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\begin{document}

\title{\bf ARIES-ST Safety Design and Analysis}

\author{H.~Y.~Khater, E.~A.~Mogahed, \\
Fusion Technology Institute, University of Wisconsin-Madison, WI \\
D. K. Sze
Argonne National Laboratory, Argonne, IL \\
M.~S.~Tillack, X.~R.~Wang, \\
Fusion Energy Research Program,
University of California, San Diego, CA \\
and the ARIES Team
}

\date{}
\maketitle

\begin{abstract}

Activation and safety analyses were performed for the ARIES-ST design.  The
ARIES-ST power plant includes a water-cooled copper centerpost.  The first
wall and shield are made of low activation ferritic steel and cooled with
helium.  The blanket is also made of ferritic steel with SiC inserts and
Li$_{17}$Pb$_{83}$ breeder.  The divertor plate is made of low
activation ferritic steel and uses a tungsten brush as plasma facing component
(PFC).  The power plant has a lifetime of 40 full power years (FPY).  However,
the centerpost, first wall, inboard shield and blanket were assumed to be
replaced every 2.86~FPY.  Neutron transmutation of copper resulted in the
production of several nickel, cobalt and zinc isotopes.  The production of
these isotopes resulted in an increase of the time-space average electrical
resistivity of the centerpost by about 6\% after 2.86~FPY.  All of the plant
components met the limits for disposal as Class C low-level waste (LLW).  The
off-site doses produced at the onset of an accident are caused by the
mobilization of the radioactive inventory present in the plant.  Analysis of a
loss of coolant accident (LOCA) indicated that the centerpost would reach a
maximum temperature of about 1000\SdegC during the accident.  In the meantime,
the first wall and shield would reach a maximum temperature of about 800\SdegC.
A similar divertor LOCA analysis indicated that the front tungsten layer would
also reach a maximum temperature of about 800\degC.  The calculated temperature
profiles and available oxidation-driven volatility experimental data were used
to calculate the dose at the site boundary under conservative release
conditions.  The current design produces an effective whole body early dose of
1.88~mSv at the site boundary.  In addition, a divertor disruption would only
produce an effective whole body early dose of 7.68~$\mu$Sv at the site
boundary.

\end{abstract}

\newpage

\section{Introduction}
\label{s-1}

ARIES-ST is a low-aspect-ratio spherical torus power plant, which is one of
several fusion power plant designs being assessed within the ARIES project.
The ARIES-ST power plant~\cite{Naj00} produces a 1000~MW of net electric power
and is assumed to operate for 40 full power years.  The plant includes a
water-cooled, DS GlidCop Al-15 copper centerpost.  The first wall and shield
are made of low activation ferritic steel (9Cr-2WVTa) and cooled with
helium~\cite{Tillack00}.  The blanket is also made of ferritic steel with SiC
inserts and Li$_{17}$Pb$_{83}$ breeder.  Figure~\ref{ast-6-01} shows an
elevation view of the ARIES-ST power core.  As shown in Fig.~\ref{ast-6-02},
the design utilizes a divertor with a steel back plate and uses a tungsten
brush as plasma facing component (PFC).  Activation analysis was performed
assuming average neutron wall loadings of 2.1 and 4.6~MW/m$^2$ at the inboard
and outboard first wall, respectively.  In addition, the average neutron wall
loading at the inner divertor plate is 0.9~MW/m$^2$.  The neutron wall loadings
are the limiting factors for the lifetime of the different components of the
power plant.  The centerpost, first wall, inboard shield and blanket were
assumed to survive for 2.86~FPY~\cite{ELG99}.  The divertor plate and manifold
were assumed to have a lifetime of 2.86~FPY.  The divertor shield was assumed
to stay in place for 40~FPY due to the fact that it is exposed to a lower
neutron environment.  The activity and decay heat were calculated as a function
of time following shutdown.  Calculation of the structure activity is needed to
evaluate the potential impact of a radioactive inventory release at the onset
of an accident.  Results of the decay heat calculation are used to examine the
thermal response of the plant structure following a loss of coolant accident
(LOCA).  The waste disposal ratings (WDR) of the different plant components at
the end of their lifetimes were also evaluated.  The waste disposal rating is
needed to determine if a given structure component would satisfy the regulatory
criteria for shallow land burial as low- level waste (LLW).

A strong emphasis has been given to the environment and safety issues in the
ARIES-ST power plant design. Low activation ferritic steel (9Cr-2WVTa) has been
used in the blanket and shield to avoid a high level of induced
radioactivity. Similarly, the use of Li$_{17}$Pb$_{83}$ as a breeder eliminates
the hazard posed by the energy producing chemical reactions usually associated
with the use of lithium and hence reduces the risk of mobilizing the
radioactive inventory present in the plant. The methodology used in this
analysis dose not depend on the probability of accident initiating
scenarios. In accordance with previous ARIES safety analyses, we have rather
adopted the principle of considering a severe accident scenario, while
maintaining integrity of all containment barriers during such an accident.  To
evaluate the possible radiological hazard to the public, we used a two step
approach in calculating the possible off-site dose.  The first step in our
approach is the identification of the sources and locations of the radioactive
inventories inside the plant building.  However, since the existence of
radioactivity does not in itself represent a safety hazard, the second step in
our approach was to consider a set of pessimistic accident scenarios for
mobilizing and releasing the radioactive inventory.  A major goal of the
ARIES-ST design has been achieving the highest level of safety while
maintaining its economic attractiveness.  Taking this into account, the design
aimed at achieving the following goals:
%
\begin{enumerate}
\item
Minimizing the increase in the centerpost resistivity.

\item
Disposal of the plant structure as low-level waste.

\item
Significantly reducing the off-site doses from routine release of tritium
during operation.

\item
Minimizing the level of off-site doses at the site boundary following an
accident.

\end{enumerate}

In this paper, it is shown that the first goal could be achieved by adequate
shielding of the centerpost.  The other goals could be achieved by using low
activation materials in the first wall and shield of the plant and proper
design of the heat exchanger to minimize the amount of routine tritium release.
The 9Cr-2WVTa ferritic steel was selected because it produces a low level of
long-term radioactivity and acceptable levels of short and intermediate-term
radioactivity.  The disposal of the structure as low-level waste is dependent
on producing low levels of long-term radioactivity.  On the other hand,
off-site doses during an accident are dominated by nuclides with short and
intermediate lifetimes.  In addition, nuclides with intermediate lifetimes are
the major contributors to the decay heat and hence, the temporal variation of
the structure temperature during an accident.

