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\begin{document}

\title{\bf 1. OVERVIEW OF THE STARLITE STUDY }
\author{
\autlist{Farrokh Najmabadi}{Mark S. Tillack}{}
}
\date{}
\maketitle

\begin{titlepage}
{%baselineskip=1.8ex
\tableofcontents
}
\end{titlepage}

\setcounter{chapter}{0}
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\chapter{OVERVIEW OF THE STARLITE STUDY }

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

Fusion offers the promise of an abundant energy source that is not tied to
national resources and will have only a minimal impact on the environment.
However, fusion will be a new technology in the energy marketplace, it must
demonstrate clear advantages over other energy technologies at the time in
order to offset the inherent technical risk of a new technology or it will
never be widely endorsed.  During the past several years the ARIES Team has
investigated the feasibility and potential features of tokamak fusion power
plants~\cite{ARIESI,ARIESIII,ARIESIV,PULSAR}.  Building upon this research, the
Starlite project was initiated to develop the goals and requirements of fusion
demonstration and commercial power plants and to assess the potential of
tokamaks to meet these goals and objectives.  In addition, the research has
aimed at identifying both the trade-offs that lead to the optimal regime of
operation for a tokamak power plant and the critical plasma physics and
technology issues.

Based on interaction and advice from the U.S.\ electric utilities and industry,
a set of criteria for fusion power is derived~\cite{Naj95,StarI,StarF}. A
similar set of criteria has been developed by the EPRI fusion working
group~\cite{Kas94}.  These criteria and associated top-level requirements
and goals (Table~\ref{t-1}) can be divided in three general categories:
(1)~Cost, (2)~safety and environmental features, and (3)~reliability,
maintainability, and availability.

\begin{table}[bt]
\caption{}
\centering{TOP-LEVEL REQUIREMENTS AND GOALS }
\centering{FOR COMMERCIAL AND DEMONSTRATION FUSION POWER PLANTS}
\vspace{1.0cm}
\small
\renewcommand{\arraystretch}{1.2}
\begin{tabular}{l c c }\hline
Element			                  & Demo  & Commercial \\
\hline \\
Must use technologies 
to be employed in commercial power plant  &  Yes             & Yes \\

Net electric output must be greater than  & 75\% commercial  & N/A \\

Cost of electricity (COE)
must be competitive (in 1995 mill/kWh)    & 80 (Goal)        & 65 (Goal)  \\
                                          & 90 (Reqmt)       & 80 (Reqmt) \\
No evacuation plan
required for any credible accident:  \\
\quad Total dose at site boundary         & $< 1$~rem   & $< 1$~rem \\

Generate no rad-waste greater than        & Class C          & Class C \\
                
Must demonstrate public
day-to-day activity is not disturbed      &  Yes             & Yes \\
                
Must not expose workers
to a higher risk than other power plants  &  Yes             & Yes \\
                
Must demonstrate robotic
maintenance of power core                 &  Yes             & Yes \\
                
Must demonstrate routine
operation with less than \\
\quad (x) unscheduled shutdowns per year including disruptions &   1              & 1/10 \\

Demonstrate a closed tritium fuel cycle   &  Yes             & Yes \\
                
Must demonstrate operation
at partial load conditions at             & 50\%             & 50\% \\

\\ \hline
\end{tabular}
\label{t-1}
\end{table}

Top-level requirements and goals for cost of electricity (COE) were adopted for
the Starlite project based on estimated cost of competitive source of
electricity at the time of introduction of fusion in the market
place~\cite{Mil95}.  These requirements and goals for COE also are in line with
projections of future power plant costs based on energy forecasting
models~\cite{Cla95}.  Safety and environmental requirements are included to
circumvent the difficulties experienced by fission and, to some degree, will be
faced by fossil fuels in the future.  Fusion should be easy to license by the
national and local regulating agencies, and be able to gain public
acceptance. Fusion power plants should only generate low-level waste ({\it
i.e.,} waste storage time less than a few hundred years).  Realization of the
full safety and environmental potential of fusion will also help fusion to
achieve a cost advantage over other sources of electricity.  Fusion power
plants can be designed to achieve these criteria \us{only} through the use of
low-activation material and care in design.  However, these requirements result
in stringent constraints on the sub-system choices and design.  Lastly, it
should be demonstrated that Demo and commercial power plants can achieve the
necessary degree of reliability.  The conceptual design studies can partially
address this issue ({\it e.g.,} by including maintenance considerations in the
design). This criterion, to a large degree, should be addressed in the
development path of fusion power. Today's experiments are, by their charter,
and are not intended to provide detailed engineering data to support the
design, construction, and operation of a power plant.

During the Starlite study, an assessment of various tokamak plasma operation
modes as candidates for fusion power plants was made.  Five different regimes
of operation were considered: (1)~steady-state operation in the first-stability
regime, {\it e.g.,} \AI~\cite{ARIESI}, (2)~pulsed-plasma tokamak operation,
{\it e.g.,} Pulsar~\cite{PULSAR}, (3)~steady-state operation in the
second-stability regime, {\it e.g.,} ARIES-II and ARIES-IV~\cite{ARIESIV},
(4)~steady-state operation with reversed-shear profile, and (5)~low-aspect
ratio tokamaks (spherical tokamaks).  The extent of the plasma physics data
base as well as performance as a power plant were considered.  In
Sec.~\ref{s-1-2}, a brief summary of each physics regime of operation given.
In order to provide a common basis for comparing the cost of power plants, a
similar blanket and shield concept has been used.

In parallel, several options for engineering design ({\it e.g.,}
choice of structural material, coolant, breeder for first wall and blanket)
have been developed and assessed. In each area, this assessment was aimed at
investigating (1)~the potential to satisfy the requirements and goals, and
(2)~the feasibility {\it e.g., critical issues} and credibility ({\it e.g.,}
degree extrapolation required from present data base).  The engineering
assessment is overviewed in Sec.~\ref{s-3}. 

