keep the temperatures within acceptable limits in the long term. However, in the SEAFP reference plant design, the presence of a concrete cryostat 2 m thick.
Fusion Engineering and Design
Fusion Engineering and Design 29 (1995) 193- 197
SEAFP: passive removal of the decay heat F. Andritsos
‘, A. Angelini ‘, H.W. Bartels b, W. Daenner ’ Europeun Commission, Joint Research Centre, b Max-Planck-Institut firPlasma Physik, ’ The NET Team, Max-Planck-Institut ftir Plasma ’ Dipartimento di Energetica, Politecnico di
‘, M. Zucchetti’
I-21020 I.spra (Va), Italy Gurching, Germany Physik, Garching, Germany Torino, Torino, Italy
Abstract Passive safety is a key issue for the public acceptance of fusion power. Several studies are under way in the European Safety and Environmental Assessment of Fusion Power (SEAFP) program to analyze the possible hazards and the means to keep the system safe. The demonstration that the decay heat of the structural materials of the reactor will not cause the failure of the containment, even in the absence of any form of active cooling, is a significant contribution towards the demonstration of the passive safety of fusion.
1. Introduction One of the worst hypothetical accidents for a fusion reactor is that of losing totally the cooling capacity of the complete machine. This accident, also referred to as a total loss of coolant accident (LOCA), has indeed been indicated, by previous calculations [l-4], performed for past NET-ITER CDA designs, as potentially the most hazardous LOCA. It implies an instantaneous, complete, total loss of any cooling mechanism, other than conduction and radiation. This hypothetical accident has been chosen as an “envelope” case for the temperature transients of the whole device after a LOCA or LOFA. In the past, it has been modelled and analyzed extensively by Andritos at the Joint Research Centre (Ispra site) of the Commission of the European Communities, in the frame of the European Fusion Technology program. Calculations for the same purpose, but with a different code and with a different approach, were Elsevier Science S.A. SSDlO920-3796(94)00273-S
carried out also at IPP Garching by Konrad [7]. Results from these two models, although conceptually different, were very close to each other, thus giving a high degree of confidence in the applied codes and methods. In contrast to past studies investigating decay heat removal in next-step machines such as ITER or NET, this paper deals with purely passive decay heat removal in future fusion reactors, assuming continuous operation at a neutron wall loading of about 2 MW rne2. Three reactor designs are studied [ 81: (1) the reference design based on a helium-cooled V/LiO, breeder blanket; (2) the alternative design based on a water-cooled steel/LiPb breeder blanket; (3) the “preliminary” design based on a watercooled steel/Li,ZrO, breeder blanket. In the past “envelope” post-accidental transient simulations [2-71, a constant natural convection coefficient of 10 W Km’ m-l was assumed at the external surface
F. Andrifsos et al. / Fusion Engineering and Design 29 (1995) 193-197
194 Table 1 Global decay heat SEAFP models Time
OS 1 min 10 min Ih 6h 1 day
3 days 10 days 1 month I year
power in the reference and alternative
Decay heat (MW) for V-STi/Li,O/Be reference
48 43 10 7.8 5.7 3.7 1.8 1.3 1.0 0.4
Decay heat (MW)
Table 2 Time-integrated global decay heat in the reference tive SEAFP models Time
Decay heat (GJ) for WSTi/Li,O/Be reference
lh 1 day I week 1 month
44 520 1900 4400
for reduced activation MANET/Li,,Pb,, alternative 17 35 32 16 1.5 4.8 4.3 4.0 3.6 1.4
of the cryostat. For a steel cryostat, such as that for the CDA ITER, this boundary condition was enough to keep the temperatures within acceptable limits in the long term. However, in the SEAFP reference plant design, the presence of a concrete cryostat 2 m thick made the system practically adiabatic [9]. To have a reasonable chance of keeping the “envelope” temperatures within acceptable limits, a modification of the cryostat is proposed, introducing the concent of the ventilated cryostat. A gap of about 10 cm is introduced between the air-tight steel liner and the concrete wall 2 m thick, which serves as a biological shield. In this way, turbulent natural convection ensures the passive dissipation of the decay heat, while maintaining both the air tightness and the shielding effect of the concrete wall.