\section{Calculational Procedure}
\label{s-2}

The neutron flux used for the activation calculations was generated by the one-
dimensional discrete ordinates neutron transport code ONEDANT~\cite{Odell82}.
A 46-group neutron and 21-group gamma coupled cross section library, based on
FENDL1.0~\cite{MACF95}. The analysis uses a $P_3$ approximation for the
scattering cross sections and $S_8$ angular quadrature set.  The plant
structure calculations used toroidal cylindrical geometry models with the
inboard and outboard sides modeled simultaneously.  The average neutron wall
loading on the inboard and outboard sides are 2.1 and
4.6~MW/m$^2$~\cite{ELG99}, respectively.  The average neutron wall loading at
the inner divertor plate is 0.9~MW/m$^2$.  The activation analysis was
performed using the latest version of the activation code
DKR-PULSAR2.0~\cite{DKP-PUL}. The code combined the neutron flux with the
FENDL/A-2.0 data library~\cite{Pash97} to calculate the activity and decay heat
generated in the different regions of the plant.  The plant was assumed to
operate continuously for 40 FPY. The centerpost, first wall, inboard shield and
blanket were assumed to be replaced every 2.86~FPY.  The divertor plate and
manifold were assumed to have a lifetime of 2.86~FPY.  The divertor shield was
assumed to stay in place for 40~FPY. The decay gamma source produced by the
DKR-PULSAR2.0 code was used with the adjoint neutron flux to calculate the
biological dose rates after shutdown using the DOSE code~\cite{DKP-PUL}. The
dose rate calculations were performed at different locations inside the
containment.  The structure activation results were utilized in a radwaste
classification.  The decay heat results were used in a loss of coolant accident
analysis.  The structure and the Li$_{17}$Pb$_{83}$ breeder activation results
were used in the off-site dose calculations following the LOCA.  The activation
results have also been utilized in the off-site dose calculation. The off-site
doses are produced by the accidental release of the radioactive inventory from
the containment building assuming the worst case weather conditions.  Finally,
the EPA code AIRDOS-PC~\cite{AIRDOS89} has been used to estimate the off-site
dose due to the routine release of tritium.  The materials used in the
different regions of the plant are presented in Table~\ref{t-1}.

\section{Activation Analysis}
\label{s-3}

\subsection{Activity and decay heat}
\label{s-3-1}

After 2.86~FPY, the total activity induced in the centerpost is 4430~MCi and
drops to 994~MCi in one day and 2.7~MCi in one week.  In the meantime, the
total decay heat induced in the centerpost is 12.25~MW and drops to 1.87~MW
within a day.  The decay heat induced in the centerpost is dominated by the two
copper isotopes $^{64}$Cu ($T_{1/2}=12.7$~h) and $^{66}$Cu
($T_{1/2}=5.1$~m). The dominant nuclides at one week following shutdown are
$^{60}$Co ($T_{1/2}=5.27$~y) and $^{64}$Cu.  The total activity induced in the
inboard shield and first wall at shutdown is 959~MCi and drops to 594~MCi in
one day and 501~MCi after one week.  At shutdown, the decay heat amounts to
5.14~MW and drops to 0.56~MW within a week.  The decay heat at shutdown is
dominated by $^{56}$Mn ($T_{1/2}=2.578$~h) and $^{187}$W ($T_{1/2}=23.9$~h).
At 1 day following shutdown, the decay heat induced in the inboard shield and
first wall is dominated by $^{54}$Mn ($T_{1/2}=312.2$~d) and $^{56}$Mn.

The amounts of activity and decay heat induced in the outboard first
wall/blanket structure and shield after 2.86~FPY are 2,800~MCi and 19.2~MW,
respectively.  The activity and decay heat drop to 1,552~MCi and 1.25~MW within
a week from shutdown.  Most of the activity and decay heat are induced in the
ferritic steel part of the structure and dominated by the same nuclides as in
the inboard first wall and shield.  The amount of activity induced in the
Li$_{17}$Pb$_{83}$ blanket at shutdown is 5,252~MCi.  Due to the rapid decay of
$^{207m}$Pb ($T_{1/2}=0.8$~s), the activity of Li$_{17}$Pb$_{83}$ drops to only
140~MCi within a day from shutdown.

The tungsten plasma facing part of the divertor has a total induced activity
and decay heat after 2.86~FPY of~546 MCi and~0.8 MW, respectively.  The
dominant nuclides at shutdown are $^{187}$W ($T_{1/2}=23.9$~h), $^{185}$W
($T_{1/2}=74.8$~d), and $^{183m}$W($T_{1/2}=5.15$~s).  After 1 day following
shutdown $^{187}$W, $^{185}$W, and $^{181}$W ($T_{1/2}=121.2$~d) dominate the
activity and decay heat induced in the tungsten layer.  The amount of activity
generated in the divertor back plate at shutdown is 129~MCi and decays to
79~MCi within a week.  The decay heat generated in the back plate drops from
0.69~MW to 0.082~MW during the first day following shutdown.

\subsection{Change in centerpost electrical resistivity}
\label{s-3-2}

Interactions between high-energy neutrons and the copper centerpost lead to the
production of several nickel, cobalt, and zinc isotopes as transmutation
products.  Production of these isotopes leads to an increase in the centerpost
electrical resistivity.  The centerpost resistivity increases linearly with
the increase in time of operation.  Increase in the centerpost resistivity
would lead to an increase in the recirculating power and lower net efficiency.
The following are the most important reactions:
%
\begin{itemize}

\item
$^{63}$Cu (n,2n) $^{62}$Cu($\beta^+$) $^{62}$Ni, \hspace{1cm}
$^{65}$Cu (n,2n) $^{64}$Cu($\beta^+$)$^{64}$Ni,

\item
$^{63}$Cu (n,$\gamma$) $^{64}$Cu($\beta^+$)$^{64}$Ni, \hspace{1.35cm}
$^{63}$Cu (n,$\gamma$) $^{64}$Cu($\beta^-$)$^{64}$Zn,

\item
$^{64}$Cu (n,2n) $^{64}$Cu($\beta^-$)$^{64}$Zn, \hspace{1.1cm}
$^{65}$Cu (n,$\gamma$) $^{66}$Cu($\beta^-$)66Zn

\item
$^{63}$Cu (n,$\gamma$) $^{64}$Cu(n,$\alpha$) $^{60}$Co, \hspace{1cm}
$^{65}$Cu (n,2n) $^{64}$Cu(n,$\gamma$) $^{60}$Co

\end{itemize}

Figure~\ref{ast-6-03} shows the radial distribution of the percentage increase
in the centerpost resistivity.  As shown in the figure, the outermost 30~cm of
the 80-cm-thick centerpost exhibits the bulk of transmutation.  This is due to
the fact that the change in the copper transmutation is mostly due to the
production of the $^{64}$Ni isotope.  As already shown, $^{64}$Ni is mostly
produced via high-energy threshold reactions.  The electric current will
redistribute within the centerpost to avoid the region with high resistivity.
The space-time average increase in resistivity over the entire centerpost is
about 6\%, which is considered a tolerable value.

\subsection{Biological dose rates}
\label{s-3-3}

Biological dose rate calculations were performed at selected locations in the
space between the centerpost and the back of the inboard shield, and in the
space between the outboard blanket and the magnet.  Figure~\ref{ast-6-04} shows
the calculated dose rates as a function of time following shutdown.  At
shutdown, the biological dose rate in the space between the centerpost and the
back of the inboard shield is $4.57\times 10^7$~mSv/h.  The dose is dominated
by the $^{56}$Mn ($T_{1/2}=2.578$~h) and $^{187}$W ($T_{1/2}=23.9$~h) isotopes
produced from the ferritic steel component of the inboard shield, and $^{66}$Cu
($T_{1/2}=5.1$~m) produced from the copper centerpost.  One day following
shutdown the dose rate drops to $1.05\times 10^7$~mSv/h.  The two isotopes,
$^{187}$W ($T_{1/2}=23.9$~h) and $^{182}$Ta ($T_{1/2}=114.43$~d), dominate the
dose rate at that time. $^{187}$W and $^{182}$Ta are produced in the inboard
shield.  The dose rate continues to be high for years following shutdown.
After one year, the biological dose rate is $1.27\times 10^6$~mSv/hr and is
dominated by $^{60}$Co ($T_{1/2}=5.27$~y) produced from the centerpost, and
$^{54}$Mn ($T_{1/2}=312.2$~d) and $^{182}$Ta produced from the inboard shield.

