Nuclear Engineering and Design 230 (2004) 195–207
Medium temperature carbon dioxide gas turbine reactor
Yasuyoshi Kato∗ , Takeshi Nitawaki, Yasushi Muto
Research Laboratory for Nuclear Reactors, Tokyo Institute of Technology, 2-12-1 O-okayama, Meguro-ku, Tokyo 152-8550, Japan
Received 8 May 2003; received in revised form 2 October 2003; accepted 5 December 2003
Abstract
A carbon dioxide (CO2 ) gas turbine reactor with a partial pre-cooling cycle attains comparable cycle efficiencies of 45.8% at
medium temperature of 650 ◦ C and pressure of 7 MPa with a typical helium (He) gas turbine reactor of GT-MHR (47.7%) at high
temperature of 850 ◦ C. This higher efficiency is ascribed to: reduced compression work around the critical point of CO2 ; and
consideration of variation in CO2 specific heat at constant pressure, Cp , with pressure and temperature into cycle configuration.
Lowering temperature to 650 ◦ C provides flexibility in choosing materials and eases maintenance through the lower diffusion
leak rate of fission products from coated particle fuel by about two orders of magnitude. At medium temperature of 650 ◦ C, less
expensive corrosion resistant materials such as type 316 stainless steel are applicable and their performance in CO2 have been
proven during extensive operation in AGRs. In the previous study, the CO2 cycle gas turbomachinery weight was estimated to
be about one-fifth compared with He cycles. The proposed medium temperature CO2 gas turbine reactor is expected to be an
alternative solution to current high-temperature He gas turbine reactors.
© 2004 Elsevier B.V. All rights reserved.
1. Introduction
An overview of major gas-cooled power plant
projects is shown in Fig. 1 that have been built or are
in the planning stages.
The first gas-cooled power reactor, the MAGNOX
reactor, was developed by the UK and France using carbon dioxide (CO2 ) coolant in a closed cycle
at 2 MPa pressure. The core consisted of graphite
moderator blocks with holes and fuel elements. Fuel
elements, consisting of a natural uranium metal bar
and magnesium–aluminum alloy (called Magnox)
cladding, were placed into the holes. Coolant flowed
through moderator holes and the temperature at the
core outlet was about 400 ◦ C. Steam cycle efficiency
was about 31%. Subsequent plants, designated as
∗ Corresponding author. Tel.: +81-3-5734-3065;
fax: +81-3-5734-2959.
E-mail address: kato@nr.titech.ac.jp (Y. Kato).
advanced gas-cooled reactors (AGRs), used 2.3% enriched uranium oxide fuel pellets and stainless steel
cladding to achieve higher cycle efficiency of about
40% through elevating coolant outlet temperature to
650 ◦ C. Coolant pressure of AGRs was elevated to
4 MPa to increase average core power density 4–6
times that of MAGNOX reactors for reduction of
plant capital costs. However, AGRs were not competitive with PWRs in electricity generation costs. The
UK government announced in 1979 that the future
nuclear power generation program would be based on
construction of PWRs.
The US and Germany began high-temperature
gas-cooled reactor (HTGR) development programs in
the 1960s. Using helium (He) as coolant and small
particle fuels with enriched uranium oxide, a high
core outlet temperature is attained ranging from 800
to 850 ◦ C without any chemical attack on moderator
and fuel materials. Small particles are triso-coated
with successive layers of porous carbon, pyro-carbon,
0029-5493/$ – see front matter © 2004 Elsevier B.V. All rights reserved.
doi:10.1016/j.nucengdes.2003.12.002
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Y. Kato et al. / Nuclear Engineering and Design 230 (2004) 195–207
of a stack of graphite spheres of about 60 mm diameter with embedded coated fuel particles. The latter
reactor core consists of hexagonal graphite blocks
with holes and fuel rods. Fuel rods enclosing coated
particles are inserted into the holes. A number of
experimental and subsequent prototype HTGR plants
were built for both types; however, no commercial
plant was constructed because of the higher capital
cost than contemporary LWRs.