2. Calculation of global decay beat Neutron fluxes, calculated with the ANISN code, are fed into the FISPACT-3 code which, using the EAF-3 cross-section library, give the specific decay heat histories at each zone of the reactor [ 121. The mean inboard and outboard neutron wall loads were 1.65 and 2.4 MW mm*, respectively, for 5 full power years of continuous irradiation. The global SEAFP decay heat is calculated as follows. The specific decay densities (in watts per cubic centimetre) are integrated along the radial direction of the reactor midplane. This yields the total decay heats of the portions of the device which
and alterna-
Decay heat (GJ) for reduced activation MANET/Li,,Pb,, alternative 93 700 3000
11000
are found behind plasma chamber unit areas in the inboard and outboard regions respectively. Multiplication of the inboard and outboard areas determines the global SEAFP decay heat numbers listed in Table 1 for the reference and alternative reactor models. The “preliminary” SEAFP reactor model has decay heat values which are about 20% below those of the alternative model, since the same less structural materials, though of the same type, are used for this particular model. In Table 2, the time-integrated release decay heat energies are listed. It turns out that the decay heat is exclusively dominated by the structural materials. The only exception is a significant contribution of LiPb in the first few seconds after plasma shutdown, but this has no effect on the safety of the fusion plant. Since the breeding blankets exhibit poor shielding of the neutrons, the decay heat in the neutron shield behind the breeding part of the blanket plays a significant role. For the reference design based on a vanadium alloy for the breeding blanket, the steel-based shield dominates the decay heat for times beyond a few days after the accident. Therefore, the choice of the steel in the shielding part of the blanket is of importance. The numbers listed in Table 1 assumed an SS3 16 shield. If a manganese-based reduced activation steel is assumed, then the global decay heat powers in the time span between 1 day and 1 month are about a factor of 2 larger for the reference design, and about 25% larger for the alternative design.
3. Evaluation of passive cooling capabilities A thick concrete cryostat, such as that in the SEAFP under envelope post-accidental conditions, design, causes quasi-adiabatic conditions for a very long time after the incident [9]. To remedy this, the design of the
F. Andritsos
et al. 1 Fusion
Engineering
and Design
29 (1995)
193-197
195
Air I
Air
_ T
Air
Fig. I Schematic representation of the cooling of the cryostat liner in the SEAFP plant.
-300
1 min I
10’ cryostat was modified by Andritsos et al. [IO]: the steel liner which covers completely the internal surface of the cryostat was detached by several centimetres from the concrete wall. Air, driven by gravity through openings at the top and bottom of the concrete wall, provides natural cooling for the liner, while maintaining the tightness of the volume inside the cryostat. The system is shown schematically in Fig. 1. The convection coefficient at the liner has been calculated, taking into account a maximum temperature of 260 “C for the liner, while the air enters at 20 “C and leaves at a maximum of 200 “C. The thermal power to be evacuated in the long run (month timescale) is assumed to be about 3 MW. For a total heat the heat load is exchange area of 3000 m’, 0.7 kW mm*. At these temperatures, for a gap 10 cm wide, we have the Prandtl number Pr = 0.73 and the Grashof number Gr = 1.73 x 1014. Since Gr x Pr = 1.27 x 10“’ > 109, the flow will be turbulent with an air velocity of about 1 m SK’, yielding a Nusselt number of Nu = 0.129( Gr x Pr) ‘I3 = 6482 Thus we have a convection coefficient of CI= 6.77 W m* “C’. This z value has been used as a reference in the following calculation of the temperature transients. It yields a cooling power of 3 MW. As mentioned by Vossebrecker of Siemens AG ,[131, a more detailed analysis might allow the application of heat transfer correlations for forced convection driven by gravity. The heat removed under these assumptions is about a factor of 3 larger than that given above, which adds considerable margin to our discussion.
,,,,,.,I
1 hr .,.,,,,I
lo2
lday
,,,,,,,I
lo3
.,.,,,,
lo4
I
lo5
iwelmo 1
lo6
,,,.
4
10”
time (s) Fig. 2. Envelope post-accidental temperature histories of the inboard components of the SEAFP reference plant model: 1, central post; 2, outboard first wall; 3, blanket shield; 4, vacuum vessel; 5. toroidal magnets; 6, liner of cryostat; 7, concrete wall.
4. Temperature transients The model used was that described in detail by Andritsos [ 141. It is a two-dimensional model of a complete 11.25” midplane sector of the reactor, from the centre to after the cryostat. Thus, it takes into consideration the actual geometry of the reactor at the plane of the magnetic axis, including the spacing between the outboard magnets, openings, etc. All the properties of the materials considered are temperature dependent and the thermal shield of the magnet is degrading, passing from an effective emissivity of 0.05 to 0.5 between 200 and 300 “C. The fact that the highest temperatures manifest themselves at very long timescales, when the heat transfer is quasi-static, permits the modelling of the poloidal heat flux (i.e. mainly through the magnets and the vacuum vessel) with massless equivalent resistances. A natural convection of 6.77 W mm7 “C (see Section 3) is applied to the external part of the liner and the internal part of the concrete wall. The decay heat that serves as an input to the thermal model is calculated separately for each of the zones of the model, as described earlier in Section 2. For the thermal modelling, the P-Thermal finite difference code was used. The Patran CAE package was used to set the model up, verifying it and post-processing the results. Both codes are validated, commercially
F. Andritsos et al. 1 Fusion Engineering and Design 29 (1995) 193-197
196
Table 3 Maximum
“envelope”
Component
First wall Vacuum vessel Magnets Liner Concrete
1 min
1 hr
lday
lwelmo
time (s) Fig. 3. Evelope post-accidental temperature histories of the outboard first wall of the SEAFP reference, alternative and preliminary plants.