\section{ASSESSMENT OF TOKAMAK REGIMES OF OPERATION}
\label{s-2}

\section{First-Stability Steady-State Regime}
\label{s-2-1}

In the late 1980's, operation at high bootstrap current fraction as the
approach to steady-state operation was proposed by ARIES~\cite{ARIESI} and
SSTR~\cite{SSTR} studies simultaneously and independently.  In order to reduce
the current-drive power, the plasma current is reduced while the bootstrap
fraction is maximized.  In the first-stability regime, this can be accomplished
by operating with a moderately high plasma aspect ratio (\hbox{$A \equiv
1/\epsilon \sim 4.5$}) and low plasma current (\hbox{$I_p \sim 10$~MA}) at a
relatively high poloidal beta (\hbox{$\epsilon \beta_p \sim 0.6$}).  Detailed
MHD and current drive analysis have showed that the maximum bootstrap fraction
in this class is about $\sim$70\% with most of the driven current located near
the magnetic axis, leading to current-drive powers of about $\sim$100~MW
delivered to the plasma.

This mode of operation, however, leads to a low value of plasma
\hbox{$\beta\simeq 2\%$} because the important parameter $\epsilon\beta_p$
(which determines the bootstrap fraction) is related to achievable plasma
$\beta$ through
%
\begin{eqnarray}
\left( \epsilon \beta_p \right) \left( \beta / \epsilon \right)
 & \leq & \left( \beta_{N} / 20 \right)^{2} S
\label{e-1-2-1}
\end{eqnarray}
%
where $S=(1+\kappa^2)/2$ is the plasma shape factor, $\kappa$ is the plasma
elongation, and $\beta_N$ is defined by \hbox{$\beta \leq \beta_{N}
(I_p/B_{T}a_p)$}. For a conventional first stability configuration, optimally
shaped current and pressure distributions, and with sufficient triangularity in
the cross-sectional shape, this equation must be obeyed with a Troyon
coefficient of $\beta_{N}\simeq3.5$.  If a high $f_{BS}$ is desired for steady
state operation (such as \AI), one can reduce the current or raise $q_{*}$,
resulting in a lower $\beta$.  In order to optimize the equilibrium and
bootstrap current profile, these discharges operate with moderately elevated
central $q_0$.  For values of \hbox{$q_0\simeq 1.3$}, the discharges are stable
to kink modes without a conducting wall.  Higher plasma $\beta$ values can be
obtained by further increasing $q_0$ but these discharges would require a
conducting wall nearby and plasma rotation to stabilize the kink modes.  

There is ample experimental data base for this regime, however, operation in
discharges with durations longer than the current diffusion time is needed.

\section{First-Stability Pulsed-Plasma Regime}
\label{s-2-2}

Because the recirculating power for inductive current drive is small, it is
argued that pulsed-tokamak power plants avoid the problem of non-inductive
current drive: the constraint on $\beta_p$ is removed and plasma $\beta$ can be
higher.  However, inductive current-drive also imposes certain constraints on
the plasma operating regime.  First, because large and expensive poloidal-field
coils are needed to supply to inductive flux, these devices also tend to
optimize at moderately high-aspect ratio, low-current, and moderately high
bootstrap current fraction.  Second, because the loop voltage is constant
across the plasma, the current-density (induced and bootstrap) and pressure
profiles are set for a given pair of density and temperature profiles. In
principle, because the current-density profiles cannot be tailored, both the
normalized $\beta$ value and bootstrap fraction are limited (\hbox{$\beta_N
\sim 3.0$}, \hbox{$f_{BS} \lesim 40\%$}).  As a result, the plasma $\beta$ for
an optimum pulsed-plasma operation is only marginally ($\sim$25\%) larger than
that of a steady-state plasma.

In addition, many critical engineering issues have to be resolved.  For
example, cyclic fatigue and eddy-currents induced in the structural support of
the toroidal-field (TF) coils becomes major issues limiting the maximum
toroidal field ({\it e.g.,} the superconducting-coil technology that yields a
maximum field of 16~T in a steady-state power plant leads to a maximum field of
about 13~T in a pulsed-plasma system). This decrease in the toroidal-field
strength more than offsets the gains in plasma $\beta$ values for a pulsed
device, {\it i.e.,} even though the plasma $\beta$ is higher in a pulsed
device, the fusion power density is lower.  Another key issue is the energy
storage as the heat-transport system cannot tolerate rapid changes in the
thermal power of the power plant during recharge of the OH transformer.
Conventional designs for energy storage lead to large and expensive systems. An
innovative, sensible-heat storage system in the outer shield was developed for
the Pulsar design~\cite{PULSAR} which has removed the cost of energy-storage as
a barrier to pulsed-tokamak power plant.

As a whole, even assuming same reliability and unit cost for components, pulsed
power plants are substantially more costly than steady-state ones.  On the
other hand, this regime of operation is supported by the main body of tokamak
experimental data.

\section{Second-Stability Stead-State Regime}
\label{s-2-3}

As is seen from Eq.~(\ref{e-1-2-1}), the only way to increase the plasma
$\beta$ and still have a large bootstrap current fraction is to increase the
value of $\beta_N$. It may be possible to violate the Troyon limit if certain
conditions are met such as through elevated central safety factor ($q_{0} \geq
2$)~\cite{Ram91} as in \AII/IV designs, reversed magnetic
shear~\cite{Oze93,Kes94}, or at very low aspect ratios~\cite{Men95}.  In
addition to these requirements, stabilization of the external kink mode in
these regimes requires a close fitting conducting wall and sufficient plasma
rotation~\cite{Bon94}.

