NP version of the well known sub-channel code COBRA-IIIC, for worst case power gradients of superheater assemblies. Himmel et al. (2008) validated this subchannel analysis with CFD analyses using locally a fine mesh and globally a coarse mesh approach. As an example, we show in Fig. 7 the coolant temperature distribution of the hottest superheater 2
assembly with a linear power gradient causing 20% more power in the upper row of fuel rods compared with the average power of each cross section. The maximum coolant temperature at the outlet is just 16 ◦ C hotter than the average coolant temperature there. Full CFD analyses of Kiss et al. (2009) confirm this excellent result. Coolant mixing in the upper mixing chamber between the evaporator outlet and the first superheater inlet has been studied numerically by Wank et al. (2008). The rather blocked flow path, running around obstacles like head piece structures, moderator boxes and connection tubes of the mixing chamber made it quite difficult to mix the coolant as intended. Several design modification were tried before a convincing mixing structure could be identified. The complexity of the flow path inside the upper mixing chamber is illustrated in Fig. 8. Another cause for concern is the risk of a deterioration of heat transfer in the evaporator assemblies close to the pseudo-critical point. The phenomenon is well known from boiler tubes where it can cause hot spots in case of high heat flux and low mass flux. First CFD analyses by Chandra et al. (2009) indicate, however, that wires wrapped around the fuel rods help to prevent such effects by increasing the turbulence level. Local flow separation downstream
Fig. 8. Two head pieces shown exemplarily in the upper mixing chamber (Wank et al., 2008).
T. Schulenberg et al. / Nuclear Engineering and Design 241 (2011) 4420–4426
4425
Fig. 9. Design of the assembly and of the moderator box with honeycomb panels and location of spacer pads (Herbell et al., 2008).
of these wires, on the other hand, can cause another hot spot which will need to be examined carefully and should be included in the final hot spot evaluation. 4.3. Stresses and deformations As discussed above, the assembly and moderator boxes need to be thermally insulated to minimize the heat-up of moderator water. Herbell et al. (2008) propose a sandwich design, shown in Fig. 9, with an internal honeycomb structure to provide sufficient stiffness and to stand the pressure difference between the outside and the inside of these boxes. The honeycombs are filled with alumina or zirconia for thermal insulation. They need to be vented to the colder side as they cannot stand an outside pressure of 25 MPa. Deformations and stresses have been studied by Herbell et al. (2008) with Finite Element Analyses of an axial segment of the assembly box. As a result, massive
corner pieces, unfortunately with higher thermal conductivity, were required to keep the local peak stresses below the fatigue limit. Radial power gradients do not only cause a temperature difference of the fuel rods inside an assembly. As the assembly box is thermally insulated, it will experience similar temperature differences as the coolant, forcing it to bend to the hotter side. As a consequence, the gap between the fuel assemblies is reduced or increased from its nominal gap width of 9 mm, may be even until it contacts the neighboring assembly box. This effect influences the neutron moderation and, in turn, the power profile of the assembly. In order to avoid such bowing, two sets of spacer pads, attached to the corners of the box, have been foreseen between the assembly boxes as shown in Fig. 9 on the right hand side. The worst case considered to study the residual bowing line assumes that one side of an evaporator assembly box experiences up to 80 ◦ C hotter temperatures than its opposite side. The bowing analysis shown in Fig. 10
4426
T. Schulenberg et al. / Nuclear Engineering and Design 241 (2011) 4420–4426
Fig. 10. Bowing of the assembly box assuming up to 80 ◦ C temperature difference between the hotter and colder side.
demonstrates that even in this case the maximum box deflection is less than 0.5 mm. 5. Conclusions The thermal core design of the HPLWR, though being highly innovative and anything but easy with its different heat-up regions and its upward and downward flow path, turned out to be feasible in all analyses performed up to now. Thanks to the repeated mixing, it has sufficient margin from maximum cladding temperature limits, which have not yet been reached by any of these predictions. The wire wrapped fuel rods are expected to help not only to improve mixing inside assemblies but also to avoid a deterioration of heat transfer close to the pseudo-critical point of water. At lot more needs to be analyzed, however, before the core design can finally be assessed. The peak cladding temperatures of individual fuel rods need to be predicted taking into account different burn up stages, control rod positions, all kinds of tolerances and uncertainties as well as allowances for operation. The latter ones are just being optimized by dynamic steam cycle analyses of Schlagenhaufer et al. (2009). Accidental conditions need to be studied, not only to design the safety system but also to predict the fuel and cladding temperatures under transient and abnormal conditions. In particular, the consequences of natural convection phenomena in heated downward flow regions should be examined with care. In so far, the design is not yet considered to be final but still subject to further improvements. Acknowledgement This work has been funded by the European Commission as part of their project HPLWR-Phase 2, contract number 036230. References Bernnat, W., Conti, A., 2009. 2D and 3D assembly burnup analysis for HPLWR. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 40).