As shown in the figure, the biological dose rates in the space between the
outboard blanket and the magnet continues to be high.  For the most part these
dose rates are also dominated by the same nuclides mentioned before.  However,
all of these nuclides are generated in the ferritic steel component of the
outboard blanket.  Since the high biological dose rates continue for several
years following shutdown, only remote maintenance is feasible anywhere inside
the ARIES-ST building.

\subsection{Radwaste classification}
\label{s-3-4}

The radwaste of the different regions of the plant were evaluated according to
both the NRC 10CFR61~\cite{10CFR61} and Fetter~\cite{Fetter90} waste disposal
concentration limits (WDL). The 10CFR61 regulations assume that the waste
disposal site will be under administrative control for 100 years.  The dose at
the site to an inadvertent intruder after the 100 years is limited to less than
500~mrem/year.  The waste disposal rating (WDR) is defined as the sum of the
ratio of the concentration of a particular isotope to the maximum allowed
concentration of that isotope taken over all isotopes and for a particular
class.  If the calculated $\mbox{WDR} = 1 $ when Class A limits are used, the
radwaste should qualify for Class A segregated waste.  The major hazard of this
class of waste is to individuals who are responsible for handling it. Such
waste is not considered to be a hazard following the loss of institutional
control of the disposal site. If the $\mbox{WDR} > 1$ when Class A WDL are used
but $\mbox{WDR} \le 1$ when Class C limits are used, the waste is termed Class
C intruder waste.  It must be packaged and buried such that it will not pose a
hazard to an inadvertent intruder after the 100 years institutional period is
over. Class C waste is assumed to be stable for 500 years. Using Class C
limits, a $\mbox{WDR} > 1$ implies that the radwaste does not qualify for
shallow land burial. Fetter developed a modified version of the NRC's intruder
model to calculate waste disposal limits for a wider range of long-lived
radionuclides which are of interest for fusion researchers than the few that
currently exist in the current 10CFR61 regulations. Fetter's model included
more accurate transfer coefficients and dose conversion factors. However, while
the NRC model limits the whole body dose to 500~mrem or the dose to any single
organ (one of seven body organs) to 1.5~rem, Fetter limits are based on the
maximum dose to the whole body only.

Specific activities calculated by the DKR-PULSAR2.0 code were used to calculate
the waste disposal ratings.  The waste disposal ratings for the 10CFR61 and
Fetter limits are shown in Tables~\ref{t-2} and \ref{t-3}.  Results in the
tables are given for compacted wastes.  Compacted waste corresponds to crushing
the solid waste before disposal and thus disallowing artificial dilution of
activity. The Class C WDR values were calculated after a one-year cooling
period.  As shown in Table~\ref{t-2}, according to the 10CFR61 limits, the
centerpost WDR is dominated by $^{63}$Ni, which is produced via the
$^{63}$Cu(n,p) reaction.  Since this reaction is a high-energy threshold
reaction, providing extra shielding on the inboard side could reduce the amount
of $^{63}$Ni generated in the centerpost further.  On the other hand,
$^{94}$Nb, produced from the 0.5~wppm niobium impurities in the 9Cr-2WVTa
steel, is the dominant source of waste hazard in the first wall, blanket and
shield.  As shown in Table~\ref{t-3}, $^{108m}$Ag produced from the 20~wppm
silver impurities contained in the GlidCop Al-15 copper alloy, is the major
waste hazard in the centerpost according to Fetter limits.  In addition to
$^{94}$Nb, $^{192m}$Ir is the other waste hazard associated with ferritic
steel.  These results show that the waste classification of the centerpost is
controlled by its 10CFR61 WDR as it is entirely due to direct transmutation of
copper rather than impurities included in the Cu alloy.  All other WDR could be
limited by controlling the level of impurities in the copper and steel alloys
regardless of the waste disposal limits used.  Similarly, as shown in the
tables, the niobium and silver impurities in the divertor tungsten PFC and
ferritic steel back plate are entirely responsible for the waste hazard of the
divertor plate.

\section{Loss of coolant accident analysis}
\label{s-4}

The goal of this analysis is to determine the temperature history of the
different plant components as a function of time following a LOCA.  LOCA occurs
when one or more supply tubes outside the plant are damaged or ruptured,
preventing the coolant from reaching the first wall or plasma facing
components.  Even though neutron heating is absent, the lack of coolant in the
coolant channels causes the temperature to rise in various sections of the
blanket and shield due to decay heat.  For ARIES-ST, it is assumed that the
plasma is immediately quenched and the chamber components begin to increase in
temperature due to the decay heat generated.  Due to the large difference
between the time scale of plasma shutdown $\approx 1 ms$ and the loss of
coolant or loss of flow (several minutes-hours), it is assumed that the plasma
is immediately queched at the onset of the LOCA/LOFA and the chamber
components' tempeature begins to increase due to the decay heat generated
(worst case scenario).  The thermal response of the blanket, shield, and
divertor following a LOCA is determined by the analysis of an axisymmetric (R,
$\phi$) finite element model for the entire power plant (Fig.~\ref{ast-6-05}).
This analysis examines the thermal behavior of the in-vessel components to
determine the temperature profiles as a function of time following LOCA.

In this LOCA analysis, the centerpost is made of copper with empty coolant
channels.  The cross-sectional area of the empty coolant channels represents
15\% of the total centerpost cross-sectional area.  The same conditions are
applicable for the rest of the power plant.  The outer shell of the toroidal
field (TF) magnet is made of aluminum and is connected to the bus-bars at the
midplane.  This analysis allowed conduction up/down the centerpost, and in the
TF coil outer leg winding pack (WP).  The centerpost and the TF coil outer leg
winding pack are assumed to be connected to an ultimate heat sink at a
temperature of 35\degC.  The present centerpost/WP design is intimately
connected to very large bus-bars (an ultimate heat sink).  The centerpost is
connected to the TF coil outer legs at both ends of the centerpost.

\subsection{Methodology}
\label{s-4-1}

In order to perform the ARIES-ST LOCA analysis, several assumptions had to be
made.  The assumptions for the accident conditions were:
%
\begin{enumerate}

\item
Coolants are drained from all the blanket/shield sectors as well as the
divertor plate leaving the coolant channels empty and making it possible for
coolant channel surfaces to radiate to each other.

\item
The plasma is quenched instantaneously upon the onset of a LOCA.  An 
appropriate shutdown mechanism must be
in place for this to happen.  If the collant leaks into the chamber 
that would cause an immediate plasma
shutdown while the collant is still running.  On the other hand, if 
it takes 10 seconds for the plasma to be
deleberty shut down after the complete collant loss, the heat 
generated amounts to 1\% of the integrated
afterheat in the first week after LOCA.  Although this does not 
appear to be very much, the immediate effect
on the first wall during that 10 seconds could be significant.

\item
All facing surfaces radiate to each other.

\item
Both inboard and outboard blanket/shields are solved interactively
and can therefore radiate to each other across the plasma space.

\item
Heat transfer across the gaps is by thermal radiation and partially by thermal
conduction. In the gaps a certain percentage of the facing area is 
assumed to have
connecting structure for internal support and can, therefore, also conduct
heat.

\item
Thermal emissivity is taken as 0.5 for all surfaces.

\item
The massive copper bus-bars are considered as ultimate heat sinks at a constant
temperature of 35\degC.

\item
Temperature dependent thermo-physical properties are used for all the
materials.

\item
The material properties of the different zones in the radial build have been
adjusted (linearly proportional to volumetric ratio) to preserve the density,
heat capacitance and thermal conductivity of the actual composition.