Second-stability operation ({\it i.e.,} operation at elevated values of central
safety factor) allows for higher bootstrap current fraction than
first-stability operation.  While bootstrap current can become as large as the
total plasma current (or larger), the bootstrap current-density profile is
different than that needed for equilibrium and stability.  As a result, current
should be driven both on the axis and at the plasma edge while cancellation of
part of the bootstrap current in the middle of the plasma is required.
Essentially, the optimum configuration has a bootstrap current fraction of
$\sim$90\% . Compared to the steady-state, first-stability optimum regime,
the current-drive power is somewhat smaller and the plasma $\beta$ is about
twice as large (a conducting wall for stabilization of kink modes is required).
The critical physics issues for this regime of operation include startup and
access to the second-stable operating point, as well as complexity of the
current drive system.  In addition, there is very little experimental evidence
for this mode of operation.

\section{Reversed-Shear Steady-State Regime}
\label{s-2-4}

Reversed-shear plasma operation combines the best features of steady-state,
first- and second-stability modes.  There is a much better match between the
bootstrap current-density profile and that needed for equilibrium and stability
and, therefore, higher values of plasma $\beta$ can be achieved with moderate
current-drive power and less complicated current-drive system.  The primary
characteristics of a reversed-shear plasma are a hollow current-density
profile, a non-monotonic safety-factor ($q$) profile, and relatively peaked
pressure profiles.  The hollow current-density profile gives rise to a
safety-factor profile, which initially decreases from its value at the plasma
center to a minimum value, and then rises to its value at the plasma edge.  It
is the initial decrease in the $q$-value away from the magnetic axis that
provides the negative magnetic shear, which is responsible for stability to
$n\rightarrow\infty$ ballooning modes.  It also appears that in this regime,
the plasma transport is suppressed and the a more peaked pressure profile that
is consistent with the high $\beta$ and high bootstrap-current fraction can be
sustained.  There is ample theoretical research on this regime and some
experimental data base is available. An extensive experimental exploration of
this regime of operation is currently on-going.

Analyses has showed that to zeroth order, the Cost of the device is independent
of the plasma aspect ratio in the range of \hbox{$A \sim 3-4$} (lower plasma
$\beta$ at the higher $A$ is compensated by higher toroidal-field strength on
axis and lower current-drive power).  In the Starlite study, an aspect ratio of
\hbox{$A=4$} was chosen because of engineering considerations.  At this aspect
ratio and with a plasma current of $\sim$11~MA, the maximum theoretical plasma
$\beta$ is $\sim$5.5\% and the bootstrap-current fraction is $\sim$90\%.
Non-inductive current drive is required to supplement bootstrap current on axis
({\it e.g.,} fast wave), at the plasma edge ({\it e.g.,} lower hybrid), and in
mid plasma.  While several candidate current-drive option exists for driving
the current in mid-plasma ({\it e.g.,} high-frequency fast wave, mode
conversion), the data base for these current-drive schemes is very small.

\section{Low-Aspect-Ratio (Spherical) Tokamak}
\label{s-2-5}

Operation at a low plasma aspect ratio (low-aspect ratio or spherical tokamaks)
is another approach to achieving high plasma $\beta$ and a high bootstrap
current fraction.  Unfortunately, the low aspect ratio of plasma rules out the
use of superconducting toroidal-field coils as there is not enough space for a
think shield in the inboard.  Therefore, a low-aspect-ratio tokamak requires
equilibria with a very high plasma $\beta$ (in order to minimize the Joule
losses in the normal-conducting center-post).  In addition, because the plasma
current is very large ($\sim$30 to 40~MA) in a low-aspect-ratio tokamak, a very
high bootstrap current which is also well-aligned to the equilibrium current
profile is essential (to minimize current-drive power).  Precise alignment of
bootstrap current density with the equilibrium current density is essential.
Detailed current-drive analysis has shown that even for a bootstrap-current
fraction of 95\%, the current drive power can easily exceeds several hundreds
of MW because the mismatch is usually at the plasma edge and mid-plasma where
current-drive systems are inefficient.

Since the power requirements associated with the edge current drive turned out
to be so large, we concentrated on a series of equilibria that require only
central current drive.  We imposes the constraint that the surface averaged
parallel current density be a linear function of the poloidal flux $\psi$
interior to the point where the linear profile is tangent to the bootstrap
current profile and equal to the bootstrap current outside this point
(Figure~\ref{f-1}). This constraint completely specifies the profiles and
insures that the first derivatives of equilibrium quantities are continuous
everywhere.  An extensive scan of MHD equilibria led to an equilibrium with
$A=1.25$, \hbox{$\beta\simeq 40\%$} and over 99\% of the plasma current is
self-driven (current-drive power is $\lesim$10~MW). This equilibrium has an
elongation of 3.0, which will certainly necessitate a vertical position control
system. It is noted that these cases would require a
conducting wall to stabilize the external kink mode.  

\begin{figure}[t]
\centerline{\psfig{figure=sti-8-2-2.ps,height=3.2in}}
\caption[sti-8-2-2]
{Parallel currents as functions of poloidal flux $\psi$ for a
high-bootstrap-fraction spherical tokamak equilibrium which does not require
edge current drive.}
\label{f-1}
\end{figure}

Using the case noted above, preliminary system studies and engineering analyses
were conducted.  A major conclusion of the system studies was that, even for
these cases with low current-drive power requirements, there still was
significant recirculating power ({\it i.e.,} recirculating power fraction
$\sim$0.6) associated with the resistive losses in the copper center post and
return legs of the toroidal field coils.  Because of the high-recirculating
power fraction, the low-aspect ratio tokamak is quite sensitive to choice of
plasma $\beta$, design of the center-post, and detail analysis of the current
drive requirements.  Other studies suggest lower values of recirculating power
fraction in a low-aspect ratio power plants.  

