Contribution to « Nuclear Engineering & Design »
GAS-COOLED FAST REACTORS – STATUS OF CEA PRELIMINARY
DESIGN STUDIES
P. Dumaz, P. Allègre, C. Bassi, T. Cadiou, A. Conti, JC Garnier, J.Y. Malo & A. Tosello
CEA/DEN/DER/SESI, CEA Cadarache, 13108 Saint-Paul-lez-Durance, FRANCE
Phone: +00 (33) 4 42 25 40 98, Fax: +00 (33) 4 42 25 71 87, E-Mail: Patrick.Dumaz@cea.fr
ABSTRACT
The Gas cooled Fast Reactor (GFR) is one of the six reactor concepts selected in the frame of the
Generation IV initiative. The most significant GFR option is the use of a helium high temperature
primary coolant. The helium option is very attractive (chemical inertness, neutron transparency, etc)
but it leads to very specific thermal-hydraulic issues.
As far as the reactor core design is concerned, a ceramic fuel concept with a good thermal conductivity
has been chosen. The main requirement is to obtain an average exit core temperature of 850°C (energy
conversion efficiency) with a maximum fuel temperature of about 1200°C and with a low core
pressure drop (in order to ease the decay heat removal). The main core characteristics have been
determined for two reactor powers: a medium one (600 MWth) and a large one (2400 MWth). For
various reasons, this latter became the CEA reference choice. A consistent set of core parameters has
been determined taking into account the different constraints including the thermal-hydraulics. The
reference arrangement proposed is based on plate fuel elements.
A significant issue for the GFR is the decay heat removal. An innovative approach has been chosen in
case of Loss Of Coolant Accidents (LOCAs). A “guard containment” enclosing the primary system is
used to maintain a medium gas pressure (10 bar) in order to remove the decay heat by low power
forced convection systems in the short term and natural convection systems in the long term. This
guard containment is not pressurized during normal operations and can be a metallic structure.
As far as the energy conversion system is concerned, an indirect-combined cycle has been chosen. The
significant advantages of this choice are : a moderate core inlet temperature (400°C instead of 480°C
for the direct cycle) and an attractive compactness of the primary system (facilitating the guard
containment design).
Due to the novelty of these options, a significant effort of components pre-sizing and design
calculations has been achieved. Following this effort, a CATHARE model of the reactor system has
been made and the calculation of the reactor steady-state confirms the consistency of the overall
system pre-sizing. This model has been used for a first transient calculation. Other types of transients
have to be analyzed, however, it is thought that the proposed GFR design can reach the safety
requirements of generation IV systems.
KEYWORDS
Nuclear system, Generation IV, Gas cooled reactors, Fast reactors
1. INTRODUCTION
These studies are partly driven by the Generation IV initiative devoted to the assessment of various
technologies and possible concepts of future nuclear systems. These systems should be competitive
and mature within the next 30-40 years and in the frame of a sustainable energy development. These
have been formulated into the following objectives:
Increased economical competitiveness (reduction of construction time and cost, high energy
conversion efficiency),
Increased safety (highly resistant fuel and materials in accidental situations, use of passive
safety whenever possible),
Reduction of long term radioactive waste (lifetime, activity, volume),
Optimal use of the energy available in the fuel material,
Highly proliferation resistant technologies and processes,
Ability for other uses (hydrogen production, water desalination, etc).
These objectives are strong incentives for the development of advanced fast reactors, their hard
neutron spectrum potentially allowing self-sustaining fuel cycles and integral actinides recycling
(minimizing waste production). On the other hand, helium coolant is very attractive because this inert
and neutron transparent gas is compatible with high core outlet temperatures
This is why the Gas cooled Fast Reactor (GFR) is one of the six concepts selected for further studies in
the frame of the Generation IV initiative and especially considered by CEA. Among CEA studies
(dealing with fuel, materials, technologies, …), reactor pre-conceptual design studies have been
undertaken in order to determine by the end of 2007 the GFR pre-viability [Garnier et al, 2006] [Martin
& Chauvin, 2005].
It is worth mentioning that during the sixties and the seventies, significant gas cooled fast reactors
projects were launched. For the core viewpoint, they were very similar to sodium cooled reactors: use
of oxide fuel cladded by stainless steel, use of fertile blankets, limited outlet gas temperature
(compatible with a Rankine indirect energy conversion cycle). These reactor projects were then
abandoned (development of sodium cooled reactors). The choice to re-examine GFR concepts was
based on:
the challenging Generation IV objectives more compatible with the high potentialities of
GFR concepts, and especially, the possibility to use more innovative fuel element design
with local fission product confinement,
the interest to go to high temperature allowing high energy conversion efficiency and
coupling with high temperature innovative processes producing hydrogen (steam
electrolysis, iodine sulfur thermo-chemical cycle).
