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\begin{document}

\title{\bf 6. SAFETY AND ENVIRONMENT}
\author{
\autlist{Thomas J. Dolan}{Jeffrey Crowell}{Hesham Y. Khater}
}
\date{ }
\maketitle

\begin{titlepage}
{
\tableofcontents
}
\end{titlepage}

\setcounter{chapter}{5}
\setcounter{page}{1}

\chapter{SAFETY AND ENVIRONMENT}

\section{INTRODUCTION}
\label{s-6-1}

While the scope of the SPPS did not permit a thorough investigation of the
safety and environmental characteristics of the SPPS/MHH, several issues were
examined.  The main safety and environmental aspects considered for the
SPPS/MHH are: (1)~Routine release of tritium, (2)~Safety precautions with
lithium, (3)~Accidental release of radioisotopes driven by a lithium fire,
(4)~Accidental release of vanadium pentoxide (toxicity), and (5)~Waste
disposal.

The tritium inventory and routine release rate are not estimated here.  The
release rate can probably be kept within the desired limits ($\sim$20~GBq/s) by
using standard techniques for tritium removal from liquid metals, such as cold
trapping and molten salt extraction.

Some advantages of using a lithium breeder and coolant are: (1)~Li is a
low-activation breeder material, (2)~It provides a satisfactory breeding ratio,
{\it i.e.,} Be is not required, and Be toxicity can be avoided; (3)~Li blankets
have a low coolant pressure specially if insulating material is used.  As such,
ruptures of tubes and welds are less likely, (4)~Li blankets operates at a high
temperature operation.  The resultant high energy conversion efficiency reduces
the amount of heat rejected to the environment; (5)~Natural convection removal
of decay heat is possible because of good coolant properties (reasonable
viscosity, specific heat comparable to that of water, and thermal conductivity
about 90 times higher than water)

The main hazard associated with lithium is chemical reactivity.  The energy
release associated with a lithium fire could heat the vanadium alloy structure,
increasing the mobilization rates of radioisotopes.  The average energy release
from reaction of lithium with air (N and O) was estimated to be about
12.3~MJ/kg(Li)~\cite{r1}.  Vanadium can also react with oxygen, releasing
12.8~MJ/kg(V), but the oxidation rate is very slow.

\clearpage

\section{SAFETY PRECAUTIONS WITH LITHIUM}
\label{s-6-2}

In order to promote safe use of lithium in the stellarator power plant,
several safety precautions are recommended:

\begin{itemize}

\item
Use a low-afterheat structural material such as V-5Cr-5Ti.

\item
Minimize the fire hazard by minimizing the mass of lithium in the system.  This
can be accomplished by placing the intermediate heat exchanger close to the
torus and using a low volume fraction of lithium in the shield.

\item
Use three barriers between the lithium and the air.  coolant ducts vacuum
chamber around coolant ducts, and a stainless steel shell on ducts that are
outside of the vacuum chamber reactor hall with a steel liner to prevent
lithium contact with concrete.

\item
Use an inert cover gas (such as argon) around the torus.  Keep argon out of
 high neutron flux areas, to avoid its activation.

\item
Exclude water from the reactor hall.

\item
If liquid metal is used for the intermediate coolant loop, use double-walled
steam generator tubes.  Double-wall tube steam generator technology has been
successfully demonstrated in the EBR-II reactor~\cite{r1,r2}.

\item
Place drain tanks below the torus, so that the lithium can be quickly drained
during an accident to prevent exposure to air.  The tanks would contain argon
sealed by fire-retardant valves~\cite{r1}.

\item
Put most other lithium piping connections above the torus, so that leaks in
those tubes will not accidentally drain the torus.

\item
The floors should slope so that spilled lithium will flow along the stainless
steel liner to drains.

\item
Use an inert powder (such as hollow graphite microspheres) to cover lithium
spills.  The powder would float on top of the lithium, minimizing contact with
the air in the event that the argon cover gas were lost.

\item
Avoid missile impacts on the lithium coolant system by placing turbine
generators away, oriented to avoid missiles traveling towards the torus.
Reduce aircraft impact consequences by locating the torus below grade inside a
concrete dome.

\end{itemize}

\section{ACCIDENTAL RELEASE OF RADIOISOTOPES}
\label{s-6-3}

The energy sources present in the power plant are listed in
\hbox{Table~\ref{t-6-3-1}}.  The fractional vanadium oxidation is based on the
hypothetical accident sequenced described below.  The chemical energy of the
lithium is dominant.

\begin{table}[b]
\caption{}
\centering{\bf Some Energy Sources in the SPPS/MHH Power Plant} \\
\vspace{1.0cm}
\renewcommand{\arraystretch}{1.3}
\begin{tabular}{l c}
\hline
Source                                        & Energy (GJ) \\
\hline \\
Plasma magnetic energy                        & negligible$^{(a)}$ \\
Plasma thermal energy                         & \D\Z\Z0.54  \\
Chemical energy of 0.12\% V oxidation$^{(b)}$ & \D\Z\Z4\D\Z \\
Magnetic energy stored in the coils           &  \D\Z78\D\Z \\
Decay heat energy (7 days)                    &   1,940\D\Z \\
Chemical energy of lithium fire$^{(b)}$       &   3,500\D\Z \\
\\ \hline
\multicolumn{2}{l}{$^{(a)}$ Because of negligible plasma current.}\\
\multicolumn{2}{l}{$^{(b)}$ For 42~$m^3$ of V.}\\
\multicolumn{2}{l}{$^{(c)}$ For 600~$m^3$ of Li.}
\end{tabular}
\label{t-6-3-1}
\end{table}

The ``worst-case'' accident scenario considered here is the rupture of a
lithium duct, failure of the drain system, breach of the reactor hall allowing
air ingress, and complete burning of all the lithium.  Such a pessimistic
scenario is required for an ESECOM-type analysis.  It is used here for two
reasons: (1)~This type of analysis is needed to facilitate comparison with
previous studies, such as ESECOM, ARIES, and PULSAR; and (2)~A probabilistic
regulatory-type analysis would require more resources and failure-rate data
than are available.

As for the ESECOM V/Li power plant, it is assumed that the first wall/blanket
reaches 1573~K for 10~hours, followed by 1273~K for~40 hours.  (This case was
also used for the \AII\ tokamak analysis.)  All mobilized material is
assumed to reach the site boundary.  No credit is taken for containment,
plate-out, or deposition.  Such a pessimistic scenario would not be used for a
regulatory-type analysis.

In estimating the potential off-site radiation dose, the following pessimistic
conditions are assumed: (1)~1-km site boundary radius; (2)~1~m/s wind velocity;
(3)~Pasquill F stability conditions; and (4)~near-ground-level release
(0--10~m).

For these conditions a table of ``effective dose'' conversion factors derived
with the MELCOR Accident Consequence Code System (MACCS)~\cite{r1} is used to
estimate the off-site dose to a Maximum Exposed Individual that could result
from release of each isotope. The activities of about 400 radioactive isotopes
present in the first wall, blanket, and shield are considered.  Mobilization
fractions are estimated for some of the most important elements, based upon
INEL experimental data, assuming as was done for the \AII\ study~\cite{r1}.
Default mobilization fractions for the remaining elements are taken from
Ref.~\cite{r1}. The dominant isotopes and resulting potential off-site doses to
a Maximum Exposed Individual are listed in \hbox{Table~\ref{t-6-3-2}} to
\ref{t-6-3-5}.