Brandauer, M., Schlagenhaufer, M., Schulenberg, T., 2009. Steam cycle optimization for the HPLWR. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 36). Chandra, L., Lycklama à Nijeholt, J.A., Visser, D.C., Roelofs, F., 2009. CFD analyses on the influence of wire wrap spacers on the heat transfer at supercritical conditions. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 30). Dobashi, K., Oka, Y., Koshizuka, S., 1998. Conceptual design of a high temperature power reactor cooled and moderated by supercritical light water. In: Proc. ICONE-6, San Diego, USA, May 10–15. Fischer, K., Schneider, T., Redon, T., Schulenberg, T., Starflinger, J., 2007. Mechanical design of core components for a high performance light water reactor with a three pass core. In: Proc. of GLOBAL 07, Boise, ID, USA, September 9–13. Herbell, H., Himmel, S., Schulenberg, T., 2008. Mechanical analysis of an innovative assembly box with honeycomb structures designed for a high performance light water reactor. In: Proc. of IYNC 2008, Interlaken, Switzerland, 20–26 September 2008 (paper no. 132). Himmel, S.R., Class, A.G., Laurien, E., Schulenberg, T., 2008. Sub-channel analysis of a HPLWR fuel assembly with STAR-CD. In: 16th Pacific Basin Nuclear Conference (16PBNC), Aomori, Japan, October 13–18, 2008 (paper P16P1152). Hofmeister, J., Waata, C., Starflinger, J., Schulenberg, T., Laurien, E., 2007. Fuel assembly design study for a reactor with supercritical water. Nuclear Engineering and Design 237, 1513–1521. Ishiwatari, Y., Yamakawa, M., Oka, Y., Ikejiri, S., 2008. Research and development of a super fast reactor (1) overview and high-temperature structural design. In: 16th Pacific Basin Nuclear Conference (16PBNC), Aomori, Japan, October 13–18, 2008 (paper P16P1290). Kiss, A., Laurien, E., Aszódi, A., Zhu, Y., 2009. Improved numerical simulation of a HPLWR fuel assembly flow with wrapped wire spacers. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 03). Koehly, C., Schulenberg, T., Starflinger, J., 2009. HPLWR reactor design concept. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 37). Kunik, C., Vogt, B., Schulenberg, T., 2009. Flow phenomena in the gap volume between assembly boxes. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 04). Maráczy, C., Hegyi, G., Hordósy, G., Temesvári, E., Hegedös, C., Molnár, A., 2008. High performance light water reactor core design studies. In: 16th Pacific Basin Nuclear Conference (16PBNC), Aomori, Japan, October 13–18, 2008 (paper P16P1221). Maráczy, C., Hegyi, G., Hordósy, G., Temesvári, E., Molnár, A., 2009. HPLWR core design with upward gap flow. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 45). Monti, L., Gabrielli, F., Schulenberg, T., 2009a. Preliminary 3D burn-up analysis of the HPLWR core. In: Annual Meeting on Nuclear Technology, Dresden, Germany, May 12–14, 2009. Monti, L., Schulenberg, T., 2009. Coupled ERANOS/TRACE system for HPLWR 3 pass core analyses. In: International Conference on Mathematics, Computational Methods & Reactor Physics (M&C 2009), May 3–7, 2009. Saratoga Springs, New York. Monti, L., Starflinger, J., Schulenberg, T., 2009b. Effectiveness of an insulated fuel assembly to improve HPLWR core performance. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 46). Schlagenhaufer, M., Starflinger, J., Schulenberg, T., 2009. Steam cycle analyses and control of the HPLWR plant. In: 4thInternational Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 38). Schlagenhaufer, M., Vogt, B., Schulenberg, T., 2007. Reactivity control mechanisms for a HPLWR fuel assembly. In: Proc. of GLOBAL 07, Boise, ID, USA, September 9–13. Schulenberg, T., Starflinger, J., Heinecke, J., 2008. Three pass core design proposal for a high performance light water reactor. Progress in Nuclear Energy 50, p.526–531. Starflinger, J., Schulenberg, T., Marsault, P., Bittermann, D., Maraczy, C., Laurien, E., Lycklama, J.A., Anglart, H., Aksan, N., Ruzickova, M., Heikinheimo, L., 2007. European research activities within the project: high performance light water reactor phase 2 (HPLWR phase 2). In: Proc. of ICAPP’07, Nice, France, May 13–18, 2007 (paper 7146). Starflinger, J., Schulenberg, T., Marsault, P., Bittermann, D., Maraczy, C., Laurien, E., Lycklama, J.A., Anglart, H., Andreani, M., Ruzickova, M., Heikinheimo, L., 2009. Results of the mid-term assessment of the high performance light water reactor phase 2 project. In: Proc. of ICAPP’09, Tokyo, Japan, May 10–14, 2009 (paper 9268). Waata, C., Schulenberg, T., Cheng, X., Laurien, E., 2005. Coupling of MCNP with a subchannel code for analysis of a HPLWR fuel assembly. In: NURETH-11, Avignon, France, October 2–6, 2005 (paper 021). Wank, A., Schulenberg, T., Class, A.G., 2008. Coolant mixing in the plenum of the HPLWR three pass core. In: Proc. of ICAPP’08, Anaheim, CA USA, June 8–12, 2008 (paper 8100).