The GFR design goals and criteria have been derived from the above general objectives. These are the
following:
Fast neutron spectrum reactor with a positive breeding gain with no or very limited use of
fertile blankets, this allows :
o an optimal use of the fuel energy content,
o a fuel cycle fed first with depleted and then with natural uranium,
o no plutonium separation from the minor actinides with a homogeneous actinides
recycling (proliferation resistance).
Core plutonium inventory below 10 t/GWe in order to have a realistic reactor fleet
deployment (few decades),
Use of a refractory fuel favoring close retention of fission products (increased safety)
Fuel residence time lower than 10 years (to be compatible with the Plutonium inventory
objective and to allow a dynamic fuel materials reprocessing),
High core outlet temperatures allowing a high thermodynamic efficiency and the use of
innovative high temperature hydrogen production processes,
High fuel burn-up, with easy reprocessing and re-fabrication.
Several other objectives could be added like the minimization of the reactivity swing per cycle (the
minimization of control rod worth is interesting from the safety point of view in limiting the
consequences of a rod ejection at full power).
For the thermal-hydraulic design point of view, the most significant requirements are the zero breeding
gain (because this limits the gas fraction in the reactor core) and the high core outlet temperature
(about 850°C to have a good efficiency, so a quite high value compared to most materials limits).
CEA has a significant R&D programme about the GFR fuel. Due to the high temperature foreseen,
fuel elements based on carbide or nitride ceramics are promising concepts. But today, only preliminary
results are available from this programme, then no definitive answers about the maximum
temperatures that these fuel concepts can hold. These temperatures are necessary to make core and
system pre-conceptual design studies. Taking into account the good thermal conductivity of carbide or
nitride ceramics, a maximum fuel temperature of 1200°C has been chosen for normal operating
conditions. For incidental and accidental design basis conditions (DBC), a maximum fuel temperature
of 1600°C has been chosen. This latter temperature should insure fission products confinement for a
significant time (several hours). Of course, these criteria will be modified, if necessary, in function of
the ongoing fuel R&D.
2. FROM 600 MWTH TO 2400 MWTH
The CEA GFR studies have started for a 600 MWth reactor power [Poette et al, 2004]. This choice
was mainly driven by the idea to use the Power Conversion Unit of the Gas Turbine Modular Helium
cooled Reactor (GT-MHR) in order to limit our research efforts to the reactor core and the safety
systems. The main options considered were the following:
a 600 MWth core with a high core power density, 100 MWth/m3,
a core installed within a metallic primary pressure vessel,
a Power Conversion Unit vessel connected to the core vessel by one single cross-duct.
Decay Heat Removal systems based on natural convection.
With these options, a first consistent set of reactor characteristics has been issued [Dumaz et al, 2005].
It was found out that this power level has significant drawbacks (very challenging fuel element in
order to have a significant fraction of actinide in the core, questionable economic competitiveness). It
was therefore decided to select a larger core power, 2400 MWth, compatible with the classical
“economy of scale” viewpoint. It was also an opportunity for opening the reactor design options :
energy conversion system, safety systems ….
3. CORE DESIGN STUDIES
3.1 General characteristics of the core
Since the very beginning of the GFR project, the main study objective has been to design a core that
will produce plutonium as much as it is burnt by fission, without fertile blankets (self breeding) with a
reasonable initial plutonium inventory (less than about 10 t/GWe). Considering the possible fuel
fraction in a gas cooled reactor, it has been decided to study as a first priority dense fuel compounds
(mixed carbides and nitrides). On the other hand, to get an easier safety demonstration, other
significant design specifications have been taken into account: a low core pressure drop (about 0.5 bar)
and a low maximum fuel temperature (about 1200°C in nominal operating conditions).
The proposed 2400 MWth core is based on a plate fuel element. Its design principle is illustrated by
the figures 1 and 2. Basically, two ceramic plates enclose the honeycomb structure which contains
pellets of the elementary fuel compound. The initial material choices are a mixed carbide (U,Pu)C
ceramic for the fuel compound and a silicon carbide (SiC) ceramic for the structures.