The spherical tokamak offers unique features and is currently pursued
vigorously in the physics community. However, the level of analysis has not
reached the stage needed to make sound assessment of this concept.  Many
critical issues has to be resolved such as: the sensitivity of the current
drive power requirements to variations in plasma profiles, center-post design,
start-up, design of in-vessel components which are subject to high heat and
particle loads, {\it etc.}  The true potential of a low-aspect ratio tokamak
power plant can only be assessed after a comprehensive design study.

\section{Systems Assessment of Tokamak Physics Regimes}
\label{s-2-6}

A system assessment of the above five tokamak plasma regimes has
been performed~\cite{StarI,Bat96}.  In order to provide a common basis for
comparing the cost of power plants, a similar blanket and shield concept has
been used which uses vanadium alloy as structural material of first wall,
blanket, and divertor and liquid lithium as the breeder which is a modification
of ARIES-II fusion core design.  It should be noted that it is not clear if
such a design can be utilized in a spherical tokamak ({\it i.e.,} most
probably, the performance of thermal power system would be lower because of the
higher wall loading and presence of water as center-post coolant).  In
addition, we have utilized a maintenance scheme for radial removal sectors
(similar to those of \AIV\ design~\cite{ARIESIV}).  This integrated sector
arrangement eliminates in-vessel maintenance operations and provides a very
sturdy continuous structure able to withstand large loads.  The penalty is the
increased size of the TF coils needed to allow adequate space for sector
removal.  Lastly, in order to minimize unscheduled interruption of plant
operation, all designs operate at 90\% of maximum theoretical $\beta$ in order
to avoid plasma disruption.  The major parameters of the five Starlite
candidates are shown in Table~\ref{t-2}.

\begin{table}
\caption{}
\centering{\bf Parameters for the Five Starlite Power Plant Candidates}
\vspace{1.0cm}
\renewcommand{\arraystretch} {1.3}
\begin{tabular}{l c c c c c } \hline
        &      FS   &    PU    &    RS    &    SS$^{(a)}$   &   LAR   \\
\hline
 & \hspace{1.2cm} & \hspace{1.2cm} & \hspace{1.2cm}
 & \hspace{1.2cm} & \hspace{1.2cm} \\
Plasma aspect ratio, $A=R/a_p$ & 4.0 & 4.0 & 4.0 & 4.0 & 1.25 \\
Major radius, $R$(m) & 7.96 & 8.68 & 5.04 & 6.40 & 5.00 \\
Plasma minor radius, $a_p$(m) & 1.99 & 2.17 & 1.26 & 1.60 & 4.00 \\
Plasma elongation, $\kappa_X$ & 1.81 & 1.80 & 1.99 & 2.03 & 3.40 \\
Plasma triangularity $\delta_X$ & 0.71 & 0.50 & 0.69 & 0.67 & 0.55 \\
Cylindrical safety factor $q_*$ & 3.77 & 2.40 & 2.37 & 4.60 & 3.54 \\
Central Safety factor,~$q_0$ &1.3&0.7&2.8&2.0&2.9 \\
Stability parameter, $\epsilon\beta_p$ & 0.54&0.32&0.56&1.22&1.14 \\
Normalized beta, $\beta_N(\%)$ & 2.88&2.70&4.76&5.28&6.42  \\
Beta parameter, $\beta/\epsilon(\%)$ &8.12&10.00&21.24&12.17&45.36 \\
Confinement ratio, $\tau_{He}^*/\tau_E$ &10.0&10.0&10.0&10.0&10.0 \\
Ignition parameter, $\beta\tau_E/a_p^2$(\%s/m$^2$)&0.78&2.18&4.17&1.35&11.05
\\
ITER-89P scaling multiplier, $H$ & 1.71&2.38 &2.40&2.47&3.02 \\
Normalized confinement multiplier, $H/q_*$ &0.45&0.99&1.01&0.54&0.85 \\
Plasma current, $I_p$(MA) &12.6&15.0&10.3&7.72&40.1 \\
Bootstrap-current fraction, $f_{BS}$ &0.57 & 0.34 & 0.89 &$>$1&0.997$^{(b)}$ \\
Current-drive efficiency, $\gamma_B(10^{20}$A/W~m$^2$) &0.56&NA&2.02&0.49&34.2 \\
Current-drive power to plasma, $P_{CD}$(MW) &236.6&0&64.7&199.1&7.3 \\
On-axis toroidal field, $B_T$(T) &8.99&7.46&7.35&8.37&1.77 \\
Peak field at TF coil, $B_{TF}$(T) &16.0&13.1&15.7&15.9&14.8 \\
Normalized heat flux, $P_{HEAT}/R$(MW/m) &71.2&29.5&71.3&89.0&124.2 \\
Recirculating power fraction, $(1/Q_E)$ &0.29 & 0.06 &0.13&0.33&0.63 \\
COE (mill/kWh)&99.7&130.2&69.7&92.6&116.0 \\
\\
\hline
\multicolumn{6}{l}{$(a)$ This design is not optimized to the lowest COE.} \\
\multicolumn{6}{l}{$(b)$ Includes diamagnetic current.}
\end{tabular}
\label{t-2}
\end{table}