\begin{table}[b]
\caption{}
\centering{\bf Estimated Doses at the Site Boundary 10 Minutes after\\
a Worst-Case Accident (First-Wall Contribution$^{(a)}$)}
\vspace{1.0cm}
\renewcommand{\arraystretch}{1.3}
\begin{tabular}{l c c c c}
\hline
            & Activity       & Mobilization    & Conversion      & Effective \\
            & (PBq)          & Fraction	       & Factor (Sv/TBq) & Dose (Sv) \\
\hline \\
$^{24}$Na & $1.81\X 10^{+0}$ & $1.00\X 10^{+0}$ & $4.62\X 10^{-4}$ & 0.8379 \\
$^{48}$Sc & $4.30\X 10^{+2}$ & $3.52\X 10^{-4}$ & $1.15\X 10^{-3}$ & 0.1737 \\
T         & $1.39\X 10^{+1}$ & $1.00\X 10^{+0}$ & $2.38\X 10^{-6}$ & 0.0329 \\
$^{46}$Sc & $4.18\X 10^{+1}$ & $3.52\X 10^{-4}$ & $2.22\X 10^{-3}$ & 0.0327 \\
$^{45}$Ca & $2.97\X 10^{+1}$ & $5.25\X 10^{-3}$ & $6.96\X 10^{-5}$ & 0.0108 \\
$^{32}$P  & $6.45\X 10^{-1}$ & $3.00\X 10^{-1}$ & $5.15\X 10^{-5}$ & 0.0100 \\
\quad {\bf Total:}  &        &                  &                 & 1.14\Z\Z \\
\\ \hline
\multicolumn{5}{l}{$(a)$ For a first-wall volume of 13.6~m$^{3}$.}
\end{tabular}
\label{t-6-3-2}
\end{table}

\begin{table}[p]
\caption{}
\centering{\bf Estimated Doses at the Site Boundary 10 Minutes after\\
a Worst-Case Accident (Blanket Contribution$^{(a)}$)}
\vspace{1.0cm}
\renewcommand{\arraystretch}{1.3}
\begin{tabular}{l c c c c}
\hline
            & Activity       & Mobilization    & Conversion      & Effective \\
            & (PBq)          & Fraction	       & Factor (Sv/TBq) & Dose (Sv) \\
\hline \\
$^{24}$Na
  &$6.75\X 10^{+0}$ & $1.00\X 10^{+0}$ & $4.62\X 10^{-4}$ & $3.12\X 10^{+0}$ \\
$^{48}$Sc
  &$1.50\X 10^{+3}$ & $1.02\X 10^{-4}$ & $1.15\X 10^{-3}$ & $1.75\X 10^{-1}$ \\
T
  &$4.59\X 10^{+1}$ & $1.00\X 10^{+0}$ & $2.38\X 10^{-6}$ & $1.09\X 10^{-1}$ \\
$^{32}$P
  &$2.85\X 10^{+0}$ & $3.00\X 10^{-1}$ & $5.15\X 10^{-5}$ & $4.40\X 10^{-2}$ \\
$^{182}$Ta
  &$2.19\X 10^{+0}$ & $1.00\X 10^{-2}$ & $1.66\X 10^{-3}$ & $3.64\X 10^{-2}$ \\
$^{46}$Sc
  &$1.57\X 10^{+2}$ & $1.02\X 10^{-4}$ & $2.22\X 10^{-3}$ & $3.55\X 10^{-2}$ \\
$^{74}$As
  &$9.67\X 10^{-2}$ & $3.00\X 10^{-1}$ & $7.80\X 10^{-4}$ & $2.26\X 10^{-2}$ \\
$^{76}$As
  &$5.62\X 10^{-1}$ & $3.00\X 10^{-1}$ & $1.30\X 10^{-4}$ & $2.19\X 10^{-2}$ \\
$^{99}$Mo
  &$2.58\X 10^{-1}$ & $3.00\X 10^{-1}$ & $1.74\X 10^{-4}$ & $1.35\X 10^{-2}$ \\
$^{35}$S
  &$4.03\X 10^{+0}$ & $1.00\X 10^{+0}$ & $3.12\X 10^{-6}$ & $1.26\X 10^{-2}$ \\
$^{51}$Cr
  &$1.06\X 10^{+3}$ & $2.43\X 10^{-4}$ & $3.53\X 10^{-5}$ & $9.10\X 10^{-3}$ \\
$^{49}$V
  &$9.70\X 10^{+2}$ & $1.54\X 10^{-3}$ & $3.61\X 10^{-6}$ & $5.40\X 10^{-3}$ \\
$^{47}$Sc
  &$4.73\X 10^{+2}$ & $1.02\X 10^{-4}$ & $8.61\X 10^{-5}$ & $4.15\X 10^{-3}$ \\
$^{75}$Se
  &$2.10\X 10^{-2}$ & $3.00\X 10^{-1}$ & $5.18\X 10^{-4}$ & $3.26\X 10^{-3}$ \\
$^{45}$Ca
  &$9.99\X 10^{+1}$ & $4.22\X 10^{-4}$ & $6.96\X 10^{-5}$ & $2.93\X 10^{-3}$ \\
$^{58}$Co
  &$1.73\X 10^{-1}$ & $1.00\X 10^{-2}$ & $1.08\X 10^{-3}$ & $1.87\X 10^{-3}$ \\
$^{47}$Ca
  &$3.56\X 10^{+0}$ & $4.22\X 10^{-4}$ & $7.08\X 10^{-4}$ & $1.06\X 10^{-3}$ \\
\quad {\bf Total:}
  &                 &                  &                  & 3.63 \\
\\ \hline
\multicolumn{5}{l}{$(a)$ For a blanket volume of 374~m$^{3}$.}
\end{tabular}
\label{t-6-3-3}
\end{table}

\begin{table}[p]
\caption{}
\centering{\bf Estimated Doses at the Site Boundary 10 Minutes after\\
a Worst-Case Accident (Hot Shield Contribution$^{(a)}$)}
\vspace{1.0cm}
\renewcommand{\arraystretch}{1.13}
\begin{tabular}{l c c c c}
\hline
   & Activity       & Mobilization     & Conversion       & Effective \\
   & (PBq)          & Fraction	       & Factor (Sv/TBq)  & Dose (Sv) \\
\hline \\
$^{24}$Na
 & $2.99\X 10^{+0}$ & $1.00\X 10^{+0}$ & $4.62\X 10^{-4}$ & $1.38\X 10^{+0}$ \\
$^{54}$Mn
 & $9.49\X 10^{+3}$ & $1.16\X 10^{-4}$ & $9.19\X 10^{-4}$ & $1.01\X 10^{+0}$ \\
$^{56}$Mn
 & $9.11\X 10^{+4}$ & $1.16\X 10^{-4}$ & $3.33\X 10^{-5}$ & $3.51\X 10^{-1}$ \\
$^{182}$Ta
 & $1.83\X 10^{+1}$ & $1.00\X 10^{-2}$ & $1.66\X 10^{-3}$ & $3.05\X 10^{-1}$ \\
$^{210}$Po
 & $2.15\X 10^{-2}$ & $1.00\X 10^{-1}$ & $9.07\X 10^{-2}$ & $1.95\X 10^{-1}$ \\
$^{188}$Re
 & $2.14\X 10^{+1}$ & $3.00\X 10^{-1}$ & $2.65\X 10^{-5}$ & $1.70\X 10^{-1}$ \\
$^{124}$Sb
 & $2.47\X 10^{+0}$ & $3.00\X 10^{-2}$ & $1.98\X 10^{-3}$ & $1.46\X 10^{-1}$ \\
$^{160}$Tb
 & $7.06\X 10^{+0}$ & $1.00\X 10^{-2}$ & $1.26\X 10^{-3}$ & $8.92\X 10^{-2}$ \\
$^{45}$Ca
 & $1.08\X 10^{+1}$ & $1.00\X 10^{-1}$ & $6.96\X 10^{-5}$ & $7.50\X 10^{-2}$ \\
$^{122}$Sb
 & $4.98\X 10^{+0}$ & $3.00\X 10^{-2}$ & $2.62\X 10^{-4}$ & $3.92\X 10^{-2}$ \\
$^{60}$Co
 & $1.05\X 10^{+2}$ & $2.54\X 10^{-5}$ & $4.55\X 10^{-3}$ & $1.22\X 10^{-2}$ \\
$^{110m}$Ag
 & $1.01\X 10^{+0}$ & $1.00\X 10^{-2}$ & $3.54\X 10^{-3}$ & $3.59\X 10^{-2}$ \\
$^{113m}$Cd
 & $8.05\X 10^{-2}$ & $1.00\X 10^{-1}$ & $4.14\X 10^{-3}$ & $3.33\X 10^{-2}$ \\
$^{76}$As
 & $8.22\X 10^{-1}$ & $3.00\X 10^{-1}$ & $1.30\X 10^{-4}$ & $3.21\X 10^{-2}$ \\
T
 & $1.34\X 10^{+1}$ & $1.00\X 10^{+0}$ & $2.38\X 10^{-6}$ & $3.19\X 10^{-2}$ \\
$^{32}$P
 & $7.27\X 10^{+0}$ & $2.54\X 10^{-2}$ & $5.15\X 10^{-5}$ & $9.53\X 10^{-3}$ \\
$^{47}$Ca
 & $3.58\X 10^{-1}$ & $1.00\X 10^{-1}$ & $7.08\X 10^{-4}$ & $2.53\X 10^{-2}$ \\
$^{48}$Sc
 & $1.66\X 10^{+2}$ & $3.40\X 10^{-5}$ & $1.15\X 10^{-3}$ & $6.46\X 10^{-3}$ \\
$^{186}$Re
 & $9.35\X 10^{-1}$ & $3.00\X 10^{-1}$ & $4.26\X 10^{-5}$ & $1.20\X 10^{-2}$ \\
$^{51}$Cr	      					  
 & $3.88\X 10^{+3}$ & $2.57\X 10^{-5}$ & $3.53\X 10^{-5}$ & $3.52\X 10^{-3}$ \\
$^{192}$Ir	      					  
 & $8.75\X 10^{-1}$ & $1.00\X 10^{-2}$ & $1.14\X 10^{-3}$ & $1.00\X 10^{-2}$ \\
$^{192m}$Ir	      					  
 & $9.96\X 10^{-2}$ & $1.00\X 10^{-2}$ & $5.87\X 10^{-3}$ & $5.85\X 10^{-3}$ \\
\quad {\bf Total:}
  &                 &                  &                  & 2.66 \\
\\ \hline
\multicolumn{5}{l}{$(a)$ For a hot shield volume of 587~m$^{3}$.}
\end{tabular}
\label{t-6-3-4}
\end{table}