The coupling of the core neutronics (self breeding objective) and the core thermal-hydraulics
(temperature, pressure drop) constraints can be examined in a core height over diameter ratio versus
helium fraction diagram (figure 3). For a given core power density, the following lines can be drawn:
the thermal-hydraulic curve: the maximum fuel temperature (1200°C) will impose a
hydraulic diameter (in order to have a sufficient heat transfer coefficient); then, having a
hydraulic diameter, the pressure drop limitation (about 0.5 bar in normal operation) will
lead to a maximum height to diameter ratio (H/D). On this line, the maximum fuel
temperature and the core pressure drop are equal to the criteria considered (1200°C and 0.5
bar).
the neutronic curve: it is obvious that having all the core content defined (He coolant, fuel),
the self-sustaining condition imposes to have compact core in order to limit the neutron
losses (a key phenomenon in fast neutron reactors), then a minimum and a maximum H/D
can be obtained (in most cases, only the minimum value is used, the maximum being above
the thermal-hydraulic limit). On this line, one has a zero breeding gain.
These analyses have been conducted with the ERANOS computer code for the neutronics [Rimpault,
2002] and the COPERNIC tool for the thermal-hydraulics [Haubensack et al, 2004]. Various cases
were investigated in term of power density (50 and 100 MW/m3), fuel element arrangement and fuel
materials. For the proposed (U,Pu)C-SiC plate fuel element, the results are shown in the figure 4.
The chosen design (so-called “06/04” design) is characterized by a core height over diameter ratio of
0,35 and a 7 mm plate fuel thickness. The main core characteristics are summarized in the table 1. It is
worth noting that the neutronic characteristics are related to the first core (initial reactor start-up with
plutonium only):
The performances of the “06/04” core design were thought to be in good agreement with the high level
objectives : despite the first cycle breeding gain is found to be slightly negative, it has been shown that
the breeding gain increases for the multiple (Pu + Minor Actinides) recycling and finally becomes
slightly positive.
The core is implemented in a metallic reactor vessel (figure 5) with an internal diameter of 7.3 m. The
9CrMo steel reactor vessel thickness is about 0.19 m. It can be noticed that an upward helium flow
cools the reactor core, a choice which was initially made to ease the natural convection start-up during
accidental transients. The fuel handling is made by an articulated system in the upper plenum. In
consequence, the control rod mechanisms are implemented in the bottom of the vessel. It is worth
noting that the absorber rods are located above the core in normal operation in order to have a gravity
driven control rod insertion in case of scram.
3.2 Complementary core thermal-hydraulic studies
The core characteristics have been obtained with very simple models (COPERNIC tool) using
available correlations. Considering the importance of a good prediction of the core pressure drops,
some CFD calculations of the sub-assembly thermal-hydraulics have benn done using the STAR-CD
code. STAR-CD is based on a finite volume discretisation, the models used here are the k-ε High
Reynolds model for the turbulent zones and the laminar model for the low Reynolds zones.
Due to the sub-assembly complexity (Figure 6), the straightforward meshing would have led to a large
number of meshes, hence a prohibitive computation time. Several calculations have been done : subassembly inlet, bottom reflector, fissile zone, upper reflector, to have the overall pressure drop.
Three cases have been considered:
the nominal regime under full power at the design pressure of 70 bar,
a first Decay Heat removal regime (at 3 % of the nominal power) at 70 bar,
a second Decay Heat removal regime (at 3 % of the nominal power) at 10 bar.
The STAR-CD results and the comparisons with COPERNIC are given in the tables 2, 3 and 4.
As far as the global sub-assembly pressure drops are concerned, the comparisons are quite satisfactory.
The maximum pressure difference between STAR-CD and COPERNIC is indeed about 10 %. The
major difference, noticed in the assembly head, is explained by the fact that COPERNIC but not
STAR-CD takes into account the real broadening at the sub-assembly outlet between the upper head
and the upper plenum.
This study confirms that the sub-assembly pressure drop is close to the 0.5 bar objective.
4. GENERAL REACTOR ARRANGEMENT AND ENERGY CONVERSION
SYSTEM
The type of reactor arrangement considered here can be called “loop type” : the metallic vessel
containing the core being connected by primary loops to the main energy removal system : turbomachinery for direct cycles, intermediate heat exchanger for indirect cycles. Considering the size of
these latter and the priority given to metallic vessels (instead of concrete vessels), the “integrated
option” appeared to be an unrealistic way. The general reactor arrangement is driven by the choice of
the energy conversion system : direct or indirect cycles ?