Contents lists available at ScienceDirect
Nuclear Engineering and Design journal homepage: www.elsevier.com/locate/nucengdes
Design and analysis of a thermal core for a high performance light water reactor T. Schulenberg a,∗ , C. Maráczy b , J. Heinecke c , W. Bernnat d a
Karlsruhe Institute of Technology, Institute for Nuclear and Energy Technologies (IKET), D-76344 Eggenstein-Leopoldshafen, Germany Hungarian Academy of Sciences (KFKI), Atomic Energy Research Institute, H-1525 Budapest, Hungary AREVA NP GmbH, D-91058 Erlangen, Germany d University of Stuttgart, Institute for Nuclear and Energy Systems (IKE), Pfaffenwaldring 31, D-70569 Stuttgart, Germany b c
a r t i c l e
i n f o
Article history: Received 10 March 2010 Received in revised form 1 June 2010 Accepted 2 June 2010
a b s t r a c t The High Performance Light Water Reactor is a Generation IV light water reactor concept, operated at a supercritical pressure of 25 MPa with a core outlet temperature of 500 ◦ C. A thermal core design for this reactor has been worked out by a consortium of Euratom member states within the 6th European Framework Program. Aiming at peak cladding temperatures of less than 630 ◦ C, including uncertainties and allowances for operation, the coolant is heated up in three steps with intermediate coolant mixing to eliminate hot streaks. Different from conventional reactors, the radial power profile is intended to be non-uniform, with the highest power in the first heat-up step in the core center and the lowest power in the second superheater step to result in the same peak cladding temperatures in each region. The concept has been studied with neutronic, thermal-hydraulic and structural analyses to assess its feasibility. Coupled neutronic/thermal-hydraulic analyses are defining the initial distribution of enrichment, control rod positions and the use of burnable poisons. Sub-channel analyses predict the coolant mixing inside assemblies, and a porous media approach simulates the flow of moderator water between assembly boxes. Finally, structural analyses of the assembly boxes are needed to minimize deformations during operation. Even though the core design cannot yet considered to be final, this state of the art review shall summarize the progress achieved so far and outline the remaining challenges. © 2010 Elsevier B.V. All rights reserved.