Helium heated to such high temperatures as
800–850 ◦ C can be used as a working fluid in a closed
gas turbine direct cycle for a power generation system.
The direct cycle eliminates the intermediate cooling
circuit and steam generators required for PWR. Compared to a steam turbine, a gas turbine is far smaller
because gas turbine outlet pressure is much higher
than that of the former because of the low pressure
ratio (about four) in the gas turbine. Moreover, the
gas turbine is much simpler than that of the steam
turbine because of absence of the moisture separation and steam extraction systems necessary for the
steam turbine. The gas turbine system provides about
46–48% efficiency. These promising features indicate
the possibility of the gas turbine direct cycle as a
Nomenclature
Cp
d
hstage
k
Nu
P
Pcrit
Pred
Pr
R
Re
Tred
Tcrit
T
V
z
specific heat at constant pressure
hydraulic equivalent diameter
enthalpy drop per turbine stage
heat conductivity
Nusselt number
gas pressure
critical pressure
reduced pressure
Prandtl number
gas constant
Reynolds number
reduced temperature
critical temperature
temperature drop in a turbine
gas volume
compressibility factor
and silicon carbide to prevent fission product release.
Two types of HTGRs were developed: pebble-bed
fuel type reactors in Germany and block fuel type
reactors in the US. The former reactor core consists
Coolant
Cycle
1950
1960
1970
1980
1990
2000
Magnox reactor (60-655 MWe)
*1
UK & 1956
Steam Indirect France
(Rankine)
AGR (660 MWe)
1976*1
UK
CO2
TIT
Gas Turbine
Direct
(Brayton)
Japan
MIT
US
1967*2 Peach Bottom (42 MWe)
Steam Indirect
(Rankine)
US
1967*2 AVR (15 MWe)
Germany
He
Gas Turbine
Direct
(Brayton)
*
1982*1 Fort St. Vrain (324 MWe)
1986*2
THTR-300 (308 MWe)
PBMR (100 MWe)
S. Africa
GT-MHR (286 MWe)
Russia
GTHTR-300 (275 MWe)
JAERI
1: Start of operation, *2: Rated full power operation
Fig. 1. History of gas-cooled power plant development projects.
Y. Kato et al. / Nuclear Engineering and Design 230 (2004) 195–207
197
Table 1
Critical data of typical fluids
Net Work
Turbine Work
Turbine Work
Net Work
Pump Work
SteamTurbine
Fluids
Critical temperature
Tc (◦ C)
Critical pressure
Pc (MPa)
CO2
H2 O
N2
NH3
31.0
374.0
147.0
132.2
7.38
22.06
3.40
11.13
Compressor Work
Gas Turbine
Fig. 2. Comparison of compressor work in a gas turbine and pump
work in a steam turbine.
potential alternative to LWRs. However, extensive development is required for high-temperature resistant
materials and the He gas turbine.
The principal difference is that the steam turbine
works with a fluid that changes phase during the Rankine cycle, whereas the gas turbine works on fluids
that remain in the gaseous phase during the Brayton
cycle. As shown in Fig. 2, feed-pump work to pressurize water in the liquid phase in the steam turbine
is few percents of steam turbine output (Wilson and
Korakianitis, 1998), whereas the gas turbine compressor consumes about 50% of the power produced by the
turbine. By far, the main efforts to enhance cycle efficiency have specifically addressed increasing turbine
output by elevating turbine inlet temperature, while
enhancement of recuperator efficiency is also crucial
to the cycle in the gas turbine system.
Fluids have peculiarities at their critical points or
pseudo-critical points in thermo-mechanical properties: they exhibit strong peaks in specific heat, thermal
conductivity and viscosity. Moreover, they have a
peculiarly deep drop in their compressibility factors.
Surface tension and heat of vaporization become zero
at the critical point. This study investigated a new
method in a CO2 cycle for improving cycle efficiency
by reduction of compressor work through utilizing
the peculiar drop in the compressibility factor around
the critical point. The usual compressor inlet temperature of about 35 ◦ C, which is usually determined
by cooling sea water temperature, and critical temperatures of fluids shown in Table 1 present a great
advantage for this system: a peculiar drop in a compressibility factor around the critical point can be
utilized to reduce compressor work in a CO2 gas turbine cycle because its critical temperature (31.0 ◦ C)
is approximately equal to the usual compressor inlet
temperature (35 ◦ C). It can not be utilized practically
in the conventional He gas turbine cycle because its
critical temperature (−268.0 ◦ C) is much lower than
the usual compressor inlet temperature.