A significant effort has been made to compare various solutions. In term of thermodynamic efficiency,
for a given core outlet temperature (here 850°C) and a good heat sink (sea temperature of 15°C) the
differences are not so significant (see the table 5 for few cases).
In addition, the relationship between the efficiency and the investment cost has to be investigated more
deeply. After having chosen the direct Brayton cycle for the 600 MWth case, we made the choice of
the indirect-combined cycle for the 2400 MWth case, four main reasons explain this choice :
-
its efficiency (45 %) seems low compared to the direct cycle, but in fact significant
efficiency improvements are expected, in particular by using a supercritical steam cycle
(about 2 to 3 % of efficiency is expected from 150 to 300 bar) On the other hand, the direct
cycle efficiency is based on component efficiencies which are really challenging and require
a significant R&D;
the maximum efficiency of this cycle is obtained with a “low” core inlet temperature :
400°C (versus 480°C or more for the direct cycle). This feature greatly simplifies the
reactor vessel design (no thermal insulation required between the inlet gas flow and the
vessel wall) ;
the primary system is much more compact than one for the direct cycle, thus, it is much
easier to design the “guard containment”, a quite interesting design option considered (see
the Decay Heat Removal discussions) ;
an indirect system allows the use of a helium nitrogen mixture on the secondary side ; the
physical properties of this mixture are sufficiently close to those of air to allow the use of
existing gas turbine technology, which reduces dramatically the turbomachine development
cost and the technological risk ;
Moreover it is interesting to note that AREVA made a similar choice for its ANTARES VHTR project
[Petit et al, 2005].
The main characteristics of the chosen energy conversion system are given in the figure 7. It is worth
noting that the inherent electricity needs (primary circuit blower of 22 MWe, steam cycle pumping of
10 MWe and other auxiliary systems) are taken into account for the calculation of the net efficiency.
For this indirect-combined cycle, an IHX concept (Intermediate Heat eXchanger) including the
primary helium blower has been studied. With this IHX module and the decay heat removal loop
design, an integration of the primary circuit into the metallic guard containment has been done. A
spherical containment with a diameter of 30 m has been chosen (figure 13). Its thickness is 38 mm in
order to avoid any thermal treatment after welding. The containment size is consistent with the decay
heat removal requirements (free volume of 8600 m3 to obtain the 10 bar pressure after a large break
LOCA, see the next section).
5.
DECAY HEAT REMOVAL STUDIES
Within the framework of GFR studies, safety constraints are addressed as far as possible at the earliest
phase of conceptual design. Due to the low He thermal inertia, one of the major GFR design issues,
which impacts the overall system architecture is the Decay Heat Removal. Therefore, the Loss Of
Coolant Accident (LOCA) is a key transient to be considered. In the pre-existing solution, coming
from the 600 MWth GFR studies, the decay heat was removed by means of a fully passive natural
circulation system, system which requires to keep the reactor, in all cases, at high gas pressure. This
pressure is called the back-up pressure. Similarly, for the GFR2400 case, a back-up pressure of 30 bar
is required to have enough natural convection few minutes after the reactor scram. Such a pressure can
be obtained with a heavy pre-stressed concrete “guard containment” which encloses all the primary
system, and which has to be kept pressurized during normal operation. This solution - called the “high
back-up pressure strategy” - was not retained for the 2400 MWth case due to the significant
investment cost of such “guard containment”.
An alternative is to have a very large containment, as for the PWR. In such a case, the pressure after a
LOCA cannot be very different from the atmospheric pressure. In these conditions, it is not possible to
use natural convection; forced convection systems must be implemented with a very significant
pumping power (more than 1 MW). For such powers, only active systems seem usable. This way,
called “low back-up pressure strategy” has not been chosen for the design presented here.
Another alternative for the LOCA management, is the implementation of a “medium back-up pressure
strategy”. In fact, for laminar flow conditions, and for a given mass flowrate, the pumping power is
almost inversely proportional to the square of the gas pressure. This means that the pumping power
decreases very significantly when the pressure increases. At 10 bar, about 15 to 20 kW are required a
few minutes after the reactor scram instead of more than 1 MW at 1 bar. For such a low pumping
power diverse passive or “semi passive” energy sources are available, like batteries. In addition,
keeping such medium back-up pressure leads to recover natural convection after about 24 hours, due
to the decay heat decrease. Finally, such medium back-up pressure can be coherent with the use of
metallic guard containment.