1. Introduction In 2006, a consortium of 10 European partners from 8 European member states, supported by the European Commission in their 6th Framework Program, decided to work out a design concept of a supercritical water cooled reactor, which they called the High Performance Light Water Reactor (HPLWR). Starflinger et al. (2007) summarize the envisaged project plan and the time schedule for the 3.5 years duration of the second phase of this project. Compared with existing light water reactors, the advantages of this concept shall be lower plant erection costs, due to a once through steam cycle without the need of recirculation pumps, steam separators and dryers, high turbine enthalpies and thus smaller mass flow rates at a given turbine power, a higher thermal efficiency, and a compact containment design, all leading primarily to economic advantages. Basically, the core of such a HPLWR can have a thermal or fast neutron spectrum. During the course of the project, however, being restricted by manpower, the consortium decided to concentrate on a thermal core design, leaving the fast core studies to Ishiwatari et al. (2008) and thus to gain synergies in the Generation
∗ Corresponding author. Tel.: +49 7247 82 3450; fax: +49 7247 82 4837. E-mail address: [email protected] (T. Schulenberg). 0029-5493/$ – see front matter © 2010 Elsevier B.V. All rights reserved. doi:10.1016/j.nucengdes.2010.09.025
IV International Forum. Meanwhile, the project has passed successfully its midterm assessment (Starflinger et al., 2009), design details have been worked out and first analyses are available to get a first impression about the feasibility of the core concept. 1.1. Core design target Aiming at a net electric power of around 1000 MW and a net efficiency of almost 44%, the target thermal power of the reactor core needs to be 2300 MW, confirmed also by recent steam cycle analyses of Brandauer et al. (2009). Early cycle studies by Dobashi et al. (1998) indicated an optimum thermal efficiency at a feedwater temperature of 280 ◦ C which was kept also for the present study. The target core outlet temperature was chosen as 500 ◦ C which is still rather low for a once through steam cycle with single reheat, compared with latest fossil fired power plants, but appears to be challenging enough with regard to available fuel cladding materials. Their peak temperature limit was targeted at 630 ◦ C which is not only a challenge for oxidation and corrosion protection, but also for their creep strength and resistance to stress corrosion cracking. The fuel centerline temperature is a function of the linear power of the fuel rod. It has been limited to 39 kW/m under nominal conditions. To be competitive with respect to latest pressurized water reactors, the target burn up should be at least 45 GWd/tHM . Near future
T. Schulenberg et al. / Nuclear Engineering and Design 241 (2011) 4420–4426
4421
3. Mechanical design
Fig. 1. Core design concept with evaporator, superheater 1 and superheater 2 assembly clusters (shown exemplarily from right to left) (Koehly et al., 2009).
fuel enrichment plants for light water reactors are expected to be limited by 6% enrichment, which was chosen as the enrichment target accordingly. Like with boiling water reactors, boron acid cannot be used to compensate the excess reactivity at the beginning of a burn-up cycle, so that burnable absorbers must be used instead. The target power and temperatures result in a coolant mass flow rate of 1179 kg/s. Schlagenhaufer et al. (2009) suggest a feedwater pressure of 25 MPa for all load conditions which keeps some margin from the critical pressure of 22.1 MPa.
2. Core layout These target data differ from conventional light water reactors not only by the higher pressure and core outlet temperature, but also by a significantly higher enthalpy rise in the core. Indeed, the difference between life steam enthalpy and feedwater enthalpy of 1936 kJ/kg exceeds the one of pressurized water reactors by around a factor of 8. Assuming an overall hot channel factor of 2 between the peak and the average coolant heat-up, this enthalpy rise would result in peak coolant temperatures of 1200 ◦ C which is far beyond the target temperature limit. A strategy to overcome this issue can be learned from fossil fired boiler design. These boilers are characterized by multiple heat-up steps with intensive coolant mixing between them to eliminate hot streaks. Schulenberg et al. (2008) applied such a strategy for a core layout with a first heat-up of the coolant as moderator water, comparable with the economizer of a fossil fired boiler. The second heat-up should be in the evaporator assemblies in the center of the core, followed by coolant mixing in a plenum above the core. From there, the coolant is directed downwards in assemblies of the first superheater, surrounding the evaporator, to be mixed again in an annular chamber underneath the core. Final heat-up to the envisaged core outlet temperature of 500 ◦ C was proposed to happen in a second superheater stage with upward flow again in assemblies at the core periphery. The core design concept is shown in Fig. 1. Assuming a hot channel factor of 2 for each heat-up step, as an initial guess, the power ratio of evaporator to superheater 1 to superheater 2 should be around 4:2:1 to reach the same peak coolant temperature in each region. The proposed core layout is trying to reach this power ratio by placing the second superheater at the core periphery where the neutron leakage is reducing the neutron flux anyway.