\clearpage

\begin{table}[t]
\caption{}
\centering{\bf Estimated Doses at the Site Boundary 10 Minutes after\\
a Worst-Case Accident (Low-Temperature Shield Contribution$^{(a)}$)}
\vspace{1.0cm}
\renewcommand{\arraystretch}{1.15}
\begin{tabular}{l c c c c}
\hline
           & Activity         & Mobilization    & Conversion      & Effective \\
           & (PBq)            & Fraction	& Factor (Sv/TBq) & Dose (Sv) \\
\hline \\
T          & $8.61\X 10^{+0}$ & $1.00\X 10^{+0}$ & $2.38\X 10^{-6}$ & 0.0205 \\
$^{56}$Mn  & $1.11\X 10^{+3}$ & $3.74\X 10^{-4}$ & $3.33\X 10^{-5}$ & 0.0138 \\
$^{54}$Mn  & $1.59\X 10^{+1}$ & $3.74\X 10^{-4}$ & $9.19\X 10^{-4}$ & 0.0055 \\
$^{210}$Po & $3.82\X 10^{-4}$ & $1.00\X 10^{-1}$ & $9.07\X 10^{-2}$ & 0.0035 \\
$^{182}$Ta & $1.77\X 10^{-1}$ & $1.00\X 10^{-2}$ & $1.66\X 10^{-3}$ & 0.0029 \\
$^{160*}$Tb& $1.70\X 10^{-1}$ & $1.00\X 10^{-2}$ & $1.26\X 10^{-3}$ & 0.0022 \\
$^{124}$Sb & $2.93\X 10^{-2}$ & $3.00\X 10^{-2}$ & $1.98\X 10^{-3}$ & 0.0017 \\
\quad {\bf Total:} &          &                  &                 & 0.050\Z \\
\\ \hline
\multicolumn{5}{l}{$(a)$ For a low-temperature shield volume of 538~m$^{3}$.}
\end{tabular}
\label{t-6-3-5}
\end{table}

The dominant isotopes are $^{54}$Mn in the shield and $^{24}$Na in the blanket.
A recent estimate shows that practically all of the Na could be mobilized at
1573~K~\cite{r1}.  From this analysis, the total Effective Dose (based on MACCS
dose conversion factors) would be about 7.5~Sv, exceeding the 2~Sv threshold
dose for prompt off-site fatalities.

The ESECOM and ARIES studies used the older Porter dose conversion factors.
With the Porter dose conversion factors, the off-site Early Dose would be about
1~Sv.  It might be appropriate to consider this lower value when comparing the
SPPS with those power plants.

There are many mitigating factors that would reduce the off-site dose below the
value of the pessimistic case considered here.  The high temperatures and
oxidation leading to radioactivity releases would be prevented or reduced by
many barriers, passive safety features, and phenomena.  These include the
chamber, reactor hall, lithium drains, inert gas, fire suppression system,
building integrity, aerosol plate out in the building, stack, and aerosol
settling outside.  The approximate effectiveness of some of these safety
features is indicated in Table~\ref{t-6-3-6}.

\begin{table}[p]
\caption{}
\centering{\bf Effectiveness of Some Fusion Power Plant Safety Features$^{(a,b)}$}
\vspace{1.0cm}
\renewcommand{\arraystretch}{1.22}
\begin{tabular}{l l}
\hline \\
\bf{Design Features}
\\
No water in building  & Water ingress would require breach of the \\
                      & confinement plus a source of water.       \\
Low pressure coolant  & Coolant pressurization would require a    \\
                      & pressure source, such as a chemical reaction,\\
                      & combined with intact ducts (or very small leaks).\\
\\
\bf{Passive Features} & \bf{Typical failure rates}                 \\
Multiple barriers                                                  \\
Vessels \& ducts      & $10^{-4}$--$10^{-5}$/y for strong vessels       \\
Shell                 & $10^{-1}$--$10^{-2}$/demand, weak barriers       \\
Building              & $10^{-1}$/demand                            \\
Lithium drains to tanks& $10^{-1}$--$10^{-2}$/demand                      \\
Concrete liner        & $10^{-3}$/demand                               \\
Inert cover gas       & $10^{-1}$/demand                               \\
Stack structure       & $10^{-4}$--$10^{-5}$/y for earthquakes            \\
Aerosol plate-out     & $\sim$0.01\%--100\% releasable                  \\
\\
\bf{Active Features}  &                                                  \\
Stack exhaust  & $10^{-4}$--$10^{-5}$/h fan failure                \\
Li fire suppression system& $10^{-2}$--$10^{-3}$/demand                   \\
Air filtration        & $10^{-5}$/hour failure of filter fibers        \\
\\ \hline
\multicolumn{2}{l}{$(a)$Based on information provided by L. C. Cadwallader,
INEL.}\\
\multicolumn{2}{l}{$(b)$These approximate values are for illustrative
 purposes only, and should not}\\
\multicolumn{2}{l}{\phantom{$(b)$} be used for quantitative risk analyses.}
\end{tabular}
\label{t-6-3-6}
\end{table}

From these values, it is apparent that, if common mode failures do not
occur, the combination of all these safety features will result in very low
frequencies for severe accidents.

\section{VANADIUM RELEASE}
\label{s-6-4}

During a high-temperature lithium fire, some of the vanadium alloy would
become oxidized and mobilized as a V$_2$O$_5$ aerosol.  Since vanadium is
mildly toxic, the release of this aerosol to the site boundary could result
in adverse health effects.  Symptoms of acute vanadium exposure include
respiratory tract irritation, bronchitis, and pneumonitis.  Severe vanadium
toxicity may also cause gastrointestinal disturbance, liver, kidney and central
nervous system effects, pulmonary edema, and pneumonia.  Recovery usually
occurs with cessation of exposure, but sensitization may occur in previously
exposed individuals.  The purpose of this section is to assess the magnitude
of the vanadium toxicity hazard.

There are about 250 tonnes of vanadium alloy in the stellarator power
plant.  The inert cover gas system would suppress oxidation of lithium and
vanadium during an accident.  However, if the building and chamber were
breached, then oxidation of the Li and part of the V could occur.  After an
air ingress the lithium would tend to oxidize much faster than the vanadium,
gradually raising the temperature of the first wall, blanket, and shield.
The vanadium mobilization fraction for the first wall and blanket is estimated
to be about 0.0012.  Thus, a total of about 300 kg of V could be mobilized
as V$_2$O$_5$ in a severe accident.  What would the health effects be at the
site boundary?

Consider a pessimistic case with release of all 300 kg during 10 hours as a
very fine aerosol, with no plate-out or settling.  Near-ground-level release is
assumed, together with Pasquill class E or F stability, a steady wind speed
$u=1$~m/s, and a dispersion factor
%
\begin{eqnarray}
 uc/Q  & =  & 3\times 10^{-4}~{\rm m}^{-2}  \: ,
\label{e-7-4-1}
\end{eqnarray}
%
which is between the value from Eisenbud (\hbox{$6\X 10^{-4}$~m$^{-2}$}) and
the value from Porter (\hbox{$3.5\X 10^{-5}$~m$^{-2}$}).  The resulting
average air concentration of V$_2$O$_5$ downwind at the 1 km site boundary
would be
%
\begin{eqnarray}
 c  & = & 2.5~{\rm mg(V)/m}^{3}  .
\label{e-7-4-2}
\end{eqnarray}
%
The actual concentration would probably be lower than this value, due to
deposition and more favorable atmospheric conditions.) The National Institute
of Safety and Health (NIOSH) recommended threshold exposure limit to vanadium
compounds in air (chronic exposure) is 0.05~mg/m$^3$.  The concentration
``Immediately Dangerous to Life and Health'' (IDLH) is 70~mg/m$^3$. Thus, if a
large lithium-vanadium fire occurred with breach of the reactor building, the
chemical toxicity hazard during plume passage would be below the value that
poses an immediate threat to life and health, but possibly above the value
recommended for continuous, long-term exposure.