With this strategy, simpler and smaller decay heat removal means can be used. Therefore, in order to
limit investment, operation and maintenance costs, the decay heat removal system has been designed
based on this medium back-up pressure strategy.
So, the system design proposed is based on (see the figure 8) :
a metallic guard containment enclosing the primary system, not pressurised during normal
operation, having a free volume such as the fast primary helium expansion gives an
equilibrium pressure of 10 bar,
three dedicated loops (3*100% redundancy) with secondary loops connected to an external
water pool (the ultimate heat sink). Each of these loops can remove the decay heat :
o for high pressure transients (blackout typically) : by natural convection,
o for loss of pressure transients (large break LOCA typically) : by forced convection (in
this case the pressure is equal to the chosen back-up pressure : 10 bar); forced
convection is obtained by a blower driven by an electrical motor.
For the medium pressures obtained in case of LOCA, due to the limited pumping power, the energy
required by this electrical motor can be supplied by batteries for at least 24 hours after the reactor
SCRAM. The required pumping energy is about 1800 MJ (500 kW.h). Assuming a conversion ratio of
80% and a battery specific volume of 50 W.h/l (classical lead-acid battery) the required battery volume
per loop would be about 12.5 m3. This volume seems quite reasonable.
After these 24 hours and if necessary, it is possible to use natural convection only due to the significant
decrease of the decay heat (maintaining a pressure of 10 bar).
A set of CATHARE calculations support this design. CATHARE is the French best-estimate code used
for safety analysis of Light Water Reactors. Two-phase flows are modeled using a two-fluid 6-equation
model in 0-D, 1-D and 3-D modules. Since the year 2000, significant efforts have been made to extend
the range of CATHARE applicability to helium cooled reactors. The efforts made include the modeling
of very specific components like turbines and compressors. The CATHAR2 v2.5 version was used.
5.1 CATHARE modeling
The CATHARE model dedicated to decay heat removal system design verifications is composed of the
primarysystem, including :
the main vessel comprising the core which is consisting of plate-type assemblies,
the main primary loops connected to boundary conditions to simulate the connection to the
“IHX - main blower” vessels (not modeled here),
three dedicated decay heat removal primary loops.
Each loop has its own water-filled secondary and pool-type ternary circuits.
The reactor core is modeled by six one-dimensional thermal-hydraulic modules (table 6). The decay
heat coming from DARWIN code calculations (run with 2 wt% of Minor Actinide loading and at a
burn-up of 10% FIMA : Fissions per Initial Metal Atom). To have a better prediction of the core
pressure drops, the detailed geometrical of the sub-assembly (inlet nozzle, neutron shielding, reflector,
…) is used .
5.2 Design for the forced convection at medium pressure (10 bar)
The design situation has been assumed to be a large break LOCA leading to a rapid loss of the primary
pressure. The calculation conditions selected are therefore:
t=0s, from nominal conditions to primary pressure = 10 bar and primary mass flow rate = 0
kg/s in 0.01 s (no inertia of the primary flow) ;
Shutdown rods fully inserted (-10 $) : 0.5 s, then decay heat law ;
decay heat removal loop valves fully opened : 2 + 10 s (opening period), then startup of the
blower (with a constant rotating speed, typical of asynchronous motors).
The calculation is run during 600 s of real time, due to the decay heat decrease and the significant
decrease of the core temperatures (see the Figure 9), there is no need to prolong the calculation (the
only issue being the ultimate heat sink, which is assumed here to do not be a concern).
Finally, to respect the design criteria selected, we obtained the following maximal characteristics for
the decay heat removal loop :
a primary mass flow rate of ~ 21 kg/s ;
a heat exchanger capacity of ~ 100 MWth.
It is worth noting that for this forced convection design, the difference of elevation between the core
and the heat exchanger does not affect the results, this parameter must be determined by natural
circulation cases.
The decay heat removal blower mass flowrate remains almost constant (figure 10). As it has been
assumed that the blower rotating speed is kept constant, the slight flowrate increase comes from the
gas temperature decrease in the system, leading to lower system pressure drops (decrease of the
helium dynamic viscosity). That behavior also explains the evolution of the blower power, about 21
kW, slightly decreasing.