Coolant CO2 has preferable properties to He. Heat
transfer coefficient h between the coolant and fuel
cladding surface is calculated as
k
k
h=
Nu ∝
(Re0.8 Pr0.4 ),
d
d
where k is heat conductivity, d = hydraulic equivalent diameter, Nu = Nusselt number, Re = Reynolds
number, and Pr = Prandtl number. The above equation provides 1.5 times higher heat transfer coefficient
h in CO2 at the same gas velocity, leading to about
a 1.5-times-lower temperature difference between the
coolant and fuel rod cladding surface. The higher heat
transport capacity, measured as product of specific
heat at constant pressure Cp and molecular weight,
results in about 2.5 times more effective core decay
heat removal under natural circulation conditions than
He. The 3.6 times longer depressurization time of
CO2 (Lewis, 1977) and higher heat transport capacity (mentioned above) over that of He mitigates the
depressurization transient and simplifies design of a
passive decay heat removal system described later.
Carbon dioxide is about 250 times less expensive
than He per unit weight and 24 times less expensive per unit volume. In addition, considering lower
leakage rate characteristics of CO2 than He, coolant
leakage problems of gas cooled reactors in operation
are orders of magnitude less severe than with He.
For the He direct cycle, a He gas turbine must be
newly developed because He and air have substantially different gas properties, e.g. density (molecular weight), specific heat, and specific heat ratio.
198
Y. Kato et al. / Nuclear Engineering and Design 230 (2004) 195–207
In contrast, most current gas turbine engine technologies developed for fossil fuel power generation
systems could be borrowed because the difference
between CO2 and air is much smaller than that between He and air. Considering turbine inlet temperature of modern gas turbines, the so-called “F” class
gas turbine is 1350 ◦ C; therefore, material technology needed for a CO2 turbine up to 850 ◦ C is within
current technology of fossil gas turbines.
Carbon dioxide may react with structural materials and a graphite moderator under high temperature
and radiation fields. During extensive experience in
MAGNOX reactors and advanced gas cooled reactors
(AGRs) over more than 20 years, structural material
corrosion problems were eliminated by appropriate
material selection according to service temperatures
and by reduced vapor content (Gibbs and Popple,
1982). Regarding corrosion of the graphite moderator,
addition of methane and carbon monoxide in a range
of small concentrations (called the coolant “window”)
inhibited reaction with CO2 ; thereafter, satisfactory
operations were done without excessive corrosion
(Hewitt and Collier, 1997). An indirect Rankine CO2
cycle with a water/steam system has been applied to
AGRs and enhanced gas cooled reactors in the UK
(Gratton, 1981; Lennox et al., 1998; Abram et al.,
2000), for which the means proposed in this study
to improve cycle efficiency can not be employed because recompressing process is absent in this indirect
cycle. However, a medium temperature CO2 direct
cycle proposed in this study could be inherited with
the above advantages and experiences.
Now, worldwide interest in small and medium
size reactors continues to increase for electricity generation and local heating for cities and islands. In
December 1999, the Research Laboratory for Nuclear Reactors of the Tokyo Institute of Technology
launched a new program to develop advanced small
and medium size nuclear reactors. Under the program
framework, we are designing advanced gas-cooled
fast and thermal reactors (Kato et al., 2000, 2002,
2003; Nitawaki et al., 2001; Muto et al., 2003) and
light water reactors (Yamashita et al., 2001; Ohtsuka
et al., 2002) in collaboration with industry.
Recently, various CO2 gas turbine cycles have been
studied by us and at MIT (Kato et al., 2000, 2001,
2002, 2003; Dostal et al., 2002; Hejzlar et al., 2002).