\end{enumerate}

\subsection{Modeling and boundary conditions}
\label{s-4-2}

An axi-symmetric model is employed for finite element analysis.  The commercial
finite element code ANSYS 5.4 is used in this analysis~\cite{ANSYS}.  The
finite element model used in the analysis is shown in Fig.~\ref{ast-6-05}.  The
only thermal loads considered in this analysis are those generated as decay
heat at the onset of a LOCA.  The external surface of the plant is considered
as an adiabatic boundary.  The copper in the centerpost is thermally conducting
to the aluminum outer shell of the TF coils.  The TF outer shell is connected
to an ultimate heat sink through the massive conduction bars.  The average
decay heat variations in the outboard and inboard sides near the midplane are
shown in Fig.~\ref{ast-6-06} for the first week after shutdown.
Figure~\ref{ast-6-07} shows the change in the average decay heat behind the
inboard tungsten stabilizer plate at the inboard top and bottom.  One can
notice the effect of the additional shielding provided by tungsten on the decay
heat induced in some of the plant components.  The analysis also assumes a 1\%
conduction (1\% of the facing area is thermally in physical contact with the
other) in the gaps.  The gaps are used to maintain the physical separation
between the plant components at different temperature and they are also needed
during maintenance.  Table~\ref{t-4} shows a list of the initial thermal
conditions used in the analysis.

\subsection{Effect of thermal conductance through the gaps}
\label{s-4-3}

In the ARIES-ST design, the gaps separating the plant components are used to
maintain the physical separation between the different components operating at
different temperatures.  In addition, these gaps are also needed for the
assembly of the plant components.  Some physical contact between various
surfaces must exist to keep the surfaces shape and to transfer structural loads
between the different components.  The heat leakage from various surfaces
operating at different temperatures during normal operation is of a serious
concern.  A parametric analysis is performed to study the effect of thermal
conductance through the gaps on the thermal response of the in-vessel
components.  A simplified 2-D finite-element ($r,\theta$) model in cylindrical
coordinates is employed for thermal analysis at the midplane is shown in
Fig.~\ref{ast-6-08}.  Various levels of partial thermal conduction through the
gaps are used~\cite{Mogahed98}.  Figure~\ref{ast-6-09} shows the effect of
thermal conductance through the gaps on the maximum temperature of outboard
first wall.  As shown in the figure, the slope of the temperature is very steep
at lower values of conductance.  Therefore, to minimize the heat leakage during
normal operation, a reasonable value of 1\% is chosen for the thermal
conductance through the gaps.

\subsection{Results and discussion}
\label{s-4-4}

The temperatures of the various plant components were calculated as a function
of time following a LOCA by using the transient thermal loads due to the decay
heat, the assumed boundary conditions, and the initial temperatures of the
various ARIES-ST components.  The maximum temperature reached by the centerpost
is 1,018\degC.  The maximum temperature of the inboard first wall is about
887\degC and the maximum temperature of the inboard shield is 974\degC.  As
shown in Fig.~\ref{ast-6-10}, these maximum temperatures are reached after
18~hours from shutdown.  As shown in Fig.\ref{ast-6-11}, The maximum outboard
first wall and blanket temperatures are 878\degC and 727\degC, respectively.
The results showed elevated temperatures in the centerpost and an effective
passive heat removal solution might be needed.  Heat pipes can be used inside
the centerpost to transfer the generated decay heat to the colder massive ends
of the centerpost.  In such a case, it is required to determine how many heat
pipes are needed to reduce the maximum temperature of the centerpost and other
plant components.  Using heat pipes that occupy less than 5\% of the centerpost
total cross sectional area at the midplane would reduce the maximum temperature
of centerpost, inboard first wall, and the inboard shield to less than
823\degC, 803\degC, and 828\degC, respectively.  The maximum temperatures of
outboard first wall and blanket would be also reduced to less than 796\degC
and 682\degC, respectively~\cite{Mogahed00}.

Finally, Fig.\ref{ast-6-12} shows the temperature variation of the divertor
tungsten brush as a function of time following shutdown. The temperature drops
quickly because of the large amount of steel in the back of the divertor plate
that radiates to a colder surface at the top and the bottom of centerpost
(100\degC).

\section{Hazard assessment}
\label{s-5}

\subsection{Routine atmospheric effluents}
\label{s-5-1}

The radiological dose to the population in the vicinity of the power plant site
due to the routine release of tritium has been estimated by using the EPA
AIRDOS-PC code.  The code calculates the effective dose equivalent (EDE) as
mandated by 40 CFR 61.93 and 61.94 to the maximally exposed individual (MEI)
and at several distances from the point of release.  Dose values are computed
from ingestion, inhalation, air immersion, and ground surface pathways. The
routine releases from the several processing systems were based upon the
quantity of tritium processed per day and followed experience at TSTA, which
indicated that a barrier factor of 10$^{6}$ is an acceptable one.  Assuming the
release parameters listed in Table~\ref{t-5} and using meteorological
conditions at different cities, we calculated the dose expected at typical
locations near Boston, Chicago, Albuquerque and Los Angeles.  A summary of the
results is shown in Table~\ref{t-6}.  The worst dose was in the Los Angeles
area but was only 2.95~mrem/yr.  More than 85\% of the doses at all sites are
incurred via the ingestion pathway.  It is important to keep in mind that the
estimated dose values strongly depend on the stack height. For example, using a
35~meter stack height results in an EDE of 1.5~mrem/yr at the site boundary
(1~km) if the Los Angeles meteorological conditions were used.  Actually, the
rule of thumb for determining the necessary stack height is to use 2.5 times
the height of the nearest tall building in order to avoid downwash of the plume
into the wake of the building~\cite{NRC83}. A shorter stack must be justified
with appropriate analysis.

\subsection{Off-site dose calculations}
\label{s-5-2}

A strong emphasis was given to the environmental and safety issues in the
ARIES-ST design.  Low activation ferritic steel (9Cr-2WVTa) was used in the
first wall and shield to avoid generating high levels of induced radioactivity.
Similarly, the use of Li$_{17}$Pb$_{83}$ as a breeder reduces the hazard posed
by the energy producing chemical reactions usually associated with the use of
lithium and hence reduces the risk of mobilizing the radioactive inventory
present in the plant.  To evaluate the possible radiological hazard to the
public, a two-step approach was used in calculating the possible off-site dose.
The first step in the approach is the identification of the sources and
locations of the radioactive inventories inside the plant.  However, since the
existence of radioactivity does not in itself represent a safety hazard, the
second step in the approach was to consider a pessimistic accident scenario for
mobilizing and releasing the radioactive inventory.  The analysis assumed a
total loss of coolant accident as the worst case accident.  During the LOCA,
heat from all in-vessel components is transported to the massive copper
bus-bars.  In addition, in this analysis we also considered two possible
scenarios during which radioactivity could be released from the divertor PFC.
Following a disruption, off-site doses could be produced by direct vaporization
of the divertor surface layer or by the release of tokamak dust.  Tokamak dust
(particles smaller than 100~$\mu$m in diameter) is produced by the vaporization
of divertor surface material and accumulated during previous plasma
disruptions.