available packages. In addition, during the last 3 years, the results of similar models have been tested extens-ively, so as to have a high degree of confidence in the codes and the modelling procedures. Both codes are on an IBM R6000, mod 375 workstation, where about 40 h of CPU time is required for a complete transient of 3 months. The results of the thermal transient can be seen in Fig. 2 for the reference SEAFP model as temperatures of different components plotted against time. The vacuum vessel (AISI 316) peaks 1 week after the accident at around 500 and 550 “C for the inboard and outboard regions, respectively, while the V-5Ti first wall and front part of the blanket peak at 650 “C peak only a few hours after the accident. The maximum temperature of the liner is only around 60 “C, while the concrete remains practically at room temperature. Fig. 3 shows the time evolution of the maximum temperatures in the reference, alternative and preliminary SEAFP designs. As a result of the low activation of the V-STi, and post-accidental temperature increase is limited below 100 K for the reference reactor. The maximum temperatures in the water-cooled steel-based designs are 870 “C (alternative design) and 780 “C (preliminary design).
temperatures
Temperature (“C) for preliminary
Temperature
designs (‘C)
MANET/Li,ZrO,/Be
for reference V-STi/Li,O/Be
800 600 500 260 120
660 550 350 60 25
10 cm) between the liner and the concrete wall of the cryostat, through which turbulent natural convection takes place. In this way, the decay heat of the tokamak is removed, while maintaining the tightness of the cryostat liner. Choosing the dimensions and the calculation of the effective convection coefficient were carried out on the basis of the global decay heat produced in the tokamak 1 month after the accident. This convection coefficient was used in a detailed finite difference mode1 of the whole tokamak to calculate the temperature transient. The results confirmed that the proposed design solution can ensure the passive removal of the decay heat in the SEAFP reference reactor, even in the “envelope” case of a total LOCA. For this particular case, the temperature transient proved to be quite insensitive to the value of the convection coefficient assumed at the liner-concrete gap. The “envelope” maximum temperatures of the various components of the SEAFP reference reactor are given, for comparison purposes, together with those of the preliminary design (MANET/Li,ZrO,/Be), in Table 3. These temperatures, although they will probably cause serious damage, are not expected to compromise the structural stability of the reactor containment. Hence, we have the conclusion that the thermal effects of the decay heat alone cannot compromise the containment of the SEAFP reference reactor, even in the envelope scenario of a complete and permanent loss of every form of active cooling.
References
[II J. Raeder and H. Gorenflo, 5. Conclusions
PI A novel passive cooling system for the SEAFP reference design was proposed. It consists of a gap (about
in the two SEAFP
On NET afterheat under loss of active cooling conditions-revised transient analysis, NET Internal Note NET/88/IN/044, October 1988. F. Andritsos, Thermal modelling of the afterheat after a total LOCA of the ITER, Tech. Note 1.90.66,CEC, JRC, Ispra, September 1990.
F. Andritsos
et al. 1 Fusion Engineering
[3] F. Andritsos, J. Kaouris, P. Rocco, M. Zucchetti et al., Loss of vacuum, loss of coolant and loss of flow accident analysis for NET/ITER, Proc. 16th Symp. on Fusion Technology, London, September 1990, pp. 147771481. [4] F. Andritsos and M. Zucchetti, Afterheat transient in ITER after a total loss of coolant accident, Fusion Eng. Des., 15(2) (1991) 113-119. [5] F. Andritsos, Post LOCA transient of the outboard region of ITER-intermediate report, safety & environment subtask 3.8, Tech. Note 1.91.137, CEC. JRC, Ispra. July 1991. [6] F. Andritsos, Post LOCA temperature transient of ITER inboard and outboard, Tech. Note 1.92.30, CEC, JRC, Ispra, April 1992. [7] C. Konrad, Nachwaermeeffekte aufgrund der Neutronenaktivierung einer Fusionsanlage, Diplomarbeit. Technisthe Universitat, Munich, January 1992. [S] J. Raeder and H.W. Bartels, Reference plant model and alternative plant model for the SEAFP, SEAFP-Memo. N/N/0900/1 7/A, Garching, July 1993.
and Design 29 (1995) 193.-197
197
transients ]91 F. Andritsos, Post LOCA envelope temperature of the SEAFP reference power reactor-ceramic blanket, Tech. Note 1.93.106, CEC, JRC, Ispra, July 1993. [lOI F. Andritsos, A. Angelini and H.W. Bartels, SEAFPpassive removal of the decay heat, SEAFP Rep. I-A2.3/ 2(93), IPP Garching, December 1993. [Ill J.Ch Sublet, Scoping assessments of the activation characteristics of first wall, divertor and blanket materials, SEAFP Memo. SEAFP/A6-5-93, Culham. 1993. [I21M. Zucchetti and P. Rocco, Activations calculations to perform: temperature transient calculations & detailed waste categorization, Tech. Note 1.93.165, CEC, JRC, Ispra, December 1993; also presented at workshop SEAFP-WS. AEA Culham, April 1994. Private communication, Siemens AG, [I31 H. Vossebrecker. November 1993. SEAFP-post LOCA envelope tempera[I41 F. Andritsos, ture transients: reference & alternative plant model, Tech. Note CEC, JRC, Ispra, in press; also presented at workshop SEAFP-WS, AEA Culham, April 1994.