\section{WASTE DISPOSAL}
\label{s-6-5}

The  shield materials have been selected primarily for their neutron
attenuation capability and cost.  A secondary consideration is the generation
and disposal of long-lived radioactive wastes.

The choice of vanadium alloy with lithium coolant is particularly beneficial
from the waste disposal standpoint.  The lithium coolant will be continually
purified during operation, and it can be recycled when the power plant is
decommissioned.

The vanadium alloy will become highly radioactive during operation, but it,
too, has recycle potential.  If the alloy constituents were initially very
pure, then the dominant radioisotopes left after tens of years cooling are
$^{42}$Ar (32.9~y) and its daughter product $^{42}$K (12.4~h).  These isotopes
can be removed by remelting the alloy in vacuum.  The residual activity depends
upon activation of impurities initially present in the vanadium alloy, such as
cobalt.  In principle, with effective initial purification, the activated
impurity dose rate after 20--60 years of cooling would be $<10$~mSv/h, which is
low enough for hands-on recycling.  Such recycling would save space in waste
repositories, save money by avoiding purchase of new vanadium alloy, help
prevent vanadium price escalation, and help gain public acceptance of fusion
power.

The steel of the shield might be recycled for shielding material in subsequent
fusion power plants, or it could be disposed as waste.  The low-temperature shield
qualifies for Class C according to both NRC and Fetter limits.  On the other
hand, the high-temperature shield qualifies for Class C only according to the
NRC limits.

\section{ACTIVATION ANALYSES}
\label{s-6-6}

\subsection{Introduction}
\label{s-6-6-1}

The safety characteristics of fusion power plants would most probably be
dependent on the choice of the structural material.  The level of radioactivity
induced in any structure depends on the constituent elements of the irradiated
material, level of neutron flux and time of irradiation.  Low activation
materials (LAMs) are an attractive choice because they result in lower levels
of off-site doses in case of accidental release of their radioactive
inventories during an accident.  LAMs can also facilitate better waste
management at the end of the plant life.

Detailed activation analyses were performed to identify the safety,
environmental and rad-waste characteristics of the SPPS/MHH plant.  The
SPPS/MHH structure is made of vanadium alloy and cooled with liquid lithium.

The activity, decay heat and biological hazard potential (BHP) were calculated
for up to 1000 years following shutdown.  Such an evaluation of the structure
activity and biological hazard potential is needed to evaluate the potential
impact of radioactive inventory release at the onset of an accident.  Results
of the decay heat calculation are used to examine the thermal response of the
fusion power core structure following a loss of coolant accident (LOCA) and/or
a loss of flow accident (LOFA).  The waste disposal ratings (WDR) of the
fusion core structure at the end of its lifetime were also evaluated.  The waste
disposal rating is needed to determine if a given structure would satisfy the
regulatory criteria for shallow land burial as low level waste (LLW).  Finally,
to assess the possibility of hands-on maintenance, contact dose rates were
calculated at selected locations inside the containment building.

\subsection{Calculational Procedure}
\label{s-6-6-2}

The one-dimensional discrete ordinates neutron transport code
ONEDANT~\cite{Ode82} was used to calculate the neutron flux for the activation
analysis.  A 46-group neutron and 21-group gamma coupled cross section library
containing P$_3$ Legendre expansions of the scattering cross sections based on
the ENDF/B-V basic data files was used in the transport calculations.  The
SPPS/MHH peak and average neutron wall loadings are 2 and 1.3~MW/m$^2$,
respectively.

The SPPS/MHH activation calculation was conducted using the DKR-ICF computer
code~\cite{Hen87} with activation cross sections taken from the new
USACT93~\cite{Hen93}.  The USACT93 library was developed by Dr. Fred Mann of
Hanford Engineering Design Laboratory.  It is based on neutron transmutation
cross section and isotopic radioactive decay data from the ENDF/B-VI and EAF3
files.  The neutron transmutation data used is in a 46-group structure format.
The gamma source data is taken from the table of isotopes~\cite{Led78} and is
in a 21-group structure format.

\begin{table}[b]
\caption{}
\centering
{\bf Materials Used in the SPPS/MHH Analysis} \\
\vspace{1cm}
\renewcommand{\arraystretch}{1.1}
\begin{tabular}{l @{\hspace{1cm}} l l l}   \hline
\\
First Wall             & 28.6\% V-5Cr-5Ti, & 71.4\% Li \\
Blanket                & 10\% V-5Cr-5Ti,   & 90\% Li \\
Hot shield             & 15\% V-5Cr-5Ti,   & 5\% Li, & 80\% Tenelon \\
Low-temperature shield & 15\% V-5Cr-5Ti,   & 5\% Li, & 80\% Borated-Tenelon \\
\\  \hline
\end{tabular}
\label{t-6-6-1}
\end{table}

The structure activation results were utilized in the rad-waste classification
performed using the WDR~\cite{Kha91} computer code.  The DOSE code~\cite{Hen87}
was used to calculate the contact doses behind the first wall, blanket, and
high and low-temperature shields.  Finally, the structure activation results
were used in the off-site dose calculations presented in this chapter.  The
materials used in the blankets and shield are presented in Table~\ref{t-6-6-1}.
The elemental compositions of the vanadium alloy (V-5Cr-5Ti) and the low
activation austenitic steel (Tenelon) are taken from the Blanket Comparison and
Selection Study (BCSS) report~\cite{Smi84}. The power plant is assumed to
operate continuously for 30 full power years (FPY) which corresponds to 40
years of operation at 75\% availability.  The SPPS/MHH was assumed to have a
28~m diameter.

\subsection{Structure Activity, Decay Heat, and Biological Hazard Potential}
\label{s-6-6-3}

The activity induced in the blanket at the end of its lifetime is higher than
the activity induced in the shield after 30 full power years.  At shutdown, the
blanket and shield activities are 683~MCi and 3908 ~MCi, respectively.
Figure~\ref{f-6-6-1} shows the total activity induced in the different regions
of SPPS/MHH as a function of time following shutdown.  The blanket activity
drops to 139 and 91~MCi within the first day and the first week following
shutdown, respectively. One year after shutdown the blanket activity drops to
22.3~MCi compared to 505~MCi for the shield.  The high-temperature shield
dominates the activities induced in the SPPS/MHH shield.

The blanket short-term activity is dominated by $^{48}$Sc~($T_{1/2}=43.7$~hr),
$^{51}$Cr~($T_{1/2}=27.7$~day), $^{47}$Sc~($T_{1/2}=3.349$~day), and
$^{45}$Ca~($T_{1/2}=162.7$~day).  On the other hand, the shield short-term
activity after shutdown ($\leq$1~day) is dominated by $^{51}$Cr,
$^{54}$Mn~($T_{1/2}=312$~day), $^{56}$Mn~($T_{1/2}=2.578$~hr), and
$^{187}$W~($T_{1/2}=23.9$~hr). In the period between 1~day and 1~year after
shutdown, $^{54}$Mn, $^{60}$Co~($T_{1/2}=5.27$~yr), and
$^{3}$H~($T_{1/2}=12.3$~yr) dominate the activity induced in the shield.
During the same period of time, the blanket activity is dominated by
$^{49}$V~($T_{1/2}=337$~day), $^{45}$Ca, and
$^{46}$Sc~($T_{1/2}=83.81$~day). Finally, the long-term activities induced in
both the shield and blanket come from the steel components and are dominated by
$^{14}$C~($T_{1/2}=5730$~yr), $^{93m}$Nb~($T_{1/2}=16.1$~yr),
$^{94}$Nb~($T_{1/2}=2\X 10^4$~yr), and $^{93}$Mo~($T_{1/2}=3.5\X 10^3$~yr).