5.3 Design for natural convection at high pressure (70 bar)
For natural convection at 70 bar, the design situation has been assumed to be a rapid and total loss of
forced circulation capabilities (the primary pressure remaining close to the nominal pressure). Given
margins were expected in this situation, the objective was mainly to identify the minimum driving
heights required (for integration considerations, the maximum value is about 15 m, measured between
the core mid plane and the decay heat removal exchanger mid plane) and the corresponding heat
exchangers designs and then to provide interesting informations on optimization capabilities related to
integration viewpoint. The calculation conditions selected were :
t=0 s, from nominal conditions to : primary mass flow rate = 0 kg/s in 0.01 s (no inertia of
the primary coolant flow) ;
Shutdown rods fully inserted (-10$) : 0.5 s, then decay heat law ;
decay heat removal loop valves fully opened : 2 + 10 s (10 s of opening time).
The calculation is run during 600 s of real time, due to the decay heat decrease and the significant
decrease of the core temperatures (see Figure 11) there is no need to prolong the calculation.
With the heat exchanger designed in forced convection, the final parameters selected are :
a primary driving height = 9 m,
a secondary driving height = 5 m.
5.4 Possible swing from forced to natural convection at 24 hours
It has been explained that the choice of a backup pressure of 10 bar gives the possibility of recovering
the natural convection about 1 day after the reactor SCRAM. This has been checked with one of the
previous CATHARE calculations extended up to 100000 s. During the first 86400 s (1 day) we let the
decay heat removal blower in operation, then, the blower is switched off. In this calculation, the
blower rotating velocity is kept constant.
The fuel temperatures evolutions can be seen in the Figure 12. During the first day, given the constant
blower velocity, the forced convection mass flow rate keeps staying constant. Due to the decay heat
decrease, the fuel temperatures are significantly reduced down to 300°C (the gas temperature being
very close, about 280°C). Due to the blower switch off, a significant decrease of the flow rate is
observed : from 28 kg/s down to 5 kg/s. This leads to an increase of the fuel temperature up to 1100°C
for the maximum. A new quasi steady-state is obtained after, the decay heat being removed by natural
convection only. This calculation confirms the capability of the decay heat removal system to recover
the natural convection operation after one day.
6.
CONCLUSIONS
In comparison with previous CEA studies where the energy conversion system was similar to the one
used in the GT-MHR project (case of a 600 MWth reactor) a much significant design effort has to be
made to obtain the poroposed system (figure 13). For all these new components : blowers, IHX, …
pre-sizing has been made.
With a CATHARE model, the decay heat removal issue has been studied with a special attention. It is
showed that the loss of the primary coolant flow is well covered by the natural convection in the decay
heat removal system. As far as the LOCAs are concerned, the calculations conducted to design the
decay heat removal system give very good hope about the system resistance to such initiating events.
Other types of initiating events and transients have to be analysed, however, the very significant
analyses achieved for the 600 MWth GFR did not raise unsolvable issues.
The GFR system arrangement as described above is the new CEA reference for further studies on the
GFR. A significant effort has been made to have studies as comprehensive and consistent as possible.
This reference presents new options and it is thought that the GFR can reach the safety requirements of
generation IV systems. These options will have to be consolidated by detailed studies leading to a previability of the GFR design by the end of 2007.
REFERENCES
J.C. Garnier et al, "Contribution to GFR design option selection", ICAPP’06 conference, Reno, USA,
June 4-8 (2006)
P. Martin, N. Chauvin, “Gas cooled Fast Reactor system : major objectives and options for reactor,
fuel and fuel cycle”, GLOBAL’05 conference, Tsukuba, Japan, Oct. 9-13 (2005)
P. Dumaz et al, “The thermal-hydraulic studies in support to the GFR pre-conceptual design”,
NURETH11 conference, Avignon, France, October 2-6 (2005)
D. Haubensack et al, “The COPERNIC/CYCLOP computer tool, the pre-conceptual design of
Generation IV nuclear systems”, 2nd International Topical Meeting on HIGH TEMPERATURE
REACTOR TECHNOLOGY, Beijing, China, September 22-24 (2004)
D. Petit et al, “Overall simulation of a HTGR plant with the gas adapted MANTA code”, NURETH11
conference, Avignon, France, October 2-6 (2005)
C. Poette et al, “Advanced gas cooled fast reactor preliminary design – 300 MWe – Project status and
trends for a higher unit power selection”, Proceedings of ICAPP ’04 Pittsburgh, PA USA, June 13-17,
2004 Paper 4071
G. Rimpault, "The ERANOS Code and Data System for Fast Reactor Neutronic Analyses",
PHYSOR2002 Conference, Seoul, Korea, October 7-10 (2002)
List of figures :
Figure 1: Design principle of the plate fuel element
Figure 2: Schematic of the sub-assembly
Figure 3: Schematic of the combination of thermal-hydraulics and neutronics constraints
Figure 4: Core feasibility diagram
Figure 5; Schematic of the reactor vessel (including two of the three loops dedicated the Decay Heat
Removal, see the section 5)
Figure 6; Overview of a fuel sub-assembly
Figure 7: Schematic of the indirect-conbined cycle
Figure 8: Schematic of the DHR system
Figure 9: Maximum fuel temperatures (°C) under forced convection operation
Figure 10: DHR blower behavior
Figure 11: Maximum fuel temperatures (°C) under natural convection at 70 bar
Figure 12: Fuel temperatures (°C) for different sub-assemblies versus time and during a 100000 s
(28 h) transient at a back-up pressure of 10 bar
Figure 13: Schematic of the reactor building
Figure 1: Design principle of the plate fuel element
Figure 2: Schematic of the sub-assembly
H/D max (thermal-hydraulics)
Height/Diameter
Possible solutions
H/D min (neutronics)
He fraction
Figure 3: Schematic of the combination of thermal-hydraulics and neutronics constraints
1,50
1,40
1,30
CerCer (U,Pu)C - SiC 50% vol.