\begin{figure}[p]
\centerline{\psfig{file=sti-5-2-02.ps,height=3.truein,clip=}\hspace{1.cm}
\psfig{file=sti-5-2-07.ps,height=3.truein,clip=}}
\centerline{\hbox to 3.truein{\hfil \bf (A) \hfil } \hspace{1.cm}
\hbox to 3.truein{\hfil \bf (B) \hfil }}
\vspace{0.5cm}
\centerline{\psfig{file=sti-5-2-03.ps,height=3.truein,clip=}\hspace{1.cm}
\psfig{file=sti-5-2-05.ps,height=3.truein,clip=}}
\centerline{\hbox to 3.truein{\hfil \bf (C) \hfil } \hspace{1.cm}
\hbox to 3.truein{\hfil \bf (D) \hfil }}
\caption[sti-5-2-02]
{(A)~MHD figures of merit for long-pulse ELMy, H-mode discharges in several
existing devices; (B)~Energy confinement figures of merit for a number existing
tokamaks and the Starlite fusion power plants. (Data provided by Stan Kaye,
PPPL.)  (C)~Large tokamak experiments which display substantial bootstrap
fraction at high poloidal beta.  Note that poloidal beta definitions vary
somewhat; and plasma conditions (collisionality, density gradient, impurity
content, {\it etc.}) also vary.  (D)~Fast-wave current-drive experimental
database on DIII-D~\cite{Deg95}, Tore Supra~\cite{Lit95}, and
TFTR~\cite{Rog95}; best data point for lower-hybrid current-drive on
JT-60~\cite{Ima90} and for lower-hybrid and FWCD synergy on
JET~\cite{Gor92}. The five Starlite fusion power plant options (FS, PU, RS, SS,
and LAR) are also shown.}
\label{f-2}
\end{figure}                                                 

In addition to economic performance, an assessment of maturity of each concept
was made.  To do so, several figures of merits were
identified~\cite{StarI,Mau95}.  The MHD figures of merits are $\beta/\epsilon$
and $\epsilon\beta_p$ as they indicate progress toward high $\beta$ discharges
with a high bootstrap current fraction.  The current-drive figure merit is
defined as \hbox{$\gamma_{B} = (\overline{n_e}/10^{20})I_pR/P_{CD}$} which is
the conventional measure of current drive efficiency but with the total plasma
current appearing to account for bootstrap effects.  The energy confinement
figures of merit are chosen to be $\beta\tau_E/a^2$ as a function of
$(\beta/\epsilon)^2$ to indicate progress towards discharges with high energy
confinement time and high plasma $\beta$.  Figure~2 shows that the existing
tokamak data base in two dimensional parameter spaces of these figures of
merit.  The five Starlite candidates are also shown.

An assessment of the five tokamak physics regimes of operation was made based
on the economic performance as summarized in Table~\ref{t-2} and based on the
maturity of data base as summarized in Figure~\ref{f-2}.  The first-stability
pulsed-plasma and steady-state regimes are closest to present data base.  Of
course these regimes should be demonstrated at long-pulse discharges with
burning plasmas.  On the other hand, the economic performance of pulsed-plasma
operation is poor. First-stability steady-state did not achieve the
economic requirements for the Starlite project.  High-field TF coils can
improve attractiveness of this regime of operation.  The second-stability
regime has a better economic performance but experimental data base for this
regime is very small.  The data base for the spherical tokamaks is not mature
and, in addition, many critical issues remain. Detailed design studies are
needed to before the true potential of spherical tokamak power plant can be
assessed.  The reversed-shear of operation offers the best economic
performance.  The data base for this regime, while small, is growing rapidly.
Based on the superior economic performance and the growing experimental and
theoretical data base, the Starlite project has chosen the reversed-shear as
the reference plasma operation regime for the ARIES-RS design~\cite{StarF}.


\section{ENGINEERING ASSESSMENT OF DESIGN OPTIONS}
\label{s-6-3}

\subsubsection{Introduction}

Through its successful operation, the U.S. Fusion Demo must be
sufficiently convincing that a utility or independent power producer
will choose to purchase one as its next electric generating plant.  A
fusion power plant which is limited to the use of currently-proven
technologies is unlikely to be sufficiently attractive to a utility
unless fuel shortages and regulatory restrictions are far more
crippling to competing energy sources than currently anticipated.  In
that case, the task of choosing an appropriate set of engineering
technologies today involves trade-offs between attractiveness and
technical risk.  The design space for an attractive DT tokamak fusion
power core is not unlimited; previous studies have shown that advanced
low-activation ferritic steel, vanadium alloy, or SiC/SiC composites
are the only candidates we have for the primary in-vessel structural
material.  An assessment of engineering design options has been
performed using these three materials and the associated in-vessel
component designs which are compatible with them.

The requirements described in Chapter 2 provide goals for engineering
components that are very challenging to meet.  Safety and
environmental requirements severely limit material choices.
Performance requirements provide strong incentives to operate at high
coolant temperature.  Reliability requirements may be the most
difficult to meet.  Design solutions must be simple, incorporate
adequate performance margins, and be tested fully.

For this assessment, detailed designs were not developed.  Instead,
various design options have been examined and their potential to meet
the power plant requirements assessed.  These options include material
choices for the structure, breeder and coolant.  In order to provide a
framework for this assessment, the three primary in-vessel structural
material classes are used to distinguish design classes.  The design
options using these materials are surveyed, paying special attention
to the first wall, blanket, shield and divertor, and the support
systems required to operate them.  For each class of designs, their
ability to meet power plant requirements are summarized and the key
issues are assessed.  Key issues are those uncertainties which, if not
adequately resolved, would lead to an unattractive product which does
not meet one or more of the top-level requirements.


\subsubsection{Assessment of Ferritic Steel designs}

\noindent{\bf A. Design Options}

Conceptual designs using ferritic steel as the structural material
have been proposed and developed for both solid and liquid breeder
materials, as well as water, helium and liquid-metal coolants.  For
most breeder, multiplier and coolant materials (except for Pb-Li),
ferritic steels are compatible up to the strength-based temperature
limit of the material (550\SdegC).  The design combination which
appears to be most appealing from a performance perspective uses
beryllium multiplier with solid breeder and is cooled by helium
flowing through tubes.  Reduced activation designs involve the use of
Li$_2$O or Li$_2$TiO$_3$ solid breeders and Fe-9Cr-2WVTa ferritic
steel.