This paper addresses the possibility of a medium
temperature CO2 gas turbine reactor with a partial
pre-cooling cycle to provide alternatives to current
high-temperature He gas turbine reactors, by evaluating thermal cycle performance.
2. Results and discussion
Cycle efficiencies of He and CO2 cycles were
calculated using PROPATH as a database for
thermo-physical properties of fluids (Ito et al., 1990)
and evaluation conditions are summarized in Table 2.
Cycle efficiency of a Brayton cycle is improved
when the number of compression stages with
inter-cooling is increased as shown in Fig. 3(a)–(c).
However, the contribution of each additional stage
Table 2
Cycle efficiency evaluation conditions
Parameters
Component efficiency
Turbine adiabatic efficiency (%)
Compressor adiabatic efficiency (%)
Recuperator effectiveness (%)
Pressure drop (%)a
Usual Brayton cycle (no pre-cooling)
Partial pre-cooling cycleb
a
b
c
Values
90
90
95
Reactor
Pre-cooler
Intercooler
Recuperator (Hb /Lc )
Reactor, pre-cooler, intercooler
Recuperator (Hb /Lc )
Percent pressure drop in each component relative to the reactor outlet pressure.
High-temperature side of a recuperator.
Low-temperature side of a recuperator.
1.5
1.02
0.58
1.99/0.66
Same as above cycle
2.65/0.88
Y. Kato et al. / Nuclear Engineering and Design 230 (2004) 195–207
Turbine
Compressor
199
Low Pressure High Pressure Turbine
Compressor Compressor
Generator
Reactor
Generator
Reactor
Intercooler
Pre-Cooler
Pre-Cooler
Recuperator
Recuperator
(a)
(b)
Middle Pressure
Compressor
High Pressure Turbine
Low Pressure
Compressor
Compressor
Generator
Reactor
Intercooler I
Intercooler II
Pre-Cooler
Recuperator
(c)
Fig. 3. Variation of Brayton cycles with the number of intercoolers. (a) No intercooler, (b) one intercooler, (c) two intercoolers.
to cycle efficiency becomes less and less; use of
more than three stages with two intercoolers can not
be justified economically. Hence, comparison of cycle efficiency is done in this study for one and two
compression stages between He and CO2 in closed
gas-turbine direct cycles. It is difficult to apply the
two-intercooler cycle to the He cycle because gas
turbine rotor length is much longer in a He cycle
than in a CO2 cycle; rotor dynamics design becomes
much more difficult if another compressor is added.
Cycle configurations of Fig. 3(a) and (b) were applied
to actual designs of He cycle HTGRs (Muto, 2000;
Kumar et al., 2001; Baydakov et al., 2001).
Cycle efficiencies in these usual Brayton cycle configurations are almost identical between He and CO2
cycles as given in Table 3, although the configurations
are not preferable for the CO2 cycle, as explained
hereafter. Specific heat at constant pressure Cp of
CO2 is dependent upon pressure and temperature,
whereas Cp of He is constant and Cp of CO2 in the
recuperator is considerably lower in the low-pressure
high-temperature side (connected to the turbine outlet) than in the high-pressure low-temperature side
(connected to the compressor outlet). Consequently,
CO2 can not be pre-heated to such temperature at
the core inlet because it provides maximum cycle
efficiency. If flow is by-passed to the compressor
before pre-cooling, as shown in Fig. 4(a), this temperature mismatch problem is avoided as schematically explained in Fig. 5. Hereafter, a cycle with a
bypass flow cycle is called a “partial pre-cooling”
cycle. Partial pre-cooling is also achieved in the
200
Y. Kato et al. / Nuclear Engineering and Design 230 (2004) 195–207
Table 3
Cycle thermal efficiencies of He and CO2 cycles at reactor outlet temperature of 800 ◦ C
Cycle thermal efficiency (%)
Items
No partial pre-cooling
No partial pre-cooling
Partial pre-cooling
One intercooler
Two intercoolers
7 MPa
12.5 MPa
7 MPa
12.5 MPa
7 MPa
12.5 MPa
45.3
45.8
49.2
45.1
45.5
49.7
47.5
47.8
51.4
47.4
47.4
51.9
48.4
48.7
–
48.3
48.1
–
High Pressure
Compressor
Bypass
Low Pressure
Compressor
Compressor
Low Pressure Bypass Turbine
Compressor Compressor
Reactor
Reactor
Pre-Cooler
Pre-Cooler
Generator
Recuperator
Turbine
Generator
Intercooler
He
CO2
No intercooler
Recuperator
(a)
(b)
Fig. 4. Partial pre-cooling in respective cycles. (a) No intercooler, (b) one intercooler.