Neutronic
Neutronic
linecurve
Burn-up = 10 FIMA
1,20
H/D ratio
1,10
1,00
0,90
P=0.8
P= -0,8 barbar
0,80
0,70
 P=
-0,7 bar
P=0.7
bar
0,60
 P=
-0,5 bar bar
P=0.5
0,50
0,40
fisssile
Thermal-hydraulic
lines curves
Thermal-hydraulic
0,30
0,20
0,10
0,00
25
30
35
40
45
Core He fraction
Figure 4: Core feasibility diagram
50
55
Figure 5; Schematic of the reactor vessel (including two of the three loops dedicated the Decay Heat
Removal, see the section 5)
S/A head
Reflectors
Neutron shielding
Fissile zone
S/A bottom
Figure 6; Overview of a fuel sub-assembly
Figure 7: Schematic of the indirect-conbined cycle
pool
DHR-HX2
3 DHR Loops
(3x100%)
water
DHR-HX1
Emergency
blowe r
Guard
containment
He
He 20wt% + N2 80wt%
IHX
Main
blowe r
3 Main Loops
(3x800 MW)
Bypass line
water
Steam generator
Figure 8: Schematic of the DHR system
CATHARE V370: GFR2400
- 06/2004, Tin=400°C, Equ, Het, Darwin, Lam
LOCA (7 to 1 MPa in 0,01s, Q=0 in 0,0
1s, Scram at 0,5s), 1 DHR blower (at 2+10s)
1450
1350
1250
1150
1050
C ore hot assembly
Core mean assembly
CORE1 mean assembly inner core
CORE3 hot assembly outer core
CORE4 mean assembly outer core
CORE5 cold assembly
950
850
750
0
100
200
300
400
500
600
time (s)
Figure 9: Maximum fuel temperatures (°C) under forced convection operation
CATHARE V370: GFR2400-06/2004, Tin=400°C, Equ, Het, Darw in, Lam
31
24
29
22
27
20
25
18
23
16
21
14
Mass flow rate
19
12
Pumping pow er
17
10
15
8
0
60
120
180
240
300
360
420
480
time (s)
Figure 10: DHR blower behavior
540
600
Blower pumping power (KW)
Blower mass flow rate (Kg/s)
LOCA (7 to 1 MPa in 0,01s, Q=0 in 0,01s, Scram at 0,5s), 1 DHR blow er (at 2+10s)
CATHARE V370: GFR2400-06/2004, Tin=400°C, Equ, Het, Darw in, Lam
BLACK-OUT (P=7MPa, Q=0 in 0,01s, Scram at 0,5s), 1 DHR loop (at 2+10s)
maximum fuel temperature (°C)
1300
1250
1200
1150
1100
1050
1000
Core hot assembly
Core mean assembly
CORE1 mean assembly inner core
CORE3 hot assembly outer core
CORE4 mean assembly outer core
CORE5 cold assembly
950
900
850
800
0
50
100
150
200
250
300
350
400
time (s)
Figure 11: Maximum fuel temperatures (°C) under natural convection at 70 bar
Figure 12: Fuel temperatures (°C) for different sub-assemblies versus time and during a 100000 s (28
h) transient at a back-up pressure of 10 bar
Figure 13: Schematic of the reactor building
List of tables :
Table 1: Core characteristics of the 2400 MWth case
Table 2: Sub-assembly pressure drops at nominal conditions
Table 3: Sub-assembly pressure drops at for Decay Heat Removal conditions (70 bar)
Table 4: : Sub-assembly pressure drops for Decay Heat Removal conditions (10 bar)
Table 5: Energy conversion cycles comparison
Table 6: Radial modelling for the core
Pressure (bar)
70
-3
Power density (MW.m )
100
Core inlet / outlet temperature (°C)
400 / 850
Core height / diameter (m)
1.55 / 4.44
Maximum fuel temp. BOL (°C)
1260
Number of fuel S/A
387
Number of plates per S/A
27
Plate element thickness (mm)
7
(U,Pu)C fuel / Coolant volume fractions %
22.4 / 40.