\noindent{\bf B. Evaluation Highlights}

The evaluation focussed on the cost of electricity (COE),
maintainability without high-risk exposure to workers, and waste
disposal.  Factors influencing the COE are the material and
fabrication costs, the thermal efficiency achievable (limited by the
maximum structural temperature allowed), the peak surface and
volumetric heat loads allowed, and the lifetime.  More expensive than
austenitic steels, ferritic steels are only 20--25\% of the cost of
vanadium alloys.  In addition, they are design- code-qualified
materials for which there is an extensive database and industrial
experience.

The 550\SdegC\ temperature limit restricts the thermal efficiency, and
the peak surface heat load for a 5-mm wall is limited to $sim$0.8
MW/m$^2$.  This would restrict the peak neutron wall loading to
$\sim$4 MW/m$^2$ and result in a larger machine with higher capital
costs.  The surface heat flux limit also would make ferritic steels
unattractive as a divertor structural material.  It is desirable for a
structural material to tolerate $\geq$1~MW/m$^2$ of surface heat load.

Lifetime limits depend on operating stresses and irradiation effects.
By reducing the operating stresses, lifetimes of 12 MW-y/m$^2$ (3 full
power years) can be achieved with ferritic steels.  Preliminary cost
analyses indicate that the required cost of electricity ($\leq$90
mills/kWh) could be achieved with ferritic steel as a structural
material because the lower material and fabrication costs would tend
to off-set the cost associated with larger machine size and reduced
thermal efficiency.  These results assume that a different material is
used for the divertor.

Maintenance of any of the candidate structural materials would have to
be performed remotely because of the high level of radioactivity and
heat following shut-down.  In terms of waste disposal,
reduced-activation ferritic steels are better than austenitic steels
and on a par with vanadium alloys.  They qualify as Class C waste
according to NRC guidelines.

\noindent{\bf C. Issues}

The lower temperature limit for ferritic steels is not well
established.  Ferritic steels irradiated at temperatures
$\leq$400\SdegC\ tend to embrittle, as indicated by a decreasing
tensile ductility and an increasing ductile-to-brittle-transition
temperature (DBTT) with irradiation.  Experiments in mixed-spectrum
fission reactors result in DBTT values of about 250\SdegC\ under
conditions of simultaneous helium production and neutron damage, while
experiments at $\geq$365\SdegC\ in fast fission reactors with little
helium production give DBTT $<$ 0\SdegC\ for some of these steels.
The He production ($\geq$10 appm-He/dpa) would be relatively high in a
fusion power core.  Without fully understanding the role of helium in
raising the DBTT, it is difficult to extrapolate the fission reactor
results to fusion reactor conditions.  If the DBTT were to reach
250\SdegC, then the minimum operating temperature would have to be set
at 285\SdegC, leaving a relatively narrow temperature window for
design, and possible problems with start-up, shut-down and off-normal
events during shut-down.  Other problems, such as ferromagnetic
effects, were judged serious but probably solvable.

\subsubsection{Assessment of V designs}

\noindent{\bf A. Design Options}

A key design issue associated with V blankets is the compatibility of
V-alloy with either hydrogen or oxygen environments at higher
temperatures ($>$300--400\SdegC).  V has strong affinity for both
hydrogen and oxygen, and the material becomes brittle with moderate
hydrogen or oxygen concentration.  Lithium is one of the few materials
which has even stronger affinity to hydrogen and oxygen than vanadium
\cite{r2}.  Thus, it is natural to design the vanadium blanket with
lithium acting as both coolant and breeding material \cite{r3,r4}.  A
self-cooled blanket simplifies the blanket design.  However,
electrical insulation between the coolant and metallic walls may be
needed to reduce the MHD pressure drop, and the large lithium
inventory causes safety concerns.

To reduce or eliminate lithium safety concern, a He-cooled lithium
breeding blanket was examined in the BCSS project \cite{r5} and also
was proposed for the Demo \cite{r6}.  A closed-cycle gas turbine was
proposed for power conversion in the latter design.  The elimination
of the steam cycle reduces the source term for both hydrogen and
oxygen.  A surface modification approach is proposed to protect the
V-alloy, if necessary.  Another design has been proposed using Li$_2$O
as the breeding material to enhance even further the safety
characteristis \cite{r7}.  The use of SiC or Al$_2$O$_3$ for V-alloy
protection was proposed, but not discussed in detail.

\noindent{\bf B. Evaluation Highlights}

Vanadium-based design options have been evaluated relative to the
requirements and in all cases the component designs either meet the
requirements or are capable of meeting them, depending on the results
of complete system integration which is yet to be performed.

\noindent{\bf C. Issues}

For any fusion power core design, a key issue is associated with the
development of the structural material.  There is no intense 14-MeV
neutron source available.  Therefore, the material response to fusion
neturons can not be quantified.  For a more advanced structural
material, there is also a lack of industrial experience.  Therefore,
large scale fabrication and joining to other structural materials also
provides cause for concern.

For the self-cooled liquid metal design, the key issue is to develop
and assure long term reliability of the insulating coating.  Only
limited results in this area are available.  Although preliminary
results may be encouraging, some important effects, such as neutron
irradiation and thermal cycling, have not been included in
experiments.  Tritium recovery from lithium to the design goal of
$\sim$1 appm has yet to be demonstrated.  For a blanket with lithium
as the breeding material, safety is always a key concern.  Lithium
will react with water, air, concrete, nitrogen, {\it etc.}  The
chemistry of lithium can not be changed.  However, design of the power
plant can minimize the frequency and the consequence of the lithium
reaction when an accident occurs.

\subsubsection{Assessment of SiC Designs}

\noindent{\bf A. Design Options}

SiC composite material with a projected allowable temperature
capability of $\geq$1000\SdegC\ was selected as the structural
materials for ARIES-I and ARIES-IV reference and alternate blanket and
divertor designs.  These designs have been used as the basis for
assessing the advantages and issues associated with a possible power
plant based on SiC composites as the primary in-vessel structural
material.