The radioactive inventory calculated by the DKR-PULSAR2.0 code was 
used in conjunction with the isotope-specific dose data
calculated by the MACCS code~\cite{Chanin90} to calculate effective 
whole body off-site dose inventory (dose caused by
100\% release of radioactivity) under worst release conditions. 
These conditions are ground release, atmospheric
stability class F, 1~km site boundary and 1~m/s wind speed.  Doses calculated
are produced through all of the following pathways:
%
\begin{itemize}
\item Inhalation of radionuclides during plume passage.
\item Inhalation of resuspended radionuclides.
\item External exposure to the plume.
\item External exposure from ground deposition.
\item Cloudshine or groundshine.
\item Ingestion of contaminated food.
\end{itemize}

\subsubsection{Off-site doses during a LOCA}
\label{s-5-2-1}

During a LOCA, a large increase in the structure temperature could lead to the
mobilization and partial release of the radioactive inventory.  As previously
discussed, the decay heat generated during the first day following a LOCA would
increase the center post temperature by $<1,000$\degC, the first wall/blanket
temperature by $<800$\degC, and the divertor temperature by $<800$\degC.  Under
these conditions, the full mobilization of the structure radioactive products
is impossible.  The highest temperature the structure would reach determines
the release fraction of its radioactive products (very conservative 
assumption).

Most of first wall/blanket radioactivity is generated in its steel component.
Off-site dose calculations were performed using ferritic steel experimental
volatility rates~\cite{Mccarthy94,Mccarthy95,Mccarthy96}.  Low activation
ferritic steel volatility rates at 700\degC in dry air were used in this
analysis.  To estimate conservative release fractions, a 24-hour LOCA was
assumed.  The previously mentioned temperatures were assumed to last 
for 24 hours and the voltilization rates were
those found in one-hour experiments where no protetctive oxide has 
formed. Similar analyses (24-hour LOCA) were performed
for the centerpost and the divertor's tungsten PFC using experimental 
volatility rates for copper (using copper volatility
rates at 1,000\degC in dry air) and tungsten (using tungsten 
volatility rates at 800\degC in dry air), respectively.

The two sources of radiological hazard in a Li$_{17}$Pb$_{83}$ blanket are
tritium and the activation products of Li$_{17}$Pb$_{83}$.  As shown in
Table~\ref{t-7}, The steady state tritium inventory in Li$_{17}$Pb$_{83}$ is
kept very low, on the order of 16~g, by its continuous removal during the plant
operation.  The activation products of major radiological hazard in a
Li$_{17}$Pb$_{83}$ blanket are the two isotopes, $^{203}$Hg ($T_{1/2}=46.61$~d)
and $^{210}$Po ($T_{1/2}=138.38$~d).  Both $^{203}$Hg ($\gamma$ and $\beta$
emitter) and $^{210}$Po ($\alpha$ emitter) are highly volatile
materials. Isotope $^{203}$Hg produces a high prompt bone marrow dose and
$^{210}$Po results in high values of prompt, early as well as chronic
doses. Isotope $^{210}$Po is produced via nuclear transmutation of bismuth and
is considered as the main safety hazard in Li$_{17}$Pb$_{83}$ blankets.
Bismuth is a major impurity of commercial lead and is also produced as a
transmutation product of Pb.  Commercial Pb contains 500--1,500~wppm of Bi and
high-purity Pb contains less than 10~wppm of Bi.  The Li$_{17}$Pb$_{83}$ used
in this analysis contains 43~wppm of Bi impurities.  It is desirable to keep
the Bi impurity in lead below 10~wppm.  The amount of Po generated can be
controlled by limiting the Bi impurities initially present in Pb as well as the
on-line continual removal of Bi atoms produced by neutron-lead reactions.
Fortunately, Po evaporates into the form of an intermetallic compound PbPo,
whose evaporation rates are very small because of the low vapor pressure of
this polonide~\cite{Wittenberg}.  Similarly, Hg evaporates into the form of an
intermetallic compound LiHg, whose evaporation rates are orders of magnitude
lower than Hg~\cite{Wittenberg}.  It is estimated that Po retention 
in a Li$_{17}$Pb$_{83}$ melt
is in the range of 96.4\% to 99.2\%~\cite{Hoffman85}.  In addition, under
accidental spill conditions, the dilution of Po is such that $\alpha$ and
$\gamma$ radiation will be shielded by the large amount of lead atoms
surrounding Po atoms.

A major advantage of using Li$_{17}$Pb$_{83}$ as a blanket is its low chemical
reactivity.  During an accident, a leak of water into the Li$_{17}$Pb$_{83}$
region will result in a chemical reaction between water  from the center 
post and the Li in the
molten Li$_{17}$Pb$_{83}$.  The reaction potential is much smaller than a
water/liquid Li reaction.  A Li$_{17}$Pb$_{83}$/water reaction tends to be
self-limiting due to the fact that the liquid metal is formed by 83\% Pb that
does not react with water and which after initial depletion of Li, tends to
shield the remaining amount of alloy from further interaction with
water~\cite{Wittenberg}.  In addition, solid products Li2O and LiOH are
produced and provide shielding for the remaining liquid metal from the rest of
the water.  The Li$_{17}$Pb$_{83}$/water reaction is an exothermic reaction,
which leads to an increase in temperature on the order of 200--400\degC.  A
complete reaction between water and Li$_{17}$Pb$_{83}$ would lead to the
production of 55.6~mole of H$_2$ per kg of water.  However, because the Li
oxidation is the source of H$_2$ production, no oxygen is present and therefore
explosion cannot occur~\cite{Ginacarli95}.

The off-site doses were calculated by combining the total off-site dose
inventory with the different volatility data under the LOCA condition discussed
previously.  Since no volatility data are available for Li$_{17}$Pb$_{83}$,
very conservative release rates were adopted for the release of $^3$H,
$^{203}$Hg and $^{210}$Po.  All of the $^3$H, 30\% of the $^{203}$Hg and 10\%
of the $^{210}$Po were assumed to mobilize during an accident.  Minor air
ingress into the coolant channel results in the volatilization of in-vessel
materials as previously discussed.  Once airborne, these particles could be
transported to the site boundary.  Assuming that the vacuum vessel and the
containment would stay intact during accidents, they would be expected to act
as release barriers.  For a vacuum vessel leak rate of 1\% per day, a
containment factor of 99\% could be considered.  Considering the vacuum and
containment boundaries as two independent barriers leads to an overall
radioactivity containment factor of 99.99\%~\cite{Taylor96,Petti94,Petti96}.
As shown in Table~\ref{t-8}, the LOCA based accident produces an effective
whole body early dose of 1.88~mSv at the site boundary.

\subsubsection{Off-site doses due to divertor disruption}
\label{s-5-2-2}

Assumptions used for the energy dissipated in the divertor during a disruption
are similar to assumptions used in the ITER safety
analysis~\cite{ITER97}. Following a disruption, off-site doses could be
produced by direct vaporization of the divertor surface layer. The total energy
dissipated during the energy quench phase is 2.1~GJ. During a disruption, very
small particles (0.1~$\mu$m) are mobilized by direct vaporization. Assuming an
interaction surface area of 15~m$^2$, the total amount of tungsten vaporized
during a disruption is 1~kg~\cite{Hoffman85}.  In addition, tokamak dust
(particles smaller than 100~$\mu$m in diameter) is produced by the vaporization
of the divertor surface material and accumulated during previous plasma
disruptions.  This dust could be mobilized during an accident or during
in-vessel maintenance.  Assuming an interaction surface area of 15~m$^2$, the
total amounts of tungsten vaporized and melted during a disruption are 1.16 and
60.8~kg, respectively.  All of the vaporized tungsten and 10\% of the melted
tungsten (melt layer splash) are assumed to form dust resulting in the total
production of 7.24~kg of tungsten per full-power
disruption~\cite{Hassanein98}. As shown in Table~\ref{t-9}, considering the
vacuum and containment boundaries as two independent barriers with an overall
radioactivity containment factor of 99.99\%, the divertor disruption would only
produce an effective whole body early dose of 7.68~$\mu$Sv at the site
boundary.