Temp.
Reactor
Turbine
A
E
High pressure
& low temp.
side
D
Low pressure
& high temp.
side
C
B
Pre-cooler
Entropy
T- s Diagram in a recuperator
Larger Cp value at low pressure &
high temperature side results in
temperature mismatch (∆ T C-D < ∆ T A-B )
not maximizing cycle efficiency .
Bypassing gas flow from B to E
before a pre-cooler leads to
∆ T C-D = ∆T A-B
improving
cycle
efficiency
by
about 4%-6% .
Fig. 5. Bypass flow to eliminate temperature mismatch problem in CO2 cycles. (A) Recuperator inlet of the low-pressure and high-temperature
side, (B) recuperator outlet of the low-pressure and high-temperature side, (C) recuperator inlet of the high-pressure and low-temperature
side, (D) and (E) recuperator outlet of the high-pressure and low-temperature side.
Y. Kato et al. / Nuclear Engineering and Design 230 (2004) 195–207
201
60
Reactor Outlet Pressure
20.0MPa
15.0MPa
10.0MPa
5.0MPa
Cycle Efficiency (%)
CO2 (827˚C)
50
He (827˚C)
40
CO2 (527˚C )
30
He (527˚C)
Pre-Cooler Outlet Temperature : 35˚C
Compressor, Turbine Efficiency : 90%
Effectiveness of Recuperator : 95%
20
0
0.1
0.2
0.3
0.4
Bypass Flow Ratio (-)
Fig. 6. Cycle efficiency change with bypass flow fraction.
65
60
55
Cycle Efficiency (%)
one-intercooler cycle by bypassing flow to the
third compressor before the pre-cooler, as shown in
Fig. 4(b).
The optimum bypass flow fraction depends on turbine inlet temperature and pressure, as shown in Fig. 6.
As expected from the constant Cp of He, bypass flow
is not necessary for the He cycle, but it degrades cycle
efficiency. Bypass flow reduces heat rejected to cooling water through the pre-cooler and increases compressor work. Taking account of the opposing effect
on cycle efficiency resulted from the bypass flow, cycle efficiency is improved by about 4% at 800 ◦ C, as
shown in Table 3.
Fig. 7 shows that cycle efficiency of the CO2
partial pre-cooling cycle is about 3–9% higher than
that of the conventional He Brayton cycle, depending on turbine inlet temperature and pressure. At
medium core outlet temperature of 650 ◦ C and turbine inlet pressure of 7 MPa, the CO2 cycle achieves
cycle efficiency of 45.8%. This cycle efficiency value
is comparable with that of a typical He cycle of
GT-MHR (47.7%) at high temperature of 850 ◦ C and
7 MPa, as shown in Fig. 8. Its higher efficiency is
ascribed to reduced compression work around the
critical point of CO2 as explained below and consideration of variation in CO2 specific heat with pressure and temperature. Typical plant data are given in
Fig. 9.
P re-C ooler T em p eratu re : 35˚C
C om p ressor E fficien cy
: 90%
T u rb in e E fficien cy
: 90%
R ecu p erator E ffectiven ess: 95%
C O 2 P artial
P re-C oolin g
C ycle
50
He Cycle
45
Reactor Pressure
40
20.0 M P a
15.0 M P a
10.0 M P a
5.0 M P a
35
30
500
600
700
800
900
1000
1100
Reactor Outlet Temperature (˚C)
Fig. 7. Cycle efficiencies of a CO2 partial pre-cooling cycle and
a He cycle in a one-intercooler system.