The temporal variation of the decay heat generated in the blanket and shield
are shown in Figure~\ref{f-6-6-2}.  The total decay heat generated in the
blanket at shutdown is 7.9~MW and drops to 1.2~MW within an hour and to only
0.2~MW within one week.  At shutdown, 46.6~MW of decay heat are generated in
the shield.  The decay heat drops to only 1.71~MW within a day and 0.77~MW
within the first year.

\begin{figure}[p]
\centerline{\psfig{file=spps-6-6-1.ps,height=3.2in,bbllx=55pt,bblly=180pt,%
            bburx=565pt,bbury=535pt,clip=}}
\caption[FIG-6-6-1]
{Activity induced in the SPPS/MHH.}
\label{f-6-6-1}
\vspace{1.cm}
\centerline{\psfig{file=spps-6-6-2.ps,height=3.7in,bbllx=60pt,bblly=180pt,%
            bburx=520pt,bbury=535pt,clip=}}
\caption[FIG-6-6-2]
{Decay heat induced in the SPPS/MHH.}
\label{f-6-6-2}
\end{figure}

The decay heat generated in the SPPS/MHH is almost dominated by the same
isotopes that dominate the level of activity.  The short-term decay heat
generated in the blanket is due to $^{48}$Sc and $^{52}$V~($T_{1/2}=3.76$~min).
$^{46}$Sc and $^{49}$V are the dominant nuclides up to one year following the
shutdown.  $^{94}$Nb and $^{14}$C dominate the decay heat generated in the
blanket several hundred years following the end of its lifetime.  In the shield
case, $^{56}$Mn and $^{52}$V produce most of the decay heat generated within
the first 8~hours.  Within the first year after shutdown, $^{56}$Mn and
$^{60}$Co are the major sources of decay heat.  The long-term decay heat is
governed by the decay of $^{94}$Nb and $^{108{\rm m}}$Ag~($T_{1/2}=130$~yr).

Figure~\ref{f-6-6-3} shows the total integrated decay heat generated in the
different regions of the SPPS/MHH.  One week after shutdown, the values of the
integrated decay heat are 333~GJ for the blanket and 1,605~GJ for the shield.
These results are useful for predicting the thermal response of the blanket and
shield to a LOCA and/or LOFA.

\begin{figure}[p]
\centerline{\psfig{file=spps-6-6-3.ps,height=3.4in,bbllx=57pt,bblly=190pt,%
            bburx=530pt,bbury=530pt,clip=}}
\caption[FIG-6-6-3]
{Integrated decay heat in the SPPS/MHH.}
\label{f-6-6-3}
\vspace{1.cm}
\centerline{\psfig{file=spps-6-6-4.ps,height=3.55in,bbllx=50pt,bblly=185pt,%
            bburx=525pt,bbury=535pt,clip=}}
\caption[FIG-6-6-4]
{Biological hazard potential in the SPPS/MHH.}
\label{f-6-6-4}
\end{figure}

The biological hazard potentials were calculated using the maximum permissible
concentration limits in air for the different isotopes according to the NRC
regulations specified in 10CFR20~\cite{Nuc75}. Figure~\ref{f-6-6-4} shows the
biological hazard potentials in air as a function of time following shutdown
for the blanket and shield.  The total BHP in the blanket at shutdown is
\hbox{$469\X 10^6$~km$^3$}.  On the other hand, the total BHP generated in the
shield at shutdown is \hbox{$492\X 10^6$~km$^3$} air.

The short-term BHP is dominated by $^{49}$V and $^{48}$Sc in the case of the
blanket, and $^{54}$Mn, $^{56}$Mn, and $^{52}$V in the case of the shield.
While $^{49}$V is responsible for most of the BHP in the blanket for times
($\leq 10$~years), $^{60}$Co and $^{54}$Mn are the major sources of mid-term
BHP generated in the shield.  Finally, in addition to $^{94}$Nb, the long-term
BHP is produced by $^{93}$Mo and $^{108m}$Ag in case of the blanket and shield,
respectively.

\subsection{Rad-waste Classification}
\label{s-6-6-4}

The rad-waste of the blanket and shield of the SPPS/MHH were evaluated
according to both the NRC 10CFR61~\cite{Nuc82} and Fetter~\cite{Fet90} waste
disposal concentration limits (WDL).  The 10CFR61 regulations assume that the
waste disposal site will remain under administrative control for 100~years.
The dose at the site to an inadvertent intruder at the end of the 100 year
period 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 the particular class.  If the calculated WDR $\leq 1$ when Class~A
limits are used, the rad-waste should qualify for Class~A segregated waste.
The major hazard of this class of waste is to the individuals responsible for
handling it.  However, such waste is not considered to be hazardous following
the loss of institutional control over the disposal site.  If the WDR is $> 1$
when Class~A WDL are used but $\leq 1$ 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~year
institutional period is over.  Class~C waste is assumed to be stable for
500~years.  Using Class~C limits, a WDR $> 1$ implies that the rad-waste 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 recent 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-ICF code were also used to calculate
the waste disposal ratings for the blanket and shield of the SPPS/MHH.  The
waste disposal ratings for Class~A and Class~C low level waste are shown in
\hbox{Table~\ref{t-6-6-2}}.  The values in the table are given for both
non-compacted and compacted (in parenthesis) waste.  Non compacted values are
based on averaging the specific activities over the total volume of a
particular region assuming that internal voids will be filled with concrete
before disposal.  On the other hand, compacted values correspond to crushing
the solid waste before disposal.  The 10CFR61 Class~A WDR is given after a
waiting period of about 10~years to allow for the specific activity of
short-lived nuclides ($T_{1/2}\leq5$~years) to drop below 7,000~Ci/m$^3$.  The
7000~Ci/m$^3$ limit is 10 times larger than the limit specified by the NRC for
Class~A disposal of short-lived nuclides where the waste form is not specified.
In comparison with other isotopes for which limits are given for different
waste forms, the factor of 10 is used for isotopes contained in metal waste.
Since the NRC regulations do not specify any limit for short-lived activity for
Class~C LLW, the Class~C WDR values were calculated after 1~year cooling period
for both 10CFR61 and Fetter limits.

\begin{table}[t]
\caption{}
\centering
{\bf SPPS/MHH Waste Disposal Ratings$^{(a)}$} \\
\vspace{1cm}
\renewcommand{\arraystretch}{1.3}
\begin{tabular}{l c c c c} \hline
WDR     &  First wall    & Blanket       &  Hot shield       & LT-Shield \\
\hline \\
Class~A (10CFR61) 
  &  48 (166)        & 7.3 (73)         & 5.23              & 0.68  \\
\vspace{0.3cm}
  & $^{3}$H (80\%)   & $^{3}$H (65\%)   & $^{94}$Nb (65\%)  & $^{3}$H (90\%) \\
Class~C (10CFR61)  
  & 0.83 (2.9)       & 0.25 (2.5)       & 0.4 & 0.0055  \\
\vspace{0.3cm}
  & $^{94}$Nb (80\%) & $^{94}$Nb (85\%) &  $^{94}$Nb (85\%) & $^{94}$Nb (90\%)\\
Class~C (Fetter)
  & 0.89 (3.1)       & 0.24 (2.4)       & 5.27 & 0.1  \\
  & $^{94}$Nb (70\%) & $^{94}$Nb (90\%) &  $^{192{\rm m}}$Ir (85\%) &
                                                    $^{192{\rm m}}$Ir (90\%)  \\
\\
\hline
\multicolumn{5}{l}{$(a)$ The values in the table are given for both
non-compacted and compacted}\\
\multicolumn{5}{l}{\phantom{$(a)$} (in parenthesis) waste. For each case, the
dominant isotope is also listed.}
\end{tabular}
\label{t-6-6-2}
\end{table}

As shown in the table, if the 10CFR61 limits are used, $^{3}$H produces more
than 80\% of the Class~A WDR for the first wall.  $^{94}$Nb is the second major
contributor to the waste disposal rating.  $^{3}$H is also the main contributor
to Class~A in the case of the blanket and low-temperature shield with its high
boron content.  On the other hand, $^{94}$Nb is the main contributor to Class~A
in the case of the shield.  The other major contributor is $^{60}$Co produced
from the cobalt, nickel and copper impurities in the steel.  According to the
same NRC regulations, the Class~C WDR of both the blanket and shield are
dominated by $^{94}$Nb.  If Fetter limits are used, the first wall and blanket
WDR are dominated by $^{94}$Nb, $^{108{\rm m}}$Ag~($T_{1/2}=130$~yr) and
$^{26}$Al.  On the other hand, the shield rating is dominated by $^{192{\rm
m}}$Ir~($T_{1/2}=241$~yr) and $^{94}$Nb.

It was concluded that at shutdown, the first wall and blanket would only
qualify for Class~C LLW according to both NRC and Fetter limits if the waste is
not compacted.  The high-temperature shield could only qualify as Class~C waste
the 10CFR61 limits are used.  Finally, the low-temperature shield would have no
difficulty qualifying for Class~A (after 15 years of cooling period) or Class~C
rating according to both limits used in this analysis.