TRU content (%)
15.2
Pu inventory (t/GWe1)
8.2
Core management (eq. full power days) 3  831 = 2493
Breeding Gain*
-0.07 / -0.04
Doppler Constant* (pcm)
-1872 / -1175
He depressurization* (pcm)
212 / 253
Delayed neutron fraction* (pcm)
388 / 344
* Beginning Of Life / End Of Life
Table 1: Core characteristics of the 2400 MWth case
STAR-CD
Pressure drop (Pa)
Tin = 400°C
COPERNIC
Relative difference
Tout = 850°C
Friction +
Pressure drop (Pa)
(%)
Gravity
P=70 bar
acceleration
S/A head
860
3.5
2395
63
Neutron shielding
2574
15
Reflector
8220
14
11401
5
Fissile zone
19892
58
21838
8 .6
Reflector
4780
23
Neutron shielding
1470
25
6818
7.6
S/A bottom
353
33
390
1
Total
38149 Pa
171 Pa
42842 Pa
10.5 %
Table 2: Sub-assembly pressure drops for nominal conditions
STAR-CD
Tin = 330°C
Tout = 1530°C
P=70 bar
S/A head
Neutron shielding
Reflector
Fissile zone
Reflector
Pressure drop (Pa)
Friction +
acceleration
0.3
1.5
4.8
47
1.1
Gravity
2.2
9.5
8.6
47
25.6
COPERNIC
Pressure drop (Pa)
Relative difference
(%)
4
37
22
82
55
-10
-12
0
Neutron shielding
0.33
28
S/A bottom
0.044
36
33
-9
Total
56 Pa
156 Pa
196 Pa
-8 %
Table 3: Sub-assembly pressure drops for Decay Heat Removal conditions (70 bar)
STAR-CD
Pressure drop (Pa)
Tin = 330°C
COPERNIC
Relative difference
Tout = 1530°C
Friction+
Pressure drop (Pa)
(%)
P=10 bar
acceleration
Gravity
S/A head
2
0.3
7
67
Neutron shielding
10
1.3
Reflector
34
1.2
32
-45
Fissile zone
253
7
245
-10
Reflector
8
3.7
Neutron shielding
2.2
4
18
0
S/A bottom
0.3
5.3
6
6
Total
319 Pa
23 Pa
308 Pa
-11 %
Table 4: : Sub-assembly pressure drops for Decay Heat Removal conditions (10 bar)
Cycle
Net
Primary
(secondary)
inlet/outlet
efficiency temperatures and Pressures
48.2%
480 - 850°C, 70 bar
He
Direct Brayton cycle
Indirect, Nitrogen Brayton cycle on the secondary 46.8%
side
45.1%
He-N2
He
480 - 850°C
(444 - 820°C)
70 bar
(65 bar)
400 - 850°C
(364 - 820°C)
70 bar
(65 bar)
Steam Pressure = 150 bar
H2O hp
Indirect-combined cycle
Nitrogen/helium on the secondary
Steam on the ternary
Table 5: Energy conversion cycles comparison
CORE 0
Sub-assembly
represented
Number
of
Sub-assemblies
Equivalent
radius (m)
Normalized
power profile
CORE 1
CORE 2
7
169
17
CORE 3
“Hot” for
high Pu
content core
zone
17
0.295
1.480
1.550
1.617
2.152
2.192
0.9971
0.9254
0.8509
0.9391
0.7680
0.5673
“Hottest” Mean for low
Mean for the
SubPu content
whole core
assembies
core zone
CORE 4
Mean for
high Pu
content core
zone
162
CORE 5
“cold”
for high Pu
content core
zone
14
Table 6: Radial modelling for the core