202
Y. Kato et al. / Nuclear Engineering and Design 230 (2004) 195–207
Cycle Thermal Efficiency (%)
60
CO2 Gas Turbine
Cycle (Direct)
55
MTGR
Partial Pre-Cooling
Cycle
(650˚C, 45.8%)
50
HTGR
GT-MHR
(850˚C, 47.6%)
45
Water/Steam
Cycle (Indirect)
40
HTGR
Fort St. Vrain
(538˚C, 40.6%)
35
He Gas Turbine
Cycle (Direct)
LWR (Avg. 278˚C, about 34%)
30
200
400
600
800
1000
Turbine Inlet Temperature (˚C)
Fig. 8. Comparison of cycle efficiencies among MTGRs, direct cycle HTGRs and an indirect cycle HTGR.
Work W of one mole of real gas in isentropic expansion and compression processes is calculated as
dP
W = − V dP = − zRT ,
P
T: 99.7
P: 7.17
H: 827.7
T: 35.0
P: 3.41
H: 783.7
T: 480.2
P: 7.11
H: 1265.3
T: 35.0
P: 2.02
H: 798.9
T: 79.5
P: 3.43
H: 832.4
T: 107.2
P: 2.04
H: 870.0
Intercooler
Reactor
PV = zRT,
T: Temperature (˚C)
P: Pressure (MPa)
H: Specific Enthalpy (kJ/kg)
High
Bypass
Low
Pressure Pressure Compressor Turbine
CompressorCompressor
Generator
Pre-Cooler
T: 650.0
P: 7.00
H: 1470.2
where V = gas volume, P = gas pressure. For real
gases, an equation of state is written as
Recuperator
T: 243.9
P: 2.07
H: 1009.2
T: 502.2
P: 2.10
H: 1296.6
T: 229.5
P: 7.14
H: 977.9
Reactor Outlet Temperature: 650˚C
Reactor Outlet Pressure: 7.0 MPa
Pre-Cooler & Intercooler Outlet
Temperature: 35˚C
Turbine & Compressor Efficiency: 90%
Recuperator Effectiveness: 95%
Bypass Flow Fraction: 7.3%
↓
Cycle Efficiency: 45.8%
Fig. 9. Plant data of medium temperature CO2 partial pre-cooling cycle with one intercooler.
Y. Kato et al. / Nuclear Engineering and Design 230 (2004) 195–207
where R = gas constant, and z = compressibility factor. The compressibility factor represents the departure
from the ideal gas and is defined by
z=
At the critical temperature and pressure, the z value
dips sharply below the ideal line of unity and takes an
extremely low value as low as about 0.2 as shown in
Fig. 10. Factor z gives fractional deviation of real gas
from the ideal gas (z = 1) in isentropic expansion and
compression processes. A low z value indicates that
the gas is more compressible than the ideal gas. Compressor inlet temperature is usually around 35 ◦ C in
gas cooled reactors which is determined by available
cooling water temperature. Since compressor inlet
temperature is very close to the CO2 critical temperature (31.0 ◦ C), z values of CO2 in the compressing
condition are estimated to be smaller than those of
He, as seen from critical parameters. Both CO2 and
He z values approach unity at high-temperature and
high-pressure conditions in the turbine, so the difference is small. Judging from z values in compression
and expansion, CO2 cycle efficiency is expected to
be higher than that of the He cycle.
Cycle efficiency is dependent on the lowest temperature that appears usually at the inlet of compressors in
a Brayton cycle. Cycle efficiencies are plotted against
the compressor inlet temperature in Fig. 11, increasing linearly with temperature decrement at the rate of
1.3%/10 ◦ C.
Corrosion resistant structural materials and reliable
components in the CO2 environment can be used at
RT
.