To take full advantages of the low activation and high temperature
capability of SiC composite material, 5--10 MPa helium was used as the
coolant.  ARIES-I used Li$_2$ZrO$_3$ as the solid tritium breeder, Be
metal sphere-pac as the neutron multiplier, and W as the divertor
coating material.  Due to the induced activation of Zr and W, the
safety rating of this design was Level of Safety Assurance (LSA)-2.
Based on new experimental results, Li$_2$O was selected as the solid
breeder for the ARIES-IV designs.  If the use of W can be avoided as
the PFC coating material, then a completely passively safe design can
be obtained, if Be chemical energy and toxicity concerns are
eliminated.

Both ARIES-I and ARIES-IV alternate designs used the toroidal flow,
nested shell blanket configuration, where the SiC composite components
can be made from smaller ceramic (about 1 m $\times$ 1 m) parts.
Reliability of this configuration depends on the successful
development of ceramic joining techniques.  The ARIES-IV reference
design has a poloidal flow configuration.  This relies on the
successful development of the fabrication of large ($>$1 m wide,
$\sim$1 m deep and $>$7 m in height components) leak-tight SiC
composite components.

\noindent{\bf B. Evaluation Highlights}

Both ARIES-IV reference and alternate blanket designs satisfy all of
the neutronics and thermal hydraulic performance requirements.  Except
for the need of Be-neutron multiplier, both designs can have excellent
safety characteristics.  By increasing the helium pressure to about 12
MPa, and taking advantage of the improved recuperator performance of
the closed cycle gas turbine system at a helium outlet temperature of
about 950\SdegC, a gross thermal efficiency of about 55\% can be
achieved.  In the evaluation for the use of SiC-composite as
structural material, there was no doubt of its projected benefits; the
key question is whether its development schedule can meet the project
schedule of a Demo.

\noindent{\bf C. Issues}

Key issues for the application of SiC composite are in the areas of
material development, the behavior of composites in a fusion
environment, the need for metallic components to stabilize the plasma,
and the development of robust plasma facing components (PFC).

Fundamental improvements in material irradiated properties, such as
thermal conductivity, are needed, for example by using advanced SiC
fiber and interface materials.  Other material development issues
include economic fabrication of large SiC-composite components,
development of vacuum leak-tight components, the technique of brazing
ceramic parts, and the development of the joining techniques of
SiC-composite to metallic parts.  Under advanced physics scenarios
({\it e.g.}, the reversed shear mode), close conducting structures are
needed for plasma kink stabilization, and separate passive and active,
toroidally connected and electrically conducting rings are needed for
vertical stabilization.  This will determine the amount of metallic
elements that will be needed close to the plasma, thus impacting the
safety rating and thermal performance of the plant.  Various
techniques of leveling the surface heat flux, such as the use of the
radiative divertor approach is essential, since it dictates the
feasibility of removing surface heat flux by helium coolant and the
utilization of low activation material for the PFC surface and
components.  The concurrent reduction in the ion energy to $<$ 5 eV is
necessary to maintain low surface erosion and therefore, adequate
component lifetime and to maintain the low activation benefits of the
SiC composite design without the use of high-Z surface material.

\subsubsection{Neutronics Assessment}

A prime goal for the breeder is to provide tritium self-sufficiency.
Neutronics analyses have been performed for 20 breeders to examine
their ability to breed adequate tritium and to multiply the neutron
energy \cite{r8}.

Among the 6 liquid breeders and 14 solid breeders considered in the
analysis, lithium provides the highest breeding, followed by PbLi and
then Li$_2$O.  The most promising liquid metal (Li,
Li$_{17}$Pb$_{83}$) and solid ceramic (Li$_2$O, Li$_2$TiO$_3$)
breeders were selected for further analysis to illustrate the impact
on breeding of the candidate structures (V, SiC, and FS) and neutron
multipliers (Be, Be$_2$C, BeO, and Pb).  The study concluded that both
Li and Li$_{17}$Pb$_{83}$ have the potential for tritium
self-sufficiency without a neutron multiplier.  Solid breeders will
most likely require a neutron multiplier to achieve net breeding in
realistic designs.

All structural materials degrade the breeding and enhance the energy
multiplication, except SiC, that degrades both.  Vanadium has the
least impact on the breeding, followed by stainless steel and SiC,
except for Li, Li$_{17}$Pb$_{83}$.  For ceramic breeders, all
multipliers enhance breeding, except BeO, and the energy
multiplication, except Pb.

The blanket supplies all the tritium needed for operation (0.2--0.3
kg/d).  An external T supply is only needed to start and for a short
time until steady state production of T is reached.  To guarantee T
self-sufficiency, the uncertainties in all design elements should be
accounted for when estimating the breeding level for the various
blanket options.  The largest source of uncertainty is the basic
nuclear data and the calculational model ($\sim$10\%).  The calculated
overall tritium breeding ratio (TBR) should exceed 1.1 to assure that
the actual achievable TBR from the blanket after operation is $>$1.01.
In order to enhance the breeding capability of the blanket, it is
recommended to reduce the structure content, maximize the blanket
coverage, and locate penetrations off midplane as much as practically
possible.  Other options, particularly for blankets with marginal
breeding, include the use of neutron multipliers and/or enriching the
lithium.

A list of top level requirements was compiled for the Demo shield
(bulk, penetration, and biological).  The bulk shield will provide a
lifetime protection for the vacuum vessel and magnets, must have low
safety and environmental impact, and should be reliable, maintainable,
and replaceable.  The shield is a lifetime component and the
blanket/reflector will provide the lifetime protection for the bulk
shield.  As the bulk shield has major impact on the overall machine
size and cost, it should be optimized for high performance, low cost,
and minimal safety impact.  Steel has the best shielding performance,
followed by V and then SiC.  General directions were identified for
designing an efficient and cost-effective bulk shield.  These include
the use of steel structure for the outer layers of the shield and
employing higher performance materials in the inboard region and
cheaper but efficient shielding materials in the less constrained
outboard and divertor regions.