Beyond 390 ◦ C, the coolant density is less than 200 kg/m3 , hardly enough to produce a thermal neutron spectrum. Therefore, colder feedwater is foreseen as moderator water to run inside moderator boxes in the fuel assemblies and in gaps between assembly boxes. As Schulenberg et al. (2008) predicted a pressure drop of about 0.5 MPa from core inlet to its outlet, but the mass of structural material in the core should be minimized to limit the neutron absorption, Hofmeister et al. (2007) concluded that the fuel assemblies should be small, preferably with 40 fuel pins each and a single moderator box in their center to enable a small wall thickness of moderator and assembly boxes. To ease handling during maintenance, they recommended to group 9 assemblies to a cluster with common head and foot piece. Fischer et al. (2007) adopted this idea to design a fuel assembly cluster for the three pass core as described above, preferably such that clusters can be exchanged between evaporator and superheater positions. Wire wraps were proposed as grid spacers to improve coolant mixing in both flow directions. The clusters can be disassembled at their foot piece to exchange single fuel rods for repair. Control rods shall be inserted from the top of the core. They run inside 5 of the 9 moderator boxes. Different control rod designs, e.g. cruciform rods and square tubes, and their shut down reactivities have been studied by Schlagenhaufer et al. (2007). A thermal insulation of assembly and moderator boxes has been proposed by Herbell et al. (2008) using a sandwich design with internal honeycomb structures to reach the envisaged stiffness of less than 0.5 mm deflection towards the fuel rods. With a pitch to diameter ratio of only 1.18, the fuel rod bundle is rather tight, which results in a mass flux of more than 1600 kg/m2 s for a nominal fuel assembly. A lower mass flux might cause concern with respect to the deterioration of heat transfer in the evaporator. The envisaged fuel rod diameter of 8 mm leads to an average core power density of about 60 MW/m3 . Mixing chambers above and underneath the core have been designed by Fischer et al. (2007) and updated by Koehly et al. (2009) to enable a downward flow of moderator water inside the moderator boxes, but an upward flow in the gaps between the assembly boxes. The latest design uses 50% of the feedwater to run at first downwards through the moderator boxes, then upwards between the assembly boxes, and then downwards again to cool the reflector around the core. The remaining fraction of feedwater is supplied through the downcomer to be mixed with the moderator water, released from the reflector, in a mixing volume underneath the core at the evaporator inlet. The upper mixing chamber is aligned near the steam outlet flanges of the reactor pressure vessel. It is separated internally by walls to serve also as the steam plenum above the core outlet. C-ring sealings avoid feedwater ingress into the plenum. Free thermal expansion of the assembly clusters is allowed at their upper ends, where the head pieces are sealed with c-rings at their penetrations through the upper chamber, whereas the foot pieces are standing on the core support plate. This core design concept is certainly quite unusual. It cannot be compared with any other core design built and tested in the past. Detailed analyses are being performed, therefore, to assess if such a three pass core will be feasible. The following chapters will summarize the current status of neutronic, thermal-hydraulic and structural analyses.
4. Analyses of the three pass core concept 4.1. Power distribution The thermal core is characterized by a large change of coolant density, ranging from more than 700 kg/m3 at the core inlet to
4422
T. Schulenberg et al. / Nuclear Engineering and Design 241 (2011) 4420–4426
Fig. 2. Core power distribution in case of fresh fuel with 5% uniform enrichment (Monti et al., 2009a).
Fig. 3. Core power distribution after 200 days burn up (Monti et al., 2009a).
less than 90 kg/m3 at its outlet. Even though additional moderator water has been foreseen to compensate missing coolant density, the local core power is changing significantly with coolant heat-up. As a consequence, neutronic codes need to be coupled with thermalhydraulic codes to account for this effect. Waata et al. (2005) were
the first who coupled the Monte Carlo code MCNP with the subchannel code STAFAS by running them iteratively to obtain the axial power profile of a single assembly of the HPLWR. Meanwhile, the method has been extended to predict the power distribution of the entire core.
Fig. 4. Radial core power distribution, relative numbers, axially averaged (Maráczy et al., 2008).
T. Schulenberg et al. / Nuclear Engineering and Design 241 (2011) 4420–4426
4423
Fig. 5. Peaking factors of the core power in each core region (Maráczy et al., 2008).
Monti and Schulenberg (2009) coupled the neutron transport code ERANOS with the system code TRACE to predict the power distribution of a quarter of the core made from fresh fuel with uniform enrichment. Maráczy et al. (2008) coupled the diffusion code KARATE with the one-dimensional thermal-hydraulic code SPROD to optimize the core power distribution with the help of control rods and burnable poisons. Bernnat and Conti (2009) coupled MCNP5 and STAFAS with ORIGEN to predict burn-up effects of a single assembly and even of a sector of the core. The physical mechanisms causing the power distribution of this strange core design can be understood when we look at the hypothetical case of fresh fuel with 5% uniform enrichment. Monti et al. (2009a) show under this assumption that the power ratio of the individual core regions (evaporator:superheater 1:superheater 2) is rather like 6:2.5:1, meaning that the evaporator is producing even more power than envisaged. This power ratio decreases quickly within 200 days of burn-up to about 3.3:2:1. The comparison is shown in Figs. 2 and 3 for 1/4 of the core. Here, the finest resolution of the power distribution is a fuel assembly cluster. In these figures, 13 of such clusters form the evaporator, the first and the second superheater each. The general strategy to reach the envisaged power ratio of 4:2:1 is therefore to add Gd absorber rods, controls rod or older fuel assemblies to the evaporator. Maráczy et al. (2009), Fig. 4, show a first example of a core power distribution in which the central power peak is suppressed by Gd rods and partially inserted control rods. Here, the finest resolution of the power distribution is a single assembly. 3 × 3 assemblies form a cluster. The evaporator