\subsection{Contact Dose}
\label{s-6-6-5}

Contact dose rates were calculated for maintenance  evaluation.   The doses
were calculated using the DOSE code, which combines the decay gamma
source and the adjoint dose field to determine the contact dose rates at
different times following shutdown.  The decay gamma source at different
times following shutdown were calculated using the DKR-ICF code.  The
adjoint dose field was determined by performing a gamma adjoint calculation
using the ONEDANT code with the flux-to-dose conversion factors
representing the source at the location where the dose was to be calculated.
The contact doses were calculated at four different locations behind the first
wall, blanket, HT-shield, and LT-shield.   A limit of 25~$\mu$Sv/hr for
hands-on maintenance was used in this  analysis, assuming that maintenance
personnel work for 40 hours a  week and 50 weeks a year.  Results in
Figure~\ref{f-6-6-5} shows that by assuming the 25~$\mu$Sv/h limit for
hands-on maintenance, only remote maintenance would be allowed at any of
the locations considered inside the containment building.  The contact dose
inside the SPPS/MHH containment is mostly produced during the first few weeks
by the decay of $^{48}$Sc, $^{52}$V, and $^{46}$Sc.

\begin{figure}[t]
\centerline{\psfig{file=spps-6-6-4.ps,height=3.3in,bbllx=53pt,bblly=185pt,%
            bburx=525pt,bbury=535pt,clip=}}
\caption[FIG-6-6-5]
{Contact dose in the SPPS/MHH.}
\label{f-6-6-5}
\end{figure}

\section{LOSS OF COOLANT ACCIDENT ANALYSIS}
\label{s-6-7}

In the event of a loss of coolant to the stellarator, there are two major
safety concerns.  First, the loss of coolant must be recognized and the plasma
must be quenched before any component is damaged.  Second, the afterheat
generated by the activated components must not result in excessive temperatures
before coolant flow can be restored.

The first of these concerns, which is generic to all magnetically confined
fusion devices, places a limit on the ``response time'' of the system.  This
limit is likely to be set by the time required to reach the maximum temperature
of the first wall during uncooled operation.  A proper evaluation requires a
more detailed first wall design than is warranted for this study.  Therefore,
the response time is treated in a general way here.  Instead, this section will
primarily address the removal of afterheat.

The thermal behavior of the fusion core following a loss of coolant was
simulated with a one-dimensional finite element model.  Several cases are
considered including adiabatic and non-adiabatic boundary conditions.  For the
non-adiabatic condition, heat losses through a containment building were
included in the model.  The adiabatic boundary condition case, which gives a
conservative bound on the actual (non-adiabatic) behavior, experiences damaging
temperatures 26 days after the accident.  Design choices which extend this time
(such as using an all vanadium hot shield) are examined.  Fusion core
temperatures never exceeded their limits in the cases that modeled heat loss
through the containment building.

The stellarator first wall, blanket, and shield modules are completely
surrounded by the evacuated cryostat boxes.  These cryostat boxes, which encase
the magnets and their support structure, are not lined with thermal insulation.
Rather, each coil case is insulated individually.  Most of the heat from the
modules escapes the fusion core by radiating across the cryostat boxes in the
spaces between the coils.  However, in order to reduce the heat load on the
magnets during normal operation, it may be necessary to insulate the cryostat
boxes.  Considerations are made for the effect of such insulation and the use
of access ports as a means of heat removal.

\subsection{Methodology}
\label{s-6-7-1}

\subsubsection{Accident conditions}
\label{s-6-7-1-1}

In the loss of coolant accident considered here, several assumptions are made
about the nature of the accident.  It is assumed that all coolant is lost from
the fusion core.  Any coolant retained in an accident would increase the heat
capacitance of the fusion core, slowing its temperature rise.  Further, the
afterheat of the coolant is negligible compared with the structural materials.
Therefore, this is a conservative assumption.

It is possible for the liquid-metal-coolant system to fail while the liquid
nitrogen and liquid helium systems which maintain the superconductors remain
functional.  In this analysis, it is assumed that the cryogenic systems fail as
well.  In any event, if the magnet cooling systems continue to operate, the
fusion core as a whole will affected only negligibly.  The low capacity of the
cryogenic system and the thermal isolation of the magnets from the modules
prevent substantial cooling of the modules via this system.

After the coolant is lost, a response time is required to identify the loss of
coolant and shut down the plasma.  During this time the surface and nuclear
heating are the same as the peak loads during regular operation.  Peak loads
were used to find the maximum temperature rise during the period of uncooled
operation.  The use of peak loads overestimates the heat added to the fusion
core as a whole.  However, the amount of heat generated during the brief
uncooled operation is negligible compared to the afterheat generated over the
ensuing weeks.  Therefore, the fusion core temperature after several weeks is
virtually independent of the choice of response time.

\subsubsection{Modeling the Stellarator}
\label{s-6-7-1-2}

The geometry of the actual stellarator is quite complex.  In this study, the
fusion core is approximated as cylindrical and the material properties are
``smeared'' (preserving their heat capacitance and thermal conductivity) so as
to depend on the radial direction only.  The temperature distribution in the
fusion core and its containment building are modeled by the finite element
method (using the commercial program ANSYS with one-dimensional axisymmetric
elements).

Several cases with different boundary conditions are analyzed.  First, an
adiabatic condition at the outside of the vacuum vessel is enforced to place a
conservative bound on the results.  In a second case, a containment building is
added to the model.  The vacuum vessel convects heat to the inert atmosphere in
the containment building and this gas convects heat to the containment
building.  (A case in which the gas temperature is held at 20\SdegC\ is
also considered.)  In addition, the vacuum vessel radiates heat to the
containment building.  On the external surface of the containment building,
heat convects and radiates to a 20\SdegC\ environment.  The heat transfer
coefficient both inside and outside the containment building is
2~Wm$^{-2}$K$^{-1}$ independent of temperature.  The true shape and
orientation of the vacuum vessel does not permit radiation and convection from
the entire surface area.  Therefore, it is assumed that one-half of the outside
surface area of the vacuum vessel can convect and radiate.  The same fraction
is applied to the size of the containment building since its area and volume
are overestimated by assuming an annular shape centered on the plasma.

Heat transfer by the support structure is neglected.  This structure
will conduct heat to the ground under the fusion core and across the gaps
reducing the maximum temperatures indicated by this analysis.  However,
these supports will presumably be designed to minimize this loss of energy.
For conservatism these losses are not included.

\begin{figure}[p]
\centerline{\psfig{file=spps-6-7-1.ps,height=7.4in,bbllx=65pt,bblly=186pt,%
            bburx=521pt,bbury=723pt,clip=}}
\caption[Fig. 7-7-1]
{Schematic of the stellarator radial build.  The radii of the various
layers are to scale.}
\label{f-6-7-1}
\end{figure}

The radial build of the fusion core is shown in \hbox{Fig.~\ref{f-6-7-1}}.  The
magnets and their support structure only partially block the thermal radiation
from the inner to outer cryostat box walls.  The fraction of the total area the
magnets block is represented in \hbox{Fig.~\ref{f-6-7-1}} by the fraction of
the total angle that they subtend (41\%).  The one-half factor applied to the
containment building area and volume is represented in the same manner.

As noted above, the model is one-dimensional.  The material properties of the
different layers in the magnets have been adjusted to preserve the heat
capacitance and thermal conductivity of the true geometry.  The azimuthal
facing sides of the magnets do not communicate thermal radiation with the
cryostat box walls.

The material compositions of the regions shown in Fig.~\ref{f-6-7-1} are listed
in \hbox{Table~\ref{t-6-7-1}}.  Note that all coolant is assumed to leave the
fusion core following the accident.  The material properties of the materials
listed in \hbox{Table~\ref{t-6-7-1}} are shown in
\hbox{Table~\ref{t-6-7-2}}~\cite{Smi85,Hol86}.