The safety features of breeders and structures will be a key factor in
selecting materials.  SiC composites possess the lowest activation and
afterheat characteristics, followed by V and then FS (modified HT-9 or
F82H alloys).  All candidate structural materials would qualify for
Class C low level waste according to NRC limits.  Safety-related
issues for the candidate materials are well documented in the
literature and design solutions can be identified.

\subsubsection{Summary}

Three candidate primary in-vessel structural materials -- ferritic
steel, vanadium alloy, and SiC composite -- have been examined to
determine if they can meet the requirements for an attractive power
plant and to characterize the remaining technical issues.

SiC composite has a unique potential for safety and high performance,
but the database requires substantial improvement in order to identify
and develop a satisfactory material composition.  Significant
improvements are needed in the basic properties, and several key
material issues such as joining and hermeticity must be resolved.  It
was judged premature to adopt such a material for Demo, which was
assumed to have a relatively near-term schedule.

Ferritic steel has the largest database, and hence the smallest
uncertainty in its performance.  However, as with all materials,
behavior after long-term operation in a complete fusion environment is
highly uncertain.  The features of ferritic steel which cause the
greatest concern is its restricted temperature window, which limits
the maximum achievable therml conversion efficiency, and its loss of
ductility.  It is not clear how one would operate a high-power device
with potentially brittle materials.  Economic studies suggest that the
lower cost of steel (as compared with vanadium alloy) does not
completely offset the lower power density and lower efficiency.

Vanadium offers significant advantages in its high temperature and
high thermal performance capability, and its low activation.  The
combination of V and Li appears particularly unique in material
chemistry, offering good compatibility up to very high temperatures.
The primary concerns with vanadium alloy are high cost and liquid
metal MHD effects (including the need for and feasibility of
insulating coatings).  Minimization of vanadium in the shield and
external systems is an important cost-reduction strategy.  The absence
of an established industrial base is also an important consideration,
but not one of the top-level requirements as elaborated in Chapter 2.
Due to its greater ultimate potential for attractive commercial power
plants and a development path which appears practical within a 25-year
timeframe, the combination of Li breeder/coolant and vanadium-alloy
structure was chosen as the next design concept to undergo full system
design and analysis in the ARIES-RS project.


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\end{thebibliography}
 
\end{document}
\bye


SPECIAL ISSUES

Several of the most critical aspects of a Demo power plant are given special
attention in this study.  These include (1) fusion economics and costing, (2)
safety and licensing, and (3) reliability, availability, maintainability and
inspectability (RAMI). These system-wide issues cross traditional borders
between physics and engineering of individual components and
systems. Historically, they have been given inadequate attention and/or have
been based on fission and other technologies without adequate updating and
adjustment for the unique attributes of a fusion power plant. Yet, they are
central issues which will play a primary role in determining the ultimate
acceptance of fusion as an electric energy source.

A. Fusion Economics and Costing

Fusion must be economically competitive at the time of its introduction into
the marketplace or no one will buy it. The economic targets for fusion Demo and
commercial power plants have been established based on extensive examination of
energy forecasting models and the expected trends for competing energy sources
[6]. The requirements specified in Table II resulted from these
investigations. One of the most difficult but important jobs of economic
forecasting is to include all relevant costs associated with an energy
source. Historically, many of the costs associated with power production (such
as R\&D costs, government subsidies, decommissioning, pollution of the
environment) have been "externalized" in the balance sheet, often resulting in
misleading conclusions. Current trends in deregulation of the industry and
expanded consideration of environmental impacts suggest that these costs will
have to be internalized by future decision-makers. Here we attempt to
explicitly identify these costs whenever possible.

In addition to economic projections, accurate costing of fusion power plants is
needed in order to determine whether or not the cost goals can be
met. Attention has been focused on a comprehensive review and updating, as
necessary, of specific unit costs and models used in the Demo Systems Code
(DSC). Use of Level of Safety Assurance (LSA) cost credits is under review as
it relates to distinctions between nuclear-safety-grade ("N-stamped") versus
conventional unit costs and labor rates.

B. Safety and Licensing

The greatest promise of fusion is its potential for enhanced safety and
environmental features. In addition to public acceptance and support, this
feature can be translated into ease of licensing and waste disposal which would
have a major impact on the cost of the power plant. These points have been
strongly emphasized by our Utility Advisory Committee and subsequently
reflected in the Demo goals and requirements.

A review of licensing procedures has been undertaken in coordination with the
Fusion Safety Steering Committee in order to provide a sensible licensing
scenario for the fusion Demo and power plants [7]. Information and necessary
guidelines for the licensing of fusion demo and power plants were provided by
the Starlite Team. By nature, this a long term task and should be continued
beyond this initial phase.

In addition to the efforts to develop appropriate licensing procedures, hazards
assessments are being performed for selected critical safety concerns. For
example, the U.S. stands alone in the world in its strong advocacy of liquid
lithium as a coolant and breeder, which offers the promise of very high
performance and simplicity which is expected to lead to enhanced
reliability. The chemical reactivity of lithium has given other programs reason
to rule out lithium a priori as an acceptable candidate. While measures have
been proposed for minimizing the hazards [8], more extensive hazards analysis
may be needed on specific designs in order to demonstrate its acceptability.

C. RAMI

Reliability goals for a commercial reactor appear extremely challenging to
meet. The systems are very complex, the operating environment for in-vessel
components is extremely harsh ({\it e.g.,} high heat flux, direct plasma exposure,
high energy neutron bombardment, strong magnetic fields), and finally, the
availability of test devices prior to Demo will be very limited due to the high
cost of an ignited fusion plasma device. In order to meet the overall cost
targets, an availability as high as 85\% may be required, and certainly
$