region and both superheater regions are separated by solid lines. The position of inserted control rods is visible in the evaporator region as local, cross shaped low power regions. These are at cluster positions (1;5), (2;2), (2;4), (3;4), (4,2), (4;3) and (5;1) counted from the core symmetry lines in x- and y-direction, respectively. Once the envisaged power ratio has been reached, the power profile needs to be homogenized in each region to minimize the peak power. We see in particular in the superheater 2 region of Fig. 4 that the radial flux gradient causes hotter assemblies towards the first superheater and colder ones towards the outer reflector. Therefore, a further optimization of the power profile by Maráczy et al. (2009) includes a water layer in the reflector and uses different enrichments to simulate different burn-up stages which are substituted by fresh fuel for this analysis. Fig. 5 shows the peaking factors, i.e. the assembly power normalized with the average power of each core region. Now, the hottest power peak in the second superheater is reduced to 1.35× the average power there. Further design studies are aiming at an optimization of the burnup cycle. Monti et al. (2009b) discuss the importance of having thermally insulated assembly and moderator box walls to keep a high density of the moderator water. Comparing solid steel boxes with insulated boxes, they got an increase of the initial excess reactivity of 8600 pcm by the thermal insulation proposed by Herbell et al. (2008). Bernnat and Conti (2009) give first indications how the burn-up can be increased by shuffling assembly clusters from the second superheater, where some Pu is produced by the harder neutron spectrum, to evaporator cluster positions. A detailed look into the power distribution inside a fuel assembly is indicating that Gd
4424
T. Schulenberg et al. / Nuclear Engineering and Design 241 (2011) 4420–4426
Fig. 7. Coolant temperature profile of the hottest superheater assembly with 20% assumed radial power gradient (Himmel et al., 2008).
Fig. 6. Flow reversal in the gap volume between the assembly boxes if moderator water is supplied from the top (Kunik et al., 2009).
rods should preferably be positioned in the corners of the assemblies. Moreover, their sensitivity study of the influence of assembly box bending on the power distribution inside assemblies will help to quantify hot channel factors based on uncertainties later on. 4.2. Flow structures and coolant mixing While the interaction of the coolant density inside assemblies with the power profile can be modeled with one-dimensional thermal-hydraulic codes, the flow structure in the gaps between assembly boxes and its density distribution need to be modeled at least in two dimensions. Kunik et al. (2009) approximated this flow structure with a porous media approach, assuming that the moderator water in the gaps was supplied from the top. Despite the thermal insulation of the assembly boxes, they conclude that a change of the flow pattern from purely downward to local recirculation may occur, caused by buoyancy of hot zones around the second superheater as shown in Fig. 6. Such change in the flow pattern, however, can effect the power profile which, in turn, can influence the flow structure again. As a consequence, to avoid instabilities, the flow path of the moderator water was changed to a supply through the foot pieces of the assembly clusters and thus to obtain a stable upward flow from there. Excellent coolant mixing inside assemblies and in the mixing chambers above and underneath the core is of crucial importance for this reactor concept. It has been the intent of the wire wrapped fuel rods to serve as mixing devices inside assemblies. The resultant coolant temperature distribution has been checked with the sub-channel code C3CLM, an AREVA NP version of the well known sub-channel code COBRA-IIIC, for worst case power gradients of superheater assemblies. Himmel et al. (2008) validated this subchannel analysis with CFD analyses using locally a fine mesh and globally a coarse mesh approach. As an example, we show in Fig. 7 the coolant temperature distribution of the hottest superheater 2
assembly with a linear power gradient causing 20% more power in the upper row of fuel rods compared with the average power of each cross section. The maximum coolant temperature at the outlet is just 16 ◦ C hotter than the average coolant temperature there. Full CFD analyses of Kiss et al. (2009) confirm this excellent result. Coolant mixing in the upper mixing chamber between the evaporator outlet and the first superheater inlet has been studied numerically by Wank et al. (2008). The rather blocked flow path, running around obstacles like head piece structures, moderator boxes and connection tubes of the mixing chamber made it quite difficult to mix the coolant as intended. Several design modification were tried before a convincing mixing structure could be identified. The complexity of the flow path inside the upper mixing chamber is illustrated in Fig. 8. Another cause for concern is the risk of a deterioration of heat transfer in the evaporator assemblies close to the pseudo-critical point. The phenomenon is well known from boiler tubes where it can cause hot spots in case of high heat flux and low mass flux. First CFD analyses by Chandra et al. (2009) indicate, however, that wires wrapped around the fuel rods help to prevent such effects by increasing the turbulence level. Local flow separation downstream
Fig. 8. Two head pieces shown exemplarily in the upper mixing chamber (Wank et al., 2008).