PV
If P, V and T are expressed in terms of the respective reduced pressure (Pred = P/Pcrit , Pcrit is critical
pressure), reduced volume Vred (Vred = V /Vcrit , Vcrit is
critical volume) and reduced temperature Tred (Tred =
T /Tcrit , Tcrit is critical temperature), the above equation is rewritten as
Pred Vred
Pcrit Vcrit
.
z=
R Vcrit
Tred
The term of (Pcrit Vcrit /RVcrit ) is known to be approximately constant for many gases; for that reason z
appears to be a universal function of Pred and Tred
because of the law of corresponding states. If z is
plotted against Pred and Tred , a single curve will be
obtained for all gases. Gases with an equal z value
display the same behavior according to the law of
corresponding states. The compressibility factor z at
various Tred is plotted against Pred shown in Fig. 10
using the data given by Hougen et al. (1960). At these
pressure and temperature conditions, the fit is good
to within about 1% for gases (Moore, 1972).
1.6
Reduced Temperature T red =1.0
1.1 1.2 1.6
2
1.4
3
4
1.4
Compressibility Factor z
203
6
1.2
8
10
15
1.0
2
He Compression Condition
1.6
1.4
0.8
0.6
Higher cycle efficiency could be attained in
CO2 cycles compared with He cycles by
utilizing reduced compression work around
the critical point.
1.2
0.4
1.1
0.2
Drawn from the data in O.A.Hougen, et al.,
"Chemical Process Principles,Part II,
Thermodynamics", John Wiley & Sons
1.0
0.0
0
2
4
6
8
10
12
14
16
18
20
22
24
Reduced Pressure P red
Fig. 10. Compressibility factor with reduced temperature and reduced pressure.
204
Y. Kato et al. / Nuclear Engineering and Design 230 (2004) 195–207
Cycle Efficiency (%)
56
54
52
Core Outlet Temperature = 800˚C
50
48
46
Core Outlet Temperature = 650˚C
44
42
25
30
35
40
System Lowest Temperature (˚C)
Fig. 11. System lowest temperature dependency on cycle efficiency
in CO2 partial pre-cooling cycle with one intercooler.
the medium temperature of 650 ◦ C, as has been proven
in AGRs during extensive operation with core outlet
temperature of 650 ◦ C. The AGR graphite moderator
was prevented from excessive oxidation reaction with
CO2 by adding methane and carbon monoxide added
to CO2 in a range of small concentrations in CO2
(Hewitt and Collier, 1997). The average plant availability of AGRs over the last 10 years is as high as
that of current LWRs (Knox, 1999).
Regarding anti-oxidation of the graphite, surface
coating with inert materials would also be promising:
it has been developed for oxidation prevention of cutting tools up to 1000 ◦ C in air atmosphere. Graphite
balls coated on surface with TiSiN, TiCN + TiN and
TiCN + Al2 O3 + TiN are shown in Fig. 12. Fracture
cross section SEM micrographs of TiCN + TiN coating on the graphite ball are shown in Fig. 13. Their irradiation testing in the JMTR (Japan Material Testing
Reactor) core has been planned.
Any type of fuel can be applicable to CO2 gas turbine cycles with partial pre-cooling such as a conventional metal cladding fuel of the type used for LWRs
and LMFRs, or a pebble bed or block fuel of the
type used for HTGRs. In the case of the conventional
fuel type encapsulated with type 316 stainless steel
cladding material, encapsulation completely (except in
the case of cladding failure) prevents release of fission
products (FPs) from the fuel to coolant and consequent contamination of turbomachinery; this reduces
radiation dosage incurred through maintenance.
Leakage of a small amount of FPs is unavoidable in
the case of particle fuel for the pebble and block type
fuel coated with porous carbon, pyro-carbon, and SiC.
The leak fraction of FPs to total quantity produced
in the fuel depends on their diffusion coefficients in
pyro-carbon and SiC. Recent irradiation experiments
(IAEA, 1997) for the coated particle fuel show that
diffusion coefficients of typical FP elements such as
cesium, strontium, silver, and iodine are evaluated
from Arrhenius plots to be lower by about two orders
of magnitude in the medium temperature of 650 ◦ C
than at the high temperature of 850 ◦ C. These results
may lead to considerable reduction of radiation dosage
Fig. 12. Graphite balls with coated surfaces.
Y. Kato et al. / Nuclear Engineering and Design 230 (2004) 195–207
205
Fig. 13. Fracture cross section SEM micrographs of TiCN + TiN coating on graphite ball surfaces.
during maintenance of turbomachinery and heat
exchangers.