One of the issues that needs to be addressed in a future analysis is the impact
of the divertor shape on the amount of material vaporized during
disruption. Because of the narrow V-shape (closed) of the ARIES-ST divertor,
the radiation energy (photons) from the vapor cloud (the vapor shielding area)
will be deposited on near surface area causing significantly higher additional
erosion~\cite{Hassanein98}.  Taking the vapor cloud effect into account could
increase the amount of dust generated by as much as an order of magnitude more
than the values reported in Table~\ref{t-9}.

\section{Summary and Conclusions}
\label{s-6}

Detailed activation and safety analyses were performed for the ARIES-ST
spherical tokamak power plant.  ARIES-ST includes a water-cooled copper
centerpost and uses a SiC/Li$_{17}$Pb$_{83}$ blanket.  The plant is assumed to
operate for 40~full power years.  Neutron irradiation resulted in the increase
of the centerpost copper resistivity due to the production of neutron-induced
transmutation.  Neutron transmutation of copper resulted in the production of
several nickel, cobalt and zinc isotopes.  The production of these isotopes
resulted in an increase of the time-space average resistivity of the centerpost
by as much as 6\% after 2.86~FPY.  Waste disposal limits were calculated for
the different plant components using the NRC 10CFR61 and Fetter waste disposal
limits.  All of the plant components met the limits for disposal as Class C
low-level waste.  The decay heat generated during the first day following a
LOCA would increase the centerpost temperature by $<1,000$\degC, the first
wall/blanket temperature by $<800$\degC and the divertor temperature by
$<800$\degC.  The calculated temperature profiles and available
oxidation-driven volatility experimental data were used to calculate doses at
the site boundary under conservative release conditions.  The vacuum vessel and
the containment were assumed to stay intact during accidents and hence act as
release barriers.  A leak rate of 1\% per day and a containment factor of 99\%
were considered for each of the two barriers.  The current design 
produces an effective whole body early
dose of 1.88~mSv at the site boundary.  Finally, a divertor disruption would
only produce an effective whole body early dose of 7.68~$\mu$Sv at the site
boundary.

{\bf Acknowledgment}

Support for this work was provided by the U. S. Department of Energy.

\clearpage

\begin{thebibliography}{99}

%1.
\bibitem{Naj00} F.~Najmabadi and the ARIES Team, ``Overview of ARIES-ST
Spherical Torus Power Plant Study,'' {\it Fusion Eng. \&
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%2.
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Power Core Engineering for the ARIES-ST Power Plant,'' ' {\it Fusion Eng. \&
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%3.
\bibitem{ELG99} Laila A. El-Guebaly and the ARIES Team, ``ARIES-ST 
NuclearAnalysis and Shield Design,'' this
issue.

%4.
\bibitem{Odell82} R. O'Dell \ea, ``User's Manual for ONEDANT: A Code Package
for One-Dimensional, Diffusion-Accelerated, Neutral Particle Transport,'' Los
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%5.
\bibitem{MACF95} R.~MacFarlane, ``FENDL/MG-1.0, Library of Multigroup 
Cross Sections in GENDF and MATXS
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IAEA-NDS-129, Rev. 2, International Atomic Energy
Agency (April 1995).

%6.
\bibitem{DKP-PUL} J.~Sisolak, Q.~Wang, H.~Khater, and D.~Henderson,
``DKR-PULSAR2.0: A Radioactivity Calculation Code that Includes
Pulsed/Intermittent Operation,'' to be published.

%7.
\bibitem{Pash97} A. Pashchenko \ea, ``FENDL/A-2.0: Neutron Activation
Cross-Section Data Library for Fusion Applications,'' International Atomic
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1997).

%8.
\bibitem{AIRDOS89} ``User's Guide for AIRDOS-PC, Version 3.0,''
U. S. Environmental Protection Agency report EPA 520/6-89-035, EPA Office of
Radiation Programs, Las Vegas, NV (December 1989).

%9.
\bibitem{10CFR61} U.~S.\ Nuclear Regulatory Commission, 10CFR part 61,
``Licensing Requirements for Land Disposal of Radioactive Waste,'' Federal
Register, FR 47, 57446 (1982).

%10.
\bibitem{Fetter90} S.~Fetter, E.~Cheng and F.~Mann, ``Long Term Radioactive
Waste from Fusion Reactors,'' {\it Fusion Eng' \& Design,} {\bf 13} (1990)
239-246.

%11.
\bibitem{ANSYS} {\it ANSYS 5.4 User's Manual,} Updo, DN-R300:50-2 (1992).

%12.
\bibitem{Mogahed98} E.~A.~Mogahed and the ARIES Team, ``Loss of Coolant
Accident (LOCA) Analysis of the ARIES-ST Design,'' \ANS{13} .

%13.
\bibitem{Mogahed00} E.~A.~Mogahed, ``Heat Pipes Analysis Applied to ARIES-ST
Centerpost,'' to be published.

%14.
\bibitem{NRC83} U.~S.\ Nuclear Regulatory Commission ,''Regulartory Guide
1.145/Rev.1'' (Feb. 1983).

%15.
\bibitem{Chanin90} D.~I.~Chanin \ea, ``MELCOR Accident Consequences Code System
(MACCS),'' Vol.\ 1-3, NUREG/CR-4691, SAND86-1562 (Feb. 1990).

%16.
\bibitem{Mccarthy94} K.~A.~McCarthy, G.~R.~Smolik, and S.~L.~Harms, ``A Summary
and Assessment of Oxidation Driven Volatility Experiments at INEL and Their
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%17.
\bibitem{Mccarthy95} K.~A.~McCarthy and M. J. Gaeta, ``Activation Source Term
Input for ESECS,'' Idaho National Engineering Laboratory Design File
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%18.
\bibitem{Mccarthy96} K.~A.~McCarthy, "Activation Product Source Term for
NSSR-1," Idaho National Engineering Laboratory Design File ITER/US/95/TE/SA-10
(July 26, 1996).

%19.
\bibitem{Wittenberg} L.~J.~Wittenberg, University of Wisconsin-Madison, private
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%20.
\bibitem{Hoffman85} N.~J.~Hoffman \ea, ``Polonium Aspects Associated with the
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%21.
\bibitem{Ginacarli95} L.~Giancarli and M.~A. F\"utterer, "Water-Cooled Pb-17Li
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%22.
\bibitem{Taylor96} N.~P.~Taylor \ea, ``Safety and Environmental Assessment of a
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%23.
\bibitem{Petti94} D.~A.~Petti, ``Estimates of Activation Product Transport
through ITER Confinements,'' Idaho National Engineering Laboratory Design File
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%24.
\bibitem{Petti96} D.~A.~Petti and D.~L.~Hagrman, ``Estimates of Activation
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(December 1997).

%26.
\bibitem{Hassanein98} A.~Hassanein, Argonne National Laboratory, private
communications, November 1998.