T. Schulenberg et al. / Nuclear Engineering and Design 241 (2011) 4420–4426
4425
Fig. 9. Design of the assembly and of the moderator box with honeycomb panels and location of spacer pads (Herbell et al., 2008).
of these wires, on the other hand, can cause another hot spot which will need to be examined carefully and should be included in the final hot spot evaluation. 4.3. Stresses and deformations As discussed above, the assembly and moderator boxes need to be thermally insulated to minimize the heat-up of moderator water. Herbell et al. (2008) propose a sandwich design, shown in Fig. 9, with an internal honeycomb structure to provide sufficient stiffness and to stand the pressure difference between the outside and the inside of these boxes. The honeycombs are filled with alumina or zirconia for thermal insulation. They need to be vented to the colder side as they cannot stand an outside pressure of 25 MPa. Deformations and stresses have been studied by Herbell et al. (2008) with Finite Element Analyses of an axial segment of the assembly box. As a result, massive
corner pieces, unfortunately with higher thermal conductivity, were required to keep the local peak stresses below the fatigue limit. Radial power gradients do not only cause a temperature difference of the fuel rods inside an assembly. As the assembly box is thermally insulated, it will experience similar temperature differences as the coolant, forcing it to bend to the hotter side. As a consequence, the gap between the fuel assemblies is reduced or increased from its nominal gap width of 9 mm, may be even until it contacts the neighboring assembly box. This effect influences the neutron moderation and, in turn, the power profile of the assembly. In order to avoid such bowing, two sets of spacer pads, attached to the corners of the box, have been foreseen between the assembly boxes as shown in Fig. 9 on the right hand side. The worst case considered to study the residual bowing line assumes that one side of an evaporator assembly box experiences up to 80 ◦ C hotter temperatures than its opposite side. The bowing analysis shown in Fig. 10
4426
T. Schulenberg et al. / Nuclear Engineering and Design 241 (2011) 4420–4426
Fig. 10. Bowing of the assembly box assuming up to 80 ◦ C temperature difference between the hotter and colder side.
demonstrates that even in this case the maximum box deflection is less than 0.5 mm. 5. Conclusions The thermal core design of the HPLWR, though being highly innovative and anything but easy with its different heat-up regions and its upward and downward flow path, turned out to be feasible in all analyses performed up to now. Thanks to the repeated mixing, it has sufficient margin from maximum cladding temperature limits, which have not yet been reached by any of these predictions. The wire wrapped fuel rods are expected to help not only to improve mixing inside assemblies but also to avoid a deterioration of heat transfer close to the pseudo-critical point of water. At lot more needs to be analyzed, however, before the core design can finally be assessed. The peak cladding temperatures of individual fuel rods need to be predicted taking into account different burn up stages, control rod positions, all kinds of tolerances and uncertainties as well as allowances for operation. The latter ones are just being optimized by dynamic steam cycle analyses of Schlagenhaufer et al. (2009). Accidental conditions need to be studied, not only to design the safety system but also to predict the fuel and cladding temperatures under transient and abnormal conditions. In particular, the consequences of natural convection phenomena in heated downward flow regions should be examined with care. In so far, the design is not yet considered to be final but still subject to further improvements. Acknowledgement This work has been funded by the European Commission as part of their project HPLWR-Phase 2, contract number 036230. References Bernnat, W., Conti, A., 2009. 2D and 3D assembly burnup analysis for HPLWR. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 40).