\begin{table}[p]
\caption{}
%\begin{center}
\centering{\bf Material Composition of the Regions in the SPPS/MHH}
\vspace{1.0cm}
\renewcommand{\arraystretch} {1.2}
\begin{tabular}{l c l } \hline
Region                 & Thickness (cm) & Composition (by volume) \\
\hline \\
First wall                 & \Z\Z0.3\Z  & V                \\
First wall coolant channel & \Z\Z1\D\Z  & 65\% Li, 35\% V  \\
First wall                 & \Z\Z0.1\Z  & V                \\
Blanket                    &  \Z35\D\Z  & 90\% Li, 10\% V  \\
Gap                        & \Z\Z2\D\Z  &                  \\
Hot shield                 &  \Z45\D\Z  & 80\% tenelon, 15\% V, 5\% Li \\
Gap                        & \Z\Z2\D\Z  &                  \\
Cold shield                &  \Z35\D\Z  & 95\% tenelon, 5\% Li \\
Gap                        & \Z\Z8.3\Z  &                  \\
Cryostat box wall          & \Z\Z1\D\Z  & stainless steel  \\
Gap                        & \Z\Z1\D\Z  &                  \\
Magnet insulation          & \Z\Z4\D\Z  & evacuated superinsulation \\
Coil case                  & \Z\Z7.5\Z  & stainless steel  \\
GFF polyimide              & \Z\Z1.25   & GFF polyimide    \\
Winding pack               &  \Z75.5\Z  & 85\% Cu, 15\% He \\
GFF polyimide              & \Z\Z1.25   & GFF polyimide    \\
Coil case                  & \Z\Z7.5\Z  & stainless steel  \\
Magnet insulation          & \Z\Z4\D\Z  & evacuated superinsulation \\
Gap                        & \Z\Z1\D\Z  &                  \\
Cryostat box wall          & \Z\Z1\D\Z  & stainless steel  \\
Gap                        &   100\D\Z  &                  \\
Vacuum vessel              & \Z\Z5\D\Z  & stainless steel  \\
Gap                        &   200\D\Z  &                  \\
Containment building       &   100\D\Z  & concrete         \\
\\ \hline
\end{tabular}
\label{t-6-7-1}
\end{table}

\begin{table}[t]
\caption{}
\centering{\bf Thermal Properties of the Materials in 
SPPS/MHH~\protect\cite{Smi85,Hol86}}
\vspace{1.0cm}
\renewcommand{\arraystretch} {1.0}
\begin{tabular}{l c c c } \hline
              & Density      & Specific heat capacity & Thermal conductivity \\
              & (kg/m$^{3}$) & (J$\cdot$kg$^{-1}\cdot$K$^{-1}$) 
                                                      & (Wm$^{-1}$K$^{-1}$) \\
\hline \\
V-15Cr-5Ti        & 6,100    & 535 @ 400\SdegC        & 26.8 @ 400\SdegC \\
                  &          & 560 @ 500\SdegC        & 28.0 @ 500\SdegC \\
\vspace*{0.2cm}   
                  &          & 575 @ 600\SdegC        & 29.5 @ 600\SdegC \\
Stainless steel   & 8,000    & 560 @ 400\SdegC        & 19.5 @ 400\SdegC \\
                  &          & 575 @ 500\SdegC        & 21.0 @ 500\SdegC \\
\vspace*{0.2cm}   
                  &          & 580 @ 600\SdegC        & 22.5 @ 600\SdegC \\
\vspace*{0.2cm}   
Magnet insulation &          & negligible             & $3\X 10^{-4}$    \\
\vspace*{0.2cm}   
Concrete          & 1,900    & 880\D                  & 1.37             \\
\vspace*{0.2cm}   
Copper            & 8,954    & 383.1                  & 407 @ -100\SdegC \\
                  &          &                        & 386 @ 0\SdegC    \\
                  &          &                        & 379 @ 100\SdegC  \\
                  &          &                        & 374 @ 200\SdegC  \\
\\ \hline
\end{tabular}
\label{t-6-7-2}
\end{table}

The only heat transfer across the gaps is by thermal radiation.  The emissivity
of polished stainless steel has been reported as 0.074~\cite{Hol86} and this
value seems to be consistent with most unoxidized metals.  Therefore, vanadium
alloy was assumed to have the same emissivity.  All radiating surfaces in the
model were given this value save the concrete, which was assumed to have an
emissivity of 0.63~\cite{Hol86}.  High emissivity coatings might be used to
improve the heat transfer in some locations.

\subsubsection{Maximum temperature}
\label{s-6-7-1-3}

There are two temperature limits of interest in this analysis.  The first is
that which causes irreversible damage to the fusion core components.  In the
stellarator, the structural material is vanadium alloy.  Although tenelon is
used in the shields, it is in form of filler material (bricks), so
microstructural changes in the tenelon due to high temperatures will not
present a problem.  The Tokamak Power System Studies~\cite{Bak85} used
1,200\SdegC\ as the maximum temperature for structural vanadium alloy.  This
will be adopted here as the maximum temperature that any vanadium alloy
component can withstand while not under load.

The second temperature limit is the temperature at which the fusion core can
pose a safety problem.  Stainless steel has a melting point of approximately
1,500\SdegC.  Since melting of the tenelon and its subsequent mobility would
present a serious safety risk, 1,500\SdegC\ is assumed for this limit.

\subsubsection{Loads}
\label{s-6-7-1-4}

The total heating in the stellarator following a loss of coolant
is shown in \hbox{Fig.~\ref{f-6-7-2}}.
%
\begin{figure}[b]
\centerline{\psfig{file=spps-6-7-2.ps,height=4.0in,bbllx=76pt,bblly=92pt,%
            bburx=539pt,bbury=425pt,clip=}}
\caption[]
{Total heat generation in the fusion core assuming plasma shutdown at
10 seconds.  Note the logarithmic time scale.}
\label{f-6-7-2}
\end{figure}
%
Note that the power of the radiant and nuclear heating during the uncooled
operation is much larger than the afterheat power.  However, the total
heat deposited during this phase becomes small on the time scales of
interest.   For example, the heat from 10 seconds of uncooled operation,
constitutes 1\% of the total heat after a week.

\subsection{Results}
\label{s-6-7-2}

In the initial seconds, when the coolant has been lost but the plasma continues
to burn, the surface heat flux drives up the first wall temperature rapidly.
The length of this period of uncooled operation must be limited in order to
prevent damage to the first wall.  The maximum permissible response time will
probably be determined by the maximum temperature limit of the first wall,
however, induced thermal stresses should also be considered.  Uncooled
heating of other components is not as severe.  Both the temperature
distribution and the stress distribution depend on the design details of the
first wall.  An approximate upper bound on the response time, determined by
this simple (one-dimensional) model and based on the 1,200\SdegC\ limit of
vanadium, is 30~s.

There is also a danger from a temporary loss of coolant during operation.   If
inlet temperature coolant is restored to a hot component, thermal shock results.
If the loss of coolant is localized, it would almost invariably be better to
shut down the plasma and wait for the hot spot to disperse before restoring
coolant flow, if possible.

The strong dependence of thermal radiation heat transfer rate on temperature
is advantageous for fusion core safety.  In normal operation, the thermal
resistance across the assembly gaps and the cryostat box accounts for most of
the total resistance to heat flow out of the fusion core.  As the temperature
rises, the thermal resistance of the radiation gaps decreases.  For example, in
normal operation 72~kW escapes from the modules (from the 250\SdegC\
cold shield) to the atmosphere outside the vacuum vessel.
If the fusion core rose to 1,000\SdegC\ then 2.2~MW would be released
(in steady-state), even assuming the heat must pass through the containment
building to the environment.

Our first analysis assumes adiabatic conditions at the outer boundary of the
fusion core (vacuum vessel).  Physically, this would approximate (on the
conservative side) a highly insulated containment building with no means of
cooling the gas around the fusion core.  It is useful as a conservative bound
on the fusion core behavior but becomes very inaccurate at long time scales due
to inevitable heat losses through the containment structure.  The temperature
at several positions in the fusion core as a function of time is shown in
\hbox{Fig.~\ref{f-6-7-3}}. Note that the time axis is logarithmic.  The results
show that serious economic losses are incurred if coolant is not restored in 26
days.  The fusion core becomes an environmental threat after 40 days without
cooling.  If adiabatic conditions were maintained long enough (months), the
entire fusion core would melt.  The temperature distribution at several times
following the accident is shown in \hbox{Fig.~\ref{f-6-7-4}}.  Note that within
a month following the accident the first wall, blanket and shields are at
roughly the same temperature.

\begin{figure}[p]
\centerline{\psfig{file=spps-6-7-3.ps,height=3.5in,bbllx=69pt,bblly=99pt,%
            bburx=539pt,bbury=426pt,clip=}}
\caption[]
{Temperature at several points in the fusion core as a function of time assuming
adiabatic conditions at the vacuum vessel.}
\label{f-6-7-3}
\centerline{\psfig{file=spps-6-7-4.ps,height=3.6in,bbllx=69pt,bblly=104pt,%
            bburx=545pt,bbury=447pt,clip=}}
\caption[]
{Temperature distribution at several times following the accident.}
\label{f-6-7-4}
\end{figure}

Three additional adiabatic cases were considered.  First, it was found that the
dominant source of afterheat in the fusion core is the tenelon in the hot
shield.  The hot shield generates between 75\% and 99\% of the afterheat
thermal power in the fusion core.  Therefore, replacement of the tenelon filler
with vanadium alloy was considered. Second, in the event of an accident it
might be desirable to allow a gas such as helium to enter the evacuated magnet
insulation, thereby increasing its thermal conductivity by 2--3 orders of
magnitude.  By breaking the thermal barrier between the coils and modules in
this way, the thermal mass of the magnets could be added to the modules.  This
modification would slow the temperature rise in the modules at the expense of
the magnets.  It has been argued that, as the insulation heats up, vaporization
of material in the cryogenic insulation will cause this effect.  If so, then it
would be a passive mechanism, rather than an active one.  The results of these
two cases and a third which considered both at once are shown in
\hbox{Fig.~\ref{f-6-7-5}}.  Note that breaking the vacuum of the insulation
delays the temperature rise: 1,200\SdegC\ at 47 days, 1,500\SdegC\ at 67 days.
The vanadium hot shield ({\i.e.,} use vanadium also as filler material) exceeds
1,200\SdegC\ after 4~years.  In reality, even small unavoidable heat losses
over this time will prevent this temperature from being reached.
Unfortunately, the vanadium hot shield option would add approximately
15~mills/kWh to the cost of electricity.