\end{thebibliography}

\clearpage


\begin{table}
\caption{}
\centering{\bf Materials Used in the Activation Analysis of ARIES-ST}
\vspace{1.0cm}
\renewcommand{\arraystretch} {1.3}
\begin{tabular}{l @{\hspace{1.cm}} l}
\hline
Component                        & Composition \\
\hline \\
Centerpost                      & 85\% GlidCop Cu and 15\% water \\
Inboard Shield                   & 80\% ferritic steel and 20\% He \\
Inboard First Wall               & 26\% ferritic steel and 74\% He \\
Outboard First Wall              & 40\% ferritic steel and 60\% He \\
Outboard Blanket                 & 6\% ferritic steel, 6\% He, 12\% SiC, \\
Outboard Blanket                 & \hspace{1.cm} and 76\% Li$_{17}$Pb$_{83}$ \\
Outboard Shield                  & 30\% ferritic steel and 70\% He \\
Divertor Tungsten Brush          & 100\% W \\
Divertor Coolant Channel         & 50\% W and 50\% He \\
Divertor Back Plate              & 100\% ferritic steel \\
Divertor Manifold                & 67\% ferritic steel and 33\% He \\
Divertor High Temperature Shield & 15\% ferritic steel, 80\% WC,
                                    and 5\% water \\
Divertor Low Temperature Shield  & 15\% ferritic steel, 70\% WC,
                                    and 15\% water \\
\\ \hline
\end{tabular}
\label{t-1}
\end{table}

\begin{table}
\caption{}
\centering{\bf Waste Disposal Ratings Using 10CFR61 Limits}
\vspace{1.0cm}
\renewcommand{\arraystretch} {1.3}
\begin{tabular}{l @{\hspace{1.cm}} c c c}
\hline
Zone                & Life (FPY) & WDR      &  Dominant Nuclides \\
\hline \\
Centerpost          & 2.86       & 0.88\Z   & $^{63}$Ni \\
Inboard shield      & 2.86       & 0.17\Z   & $^{94}$Nb \\
Inboard first wall  & 2.86       & 0.2\Z\Z  & $^{94}$Nb \\
Outboard first wall & 2.86       & 0.26\Z   & $^{94}$Nb \\
Outboard blanket    & 2.86       & 0.021    & $^{94}$Nb \\
Outboard shield     & 2.86       & $3.7x10^{-3}$ & $^{94}$Nb \\
Divertor plate      & 2.86       & 0.18\Z   & $^{94}$Nb \\
Divertor shield     & 40         & 0.012    & $^{94}$Nb, $^{14}$C \\
\\ \hline
\end{tabular}
\label{t-2}
\end{table}

\begin{table}
\caption{}
\centering{\bf Waste Disposal Ratings Using Fetter Limits}
\vspace{1.0cm}
\renewcommand{\arraystretch} {1.3}
\begin{tabular}{l @{\hspace{1.cm}} c c c}
\hline
Zone                & Life (FPY) & WDR      &  Dominant Nuclides \\
\hline \\
centerpost          & 2.86       & 0.4\Z\Z  & $^{108m}$Ag \\
Inboard shield      & 2.86       & 0.69\Z   & $^{192m}$Ir, $^{94}$Nb \\
Inboard first wall  & 2.86       & 0.55\Z   & $^{192m}$Ir, $^{94}$Nb \\
Outboard first wall & 2.86       & 0.65\Z   & $^{192m}$Ir, $^{94}$Nb \\
Outboard blanket    & 2.86       & 0.049    & $^{192m}$Ir, $^{94}$Nb \\
Outboard shield     & 2.86       & 0.015    & $^{192m}$Ir \\
Divertor plate      & 2.86       & 0.49\Z   & $^{94}$Nb, $^{108m}$Ag \\
Divertor shield     & 40         & 0.017    & $^{94}$Nb, $^{192m}$Ir \\
\\ \hline
\end{tabular}
\label{t-3}
\end{table}

\begin{table}
\caption{}
\centering{\bf Initial Thermal Conditions (Temperatures Are in \degC)}
\vspace{1.0cm}
\renewcommand{\arraystretch} {1.3}
\begin{tabular}{l @{\hspace{2cm}} c c c }
\hline
                      	& $T_{min}$ & $T_{max}$ & $T_{avg}$ \\
\hline \\
Outboard first wall \\
\quad  He		& 300       & 425       & 362.5  \\ 
\quad Ferritic Steel    & 300       & 600       & 450\D  \\ 
Outboard blanket \\
\quad He in grid plates & 425       & 500       & 462.5  \\  
\quad Ferritic Steel    & 450       & 550       & 500\D  \\
\quad Li$_{17}$Pb$_{83}$ & 500      & 700       & 600\D  \\
Inboard first wall \\
\quad  He		& 300       & 500       & 400\D  \\  
\quad Ferritic Steel    & 300       & 600       & 450\D  \\
Inboard shield \\
\quad He		& 300       & 500       & 400\D  \\ 
\quad Ferritic Steel    & 300       & 550       & 475\D  \\
Centerpost \\
\quad Water		& \Z30      & \Z90      & \Z60\D \\    
\quad Copper		& \Z70\D    & 130       & 100\D   \\
Divertor \\
Tungsten                & 820 \\
\\ \hline
\end{tabular}
\label{t-4}
\end{table}


\begin{table}
\caption{}
\centering{\bf Routine Atmospheric Release Parameters}
\vspace{1.0cm}
\renewcommand{\arraystretch} {1.3}
\begin{tabular}{l @{\hspace{1.cm}} l}
\hline \\
{\bf Site Information:} \\
\hspace{1cm} Locations:       & Albuquerque , Boston \\
                               & Chicago , Los Angeles \\
\hspace{1cm} Temperature:     & 15\degC \\
\hspace{1cm} Rainfall:        & 75~cm/y \\
{\bf Emission Information:} \\
\hspace{1cm} Year-Round Averaging \\
\hspace{1cm} Stack Height:    & 75~m \\
\hspace{1cm} Stack Diameter:  & 30~cm \\
\hspace{1cm} Momentum:        & 1~m/s \\
{\bf Tritium Pathways:} \\
\hspace{1cm} Steam Generator: & 10~Ci/day \\
\hspace{1cm} Total (adjusted for 80\% availability): & 2,920~Ci/y \\
\\ \hline
\end{tabular}
\label{t-5}
\end{table}

\begin{table}
\caption{}
\centering{\bf Dose to the Maximally Exposed Individual (MEI)}
\vspace{1.0cm}
\renewcommand{\arraystretch} {1.3}
\begin{tabular}{l @{\hspace{1.cm}} c c }
\hline
Site        & Dose (mrem/yr) & Distance (m) \\
\hline \\
Albuquerque & 0.21           & 1,000 \\
Boston      & 0.09           & 3,000 \\
Chicago     & 0.13           & 1,000 \\
Los Angeles & 0.28           & 3,000 \\
\\ \hline
\end{tabular}
\label{t-6}
\end{table}

\begin{table}
\caption{}
\centering{\bf Tritium Parameters}
\vspace{1.0cm}
\renewcommand{\arraystretch} {1.3}
\begin{tabular}{l @{\hspace{1.cm}} l}
\hline \\
In Li$_{17}$Pb$_{83}$ \\
\hspace{1cm} Li$_{17}$Pb$_{83}$ flow rate      & 47,450 kg/s \\
\hspace{1cm} Li$_{17}$Pb$_{83}$ flow rate      & $2.5\times 10^5$ mol/s \\
\hspace{1cm} Sievert's constant at 700\degC
                                   & $2\times 10^{-8}$ atom frac/Pa$^{0.5}$ \\
\hspace{1cm} Tritium source term               & $1.8\times 10^{-3}$ mol/s \\
\hspace{1cm} Tritium concentration increase per coolant pass  & 7.2 appb \\
\hspace{1cm} Tritium concentration in ARIES-ST                & 0.72 appm  \\
\hspace{1cm} Tritium partial pressure over Li$_{17}$Pb$_{83}$ & 7,400 Pa \\
\hspace{1cm} Estimated Li$_{17}$Pb$_{83}$ inventory           & 150 m$^3$ \\
\hspace{1cm} Total tritium inventory in Li$_{17}$Pb$_{83}$    & 16~g \\
\\
In He \\
\hspace{1cm} He flow rate                  & 1,444 kg/s \\
\hspace{1cm} He density at 500\SdegC       & 7.6 kg/m$^3$ \\
\hspace{1cm} He volumetric flow rate       & 190 m$^3$/s \\
\hspace{1cm} Tritium source term           & $1.5\times 10^{-5}$ mol/s \\
\hspace{1cm} Tritium pressure increase per coolant pass & 0.005 Pa \\
\hspace{1cm} Tritium pressure in He        & 5 Pa \\
\hspace{1cm} Estimated He volume           & 200 m$^3$ \\
\hspace{1cm} Tritium inventory in He       & 1 g \\
\\ \hline
\end{tabular}
\label{t-7}
\end{table}