Brandauer, M., Schlagenhaufer, M., Schulenberg, T., 2009. Steam cycle optimization for the HPLWR. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 36). Chandra, L., Lycklama à Nijeholt, J.A., Visser, D.C., Roelofs, F., 2009. CFD analyses on the influence of wire wrap spacers on the heat transfer at supercritical conditions. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 30). Dobashi, K., Oka, Y., Koshizuka, S., 1998. Conceptual design of a high temperature power reactor cooled and moderated by supercritical light water. In: Proc. ICONE-6, San Diego, USA, May 10–15. Fischer, K., Schneider, T., Redon, T., Schulenberg, T., Starflinger, J., 2007. Mechanical design of core components for a high performance light water reactor with a three pass core. In: Proc. of GLOBAL 07, Boise, ID, USA, September 9–13. Herbell, H., Himmel, S., Schulenberg, T., 2008. Mechanical analysis of an innovative assembly box with honeycomb structures designed for a high performance light water reactor. In: Proc. of IYNC 2008, Interlaken, Switzerland, 20–26 September 2008 (paper no. 132). Himmel, S.R., Class, A.G., Laurien, E., Schulenberg, T., 2008. Sub-channel analysis of a HPLWR fuel assembly with STAR-CD. In: 16th Pacific Basin Nuclear Conference (16PBNC), Aomori, Japan, October 13–18, 2008 (paper P16P1152). Hofmeister, J., Waata, C., Starflinger, J., Schulenberg, T., Laurien, E., 2007. Fuel assembly design study for a reactor with supercritical water. Nuclear Engineering and Design 237, 1513–1521. Ishiwatari, Y., Yamakawa, M., Oka, Y., Ikejiri, S., 2008. Research and development of a super fast reactor (1) overview and high-temperature structural design. In: 16th Pacific Basin Nuclear Conference (16PBNC), Aomori, Japan, October 13–18, 2008 (paper P16P1290). Kiss, A., Laurien, E., Aszódi, A., Zhu, Y., 2009. Improved numerical simulation of a HPLWR fuel assembly flow with wrapped wire spacers. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 03). Koehly, C., Schulenberg, T., Starflinger, J., 2009. HPLWR reactor design concept. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 37). Kunik, C., Vogt, B., Schulenberg, T., 2009. Flow phenomena in the gap volume between assembly boxes. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 04). Maráczy, C., Hegyi, G., Hordósy, G., Temesvári, E., Hegedös, C., Molnár, A., 2008. High performance light water reactor core design studies. In: 16th Pacific Basin Nuclear Conference (16PBNC), Aomori, Japan, October 13–18, 2008 (paper P16P1221). Maráczy, C., Hegyi, G., Hordósy, G., Temesvári, E., Molnár, A., 2009. HPLWR core design with upward gap flow. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 45). Monti, L., Gabrielli, F., Schulenberg, T., 2009a. Preliminary 3D burn-up analysis of the HPLWR core. In: Annual Meeting on Nuclear Technology, Dresden, Germany, May 12–14, 2009. Monti, L., Schulenberg, T., 2009. Coupled ERANOS/TRACE system for HPLWR 3 pass core analyses. In: International Conference on Mathematics, Computational Methods & Reactor Physics (M&C 2009), May 3–7, 2009. Saratoga Springs, New York. Monti, L., Starflinger, J., Schulenberg, T., 2009b. Effectiveness of an insulated fuel assembly to improve HPLWR core performance. In: 4th International Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 46). Schlagenhaufer, M., Starflinger, J., Schulenberg, T., 2009. Steam cycle analyses and control of the HPLWR plant. In: 4thInternational Symposium on Supercritical Water-Cooled Reactors, March 8–11, 2009, Heidelberg, Germany (paper no. 38). Schlagenhaufer, M., Vogt, B., Schulenberg, T., 2007. Reactivity control mechanisms for a HPLWR fuel assembly. In: Proc. of GLOBAL 07, Boise, ID, USA, September 9–13. Schulenberg, T., Starflinger, J., Heinecke, J., 2008. Three pass core design proposal for a high performance light water reactor. Progress in Nuclear Energy 50, p.526–531. Starflinger, J., Schulenberg, T., Marsault, P., Bittermann, D., Maraczy, C., Laurien, E., Lycklama, J.A., Anglart, H., Aksan, N., Ruzickova, M., Heikinheimo, L., 2007. European research activities within the project: high performance light water reactor phase 2 (HPLWR phase 2). In: Proc. of ICAPP’07, Nice, France, May 13–18, 2007 (paper 7146). Starflinger, J., Schulenberg, T., Marsault, P., Bittermann, D., Maraczy, C., Laurien, E., Lycklama, J.A., Anglart, H., Andreani, M., Ruzickova, M., Heikinheimo, L., 2009. Results of the mid-term assessment of the high performance light water reactor phase 2 project. In: Proc. of ICAPP’09, Tokyo, Japan, May 10–14, 2009 (paper 9268). Waata, C., Schulenberg, T., Cheng, X., Laurien, E., 2005. Coupling of MCNP with a subchannel code for analysis of a HPLWR fuel assembly. In: NURETH-11, Avignon, France, October 2–6, 2005 (paper 021). Wank, A., Schulenberg, T., Class, A.G., 2008. Coolant mixing in the plenum of the HPLWR three pass core. In: Proc. of ICAPP’08, Anaheim, CA USA, June 8–12, 2008 (paper 8100).