Various heat-resistant materials have been used for
HTGRs, as shown in Table 4: Hastelloy XR, Alloy
800H, and 2.25 Cr–1 Mo steel. Instead of these materials, less-expensive materials, such as type 316 stainless steel, can be applied to a medium temperature
reactor which has been proven to be compatible with
CO2 up to medium temperature of 650 ◦ C in AGRs,
and used for LWRs and LMFRs for long periods.
Thus, the medium temperature offers more flexibility
for structural material choice. Lowering temperature to
650 ◦ C not only eases maintenance through the lower
diffusion leak rate of fission products from coated
Fig. 14. Bird’s-eye view of a CO2 gas turbine reactor.
206
Y. Kato et al. / Nuclear Engineering and Design 230 (2004) 195–207
Table 4
Typical material selection in HTGRs and MTGRs
Positions
Core barrel, control rod
drive mechanism
Core support plate
Pressure vessel liner
Typical materials
HTGRs (850 ◦ C)
MTGRs (650 ◦ C)
Alloy 800H
316 SS (proven
in AGR)
2.25 Cr–1
Mo steel
Hastelloy XR
particle fuel, but also provides flexibility in choosing
materials.
All components (core, turbine, recuperator, pump,
and pre-cooler) are installed into pressure vessels as
typically illustrated in Fig. 14. Number (N) of turbine
stages is given as
N = Cp
T
,
hstage
where Cp = specific heat at constant pressure,
T = temperature drop in a turbine, hstage =
enthalpy drop per turbine stage.
It is seen from the above equation that the number
of the turbine stages (or turbine length) is proportional
to the Cp value at the same T and hstage conditions.
Values of T and hstage are determined by optimization of the cycle design and the turbine diameter, respectively. Because the specific heat value of CO2 at
constant pressure is smaller by a factor of about five
He
than that of He, number of stages (or length of turbine) is smaller at the same turbine diameter and then
volume (or approximately weight) of a CO2 cycle is
less by a factor of about five than that of a He cycle.
We did a comparative design study of turbomachinery
(turbine and compressor) between He and CO2 cycles
optimizing the turbine length and diameter. The optimization results shown in Fig. 15 indicate that turbine
length is 44% [= (0.81/1.84) × 100] and turbine diameter is 65% [= (1.25 + 1.33)/(1.94 + 2.01) × 100]
in a CO2 cycle compared with those of a He cycle.
Consequently, the CO2 cycle system offers one-fifth
(19%) volume (or approximate weight) of CO2 cycle
reference to that of He cycle (Muto et al., 2003). Judging from the key component size and cycle efficiency,
power generation cost per unit electricity for the CO2
cycle design might be less than that for a He cycle
design.
A gas turbine is a much simpler system than a steam
turbine because it has no moisture separators, steam
extraction systems, or coolant purity control systems
required for steam turbine cycles. Furthermore, it has
much smaller size because of its lower turbine pressure ratio of about four. Our proposed direct cycle has
only a single coolant circuit because of its direct cycle,
whereas PWRs have primary and secondary cooling
circuits with huge steam generators because of their
indirect cycle. Consequently, our CO2 gas turbine direct cycle reactor is much simpler and smaller than
PWRs.
CO2
1.84m
0.81m
1.94m
1.80m
2.01m
1.25m
1.10m
1.00m
Fig. 15. Comparison of turbine size between He and CO2 cycles.
1.33m
Y. Kato et al. / Nuclear Engineering and Design 230 (2004) 195–207
3. Conclusions
A medium temperature (650 ◦ C) CO2 gas turbine reactor with a partial pre-cooling cycle attains
comparable cycle efficiency of 45.8% with a typical
high temperature (850 ◦ C) He gas turbine reactor of
GT-MHR (47.7%). Lowering temperature to 650 ◦ C
provides flexibility in choosing materials and eases
maintenance through the lower diffusion leak rate
of fission products from coated particle fuel. The
medium temperature CO2 gas turbine reactor is expected to be a practical alternative to current high
temperature He gas turbine reactors.
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