\begin{table}
\caption{}
\centering{\bf Early Doses Released During a Severe LOCA}
\vspace{1.0cm}
\renewcommand{\arraystretch} {1.3}
\begin{tabular}{l @{\hspace{1cm}} c c c c }
\hline
Zone          & Inventory & Mobilized & Released & Dominant Nuclides \\
              & (Sv) & Inventory (Sv) & (mSv)    & Dominant Nuclides \\
\hline \\
Centerpost          &  2,502\D & 1.1 & 0.11\Z\Z  & $^{60}$Co \\
Inboard Shield      &  4,175\D & 1.7 & 0.17\Z\Z  & $^{60}$Co, 
$^{54}$Mn, $^{56}$Mn \\
Inboard first wall  &  \D306\D & 0.2 & 0.02\Z\Z  & $^{54}$Mn, 
$^{56}$Mn, $^{60}$Co \\
Outboard first wall &  5,000\D & 2 & 0.2\Z\Z\Z & $^{54}$Mn, 
$^{56}$Mn, $^{60}$Co \\
Blanket           &  2,257\D & 13.8 & 1.38\Z\Z  & $^{210}$Po, 
$^{203}$Hg, $^{54}$Mn \\
Manifold            &  \D167\D & 0.04 &0.004\Z   & $^{60}$Co \\
Outboard Shield     & \D\Z53\D & 0.008 & 0.0008    & $^{60}$Co \\
Divertor plate      & \D\Z77.8 & 0.004 & 0.0004    & $^{187}$W, 
$^{181}$W, $^{185}$W \\
\\ \hline
\end{tabular}
\label{t-8}
\end{table}

\begin{table}
\caption{}
\centering{\bf Early Doses Released During a Disruption}
\vspace{1.0cm}
\renewcommand{\arraystretch} {1.3}
\begin{tabular}{l @{\hspace{1cm}} c c c c }
\hline
Pathway          & Inventory & Mobilized & Released & Dominant Nuclides \\
                 & (Sv) & Inventory (Sv) & (mSv)    & \\
\hline \\
Direct Vaporization & 79.4 & 10.6 & 1.06 & $^{187}$W, $^{181}$W, $^{185}$W \\
Tokamak Dust        & 54.8 & 66.2 & 6.62 & $^{187}$W, $^{181}$W, $^{185}$W \\
\\ \hline
\end{tabular}
\label{t-9}
\end{table}

\clearpage

\begin{figure}[p]
\vskip1.in
\centerline{
\epsfig{file=ast-6-01.ps,height=6.5in,bbllx=96pt,bblly=111pt,%
       bburx=555pt,bbury=622pt,clip=}}
\caption
{Elevation view of the ARIES-ST power core.}
\label{ast-6-01}
\end{figure}

\begin{figure}
\centerline{\epsfig{file=ast-6-02.ps,height=3.5in}}
\caption
{Configuration of the slotted duct divertor plate.}
\label{ast-6-02}
\end{figure}
\begin{figure}
\centerline{\epsfig{file=ast-6-03.ps,height=3.5in}}
\caption
{Radial distribution of the percentage increase in the centerpost
resistivity.}
\label{ast-6-03}
\end{figure}

\begin{figure}[p]
\vskip3.in
\centerline{\epsfig{file=ast-6-04.ps,width=6.in}}
\caption
{Biological dose rates following shutdown.}
\label{ast-6-04}
\end{figure}

\begin{figure}[p]
\centerline{\epsfig{file=ast-6-05.ps,height=7.2in}}
\caption
{Finite element model for the inboard and outboard of ARIES-ST.}
\label{ast-6-05}
\end{figure}

\begin{figure}
\centerline{\epsfig{file=ast-6-06.ps,height=3.5in}}
\caption
{Average specific decay heat in a 100\% dense material at the midplane.}
\label{ast-6-06}
\end{figure}
\begin{figure}
\centerline{\epsfig{file=ast-6-07.eps,height=3.5in}}
\caption
{Average specific decay heat in a 100\% dense material behind the tungsten
shielding plate.}
\label{ast-6-07}
\end{figure}

\begin{figure}
\centerline{\epsfig{file=ast-6-08.ps,height=3.5in}}
\caption
{Simplified finite element ($r,\theta$) model for the inboard and outboard
sides.}
\label{ast-6-08}
\end{figure}
\begin{figure}
\centerline{\epsfig{file=ast-6-09.ps,height=3.5in}}
\caption
{Maximum temperature of the outboard first wall as a function of percentage of
thermal conduction through the gaps.}
\label{ast-6-09}
\end{figure}

\begin{figure}
\centerline{\epsfig{file=ast-6-10.ps,height=3.5in}}
\caption
{Temperature variation history of the inboard components during the first week
following a LOCA.}
\label{ast-6-10}
\end{figure}
\begin{figure}
\centerline{\epsfig{file=ast-6-11.eps,height=3.5in}}
\caption
{Temperature variation history of the outboard components during the first week
following a LOCA.}
\label{ast-6-11}
\end{figure}

\begin{figure}
\centerline{\epsfig{file=ast-6-12.ps,height=3.5in}}
\caption
{Temperature variation of the divertor tungsten brush.}
\label{ast-6-12}
\end{figure}

%
%       FIGURE CAPTIONS
%
\clearpage

\begin{center}
{\bf List of Figure captions}
\end{center}

\noindent
{\bf Figure~\ref{ast-6-01}}: Elevation view of the ARIES-ST power core.

\noindent
{\bf Figure~\ref{ast-6-02}}: Configuration of the slotted duct divertor plate.

\noindent
{\bf Figure~\ref{ast-6-03}}: Radial distribution of the percentage increase in
the centerpost resistivity.

\noindent
{\bf Figure~\ref{ast-6-04}}: Biological dose rates following shutdown.

\noindent
{\bf Figure~\ref{ast-6-05}}: Finite element model for the inboard and outboard
of ARIES-ST.

\noindent
{\bf Figure~\ref{ast-6-06}}: Average specific decay heat in a 100\% 
dense material at the midplane.

\noindent
{\bf Figure~\ref{ast-6-07}}: Average specific decay heat in a 100\% dense
material behind the tungsten shielding plate.

\noindent
{\bf Figure~\ref{ast-6-08}}: Simplified finite element ($r,\theta$) model for
the inboard and outboard sides.

\noindent
{\bf Figure~\ref{ast-6-09}}: Maximum temperature of the outboard first wall as
a function of percentage of thermal conduction through the gaps.

\noindent
{\bf Figure~\ref{ast-6-10}}: Temperature variation history of the inboard
components during the first week following a LOCA.

\noindent
{\bf Figure~\ref{ast-6-11}}: Temperature variation history of the outboard
components during the first week following a LOCA.

\noindent
{\bf Figure~\ref{ast-6-12}}: Temperature variation of the divertor tungsten
brush.

\end{document}