As noted, the adiabatic conditions assumed above are overly conservative.  If
the fusion core is allowed to convect and radiate heat inside the containment
building enclosing the fusion core and this containment building can convect
and radiate to the environment, then the fusion core behavior changes.  A 1~m
thick concrete containment building was placed 2~m from the vacuum vessel (the
closer this is placed, the smaller the surface area and the greater its thermal
resistance; 2~m was considered the minimum reasonable) and these heat flows
were modeled.  \hbox{Figure~\ref{f-6-7-6}} shows the temperatures in the
fusion core as a function of time for these conditions.  As discussed above,
the environment is assumed to be at 20\SdegC, the heat transfer coefficient is
2~Wm$^{-2}$K$^{-1}$ from all surfaces and a factor of one-half is applied to
the vacuum vessel surface area and the containment building area and volume.
The maximum temperature in this case does not exceed 1,000\SdegC.

\begin{figure}[p]
\centerline{\psfig{file=spps-6-7-5.ps,height=3.2in,bbllx=69pt,bblly=99pt,%
            bburx=539pt,bbury=427pt,clip=}}
\caption[]
{First wall temperature as a function of time for several cases: the reference
case (see \hbox{Fig.~\protect\ref{f-6-7-3}}), a release of helium into the
magnet insulation at 1 day following the accident, the replacement of the
tenelon filler in the hot shield with vanadium alloy, and both of these
actions.  The boundary conditions are adiabatic at the vacuum vessel.}
\label{f-6-7-5}
\centerline{\psfig{file=spps-6-7-6.ps,height=3.3in,bbllx=69pt,bblly=99pt,%
            bburx=539pt,bbury=442pt,clip=}}
\caption[]
{Temperature at several points in the fusion core as a function of time
assuming heat losses through a containment building.}
\label{f-6-7-6}
\end{figure}

It may be possible to cool the gas in the containment building, thereby cooling
the fusion core more rapidly.  If this can be done, the temperature in the first
wall is reduced somewhat as is shown in \hbox{Fig.~\ref{f-6-7-7}}.
However, since radiation heat transfer depends strongly on temperature, the
improvement is not substantial.  The steady-state heat transfer rates from the
fusion core for these two cases are shown in \hbox{Fig.~\ref{f-6-7-8}}.
(Compare to \hbox{Fig.~\ref{f-6-7-2}}.)

\begin{figure}[p]
\centerline{\psfig{file=spps-6-7-7.ps,height=3.3in,bbllx=69pt,bblly=99pt,%
            bburx=539pt,bbury=442pt,clip=}}
\caption[]
{Temperature in the first wall and vacuum vessel assuming heat losses
through a containment building, with and without cooling of the containment
building gas (to 20\SdegC).}
\label{f-6-7-7}
\centerline{\psfig{file=spps-6-7-8.ps,height=3.3in,bbllx=76pt,bblly=95pt,%
            bburx=541pt,bbury=427pt,clip=}}
\caption[]
{Total heat losses from the stellarator in steady-state with and without
cooling of the containment building gas.}
\label{f-6-7-8}
\end{figure}

\clearpage

Without insulation between the cryostat box and the cold shield, the cryostat
box inner wall is warmed to 208\SdegC\ in normal operation.  As a result,
approximately 5~kW will enter the magnets by thermal radiation.  Perhaps
several times this value may conduct into the magnets via the support
structure.  Removing this heat will be expensive.  It would therefore be
advantageous to insulate, to some extent, the cryostat box inner wall.
Unfortunately, this would make heat removal more problematic in an accident.

Some insulation may be acceptable if the design takes advantage of the large
spaces between the magnets by creating open ports through the cryostat boxes.
Through these, heat can radiate directly from the cold shield to the vacuum
vessel.  If at least 10\% of the area of the cryostat boxes are open ports,
then the cryostats could be entirely adiabatic and still the fusion core would
not exceed 1,200\SdegC\ in an loss of coolant accident.  Such openings are
planned in each of the corner cryostats for the purpose of maintenance access;
these are 2~m $\times$ 5~m.  These 4 ports together constitute 2\% of the
approximately 2,000~m$^{2}$ the cryostats cover.  With adiabatic cryostats and
without additional ports, the modules would exceed 1,200\SdegC\ in 23 days and
1,500\SdegC\ in 40 days.

Perhaps the best solution to reducing magnet heating during normal operation
while providing for heat loss in an accident will involve placing several thin
closely-spaced layers of steel as insulation for the cryostat boxes.  During
normal operation, these radiation gaps will have a high thermal resistance but
if their temperature becomes elevated, this resistance will decrease.  This
trade-off of safety and operating cost needs optimization.

Temperatures in the winding packs and coil cases rise much more slowly than the
modules because the magnets are thermally isolated from the rest of the fusion
core.  However, by neglecting conduction in the magnet supports, the true rate
of temperature rise in the magnets is underestimated.

Under adiabatic boundary conditions, the loss of all coolant in the stellarator
would lead to serious economic damage if coolant were not restored in 26 days.
After 40 days without heat losses, the fusion core would pose an environmental
threat.  However, this boundary condition is over-conservative.  A conservative
but more realistic model which considers heat losses through the containment
structure never exceeds the temperature limits.  Efforts to reduce the magnet
heating by insulating the cryostat boxes need to consider the effect on safety.

\section{SUMMARY AND CONCLUSIONS}
\label{s-6-8}

The safety and environmental aspects of the SPPS/MHH have been considered.  The
well-established safety precautions for liquid metal use can make the
probability of a large lithium fire very low.

Here an ESECOM-type analysis, using extremely pessimistic assumptions, was done
to facilitate comparison with the previous ESECOM and ARIES tokamak studies.  A
regulatory-type analysis, taking credit for ameliorating factors, would yield a
more accurate assessment of off-site dose, but such an analysis was beyond the
scope of this work.

For a very pessimistic case with lithium drain failure, air ingress, complete
lithium burning, temperatures of 1,573~K (10~hours) followed by 1,273~K
(40~hours), and breach of containment with ground-level release under
pessimistic atmospheric conditions, the estimated site boundary Effective Dose
to the Maximum Exposed Individual would be about 7.5~Sv.  If the Porter dose
conversion factors were used, as in the ESECOM and ARIES tokamak studies, then
the potential dose from the stellarator power plant would be about 1~Sv, which
is comparable to doses found in those studies.

A regulatory-type of analysis, considering ameliorating factors, would
yield a more realistic assessment of off-site dose.  For example, the
estimated dose would be much lower if:

\begin{itemize}

\item
the ducts or chamber were not ruptured
\item
the lithium were drained into storage tanks
\item
the inert cover gas prevented a fire
\item
application of graphite microspheres extinguished the fire
\item
a fire persisted, but was partially controlled, resulting in lower
temperatures
\item
partial building integrity slowed down the leak rate
\item
credit were taken for aerosol deposition inside the building
\item
the release were through a stack
\item
less pessimistic atmospheric conditions prevailed
\item
credit were taken for aerosol deposition on the ground
\item
people at the boundary took shelter or evacuated.

\end{itemize}

A probabilistic analysis would show extremely low frequencies for severe
accidents that could cause hazardous doses off-site.  Similarly, the likelihood
of significant chemical toxicity effects being produced by release of
V$_2$O$_5$ is also very small.  Another attractive feature of the SPPS/MHH is
the possibility of recycling the vanadium structure and the lithium coolant.
The tenelon shield would probably need to be disposed as radioactive waste.

\clearpage

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\end{document}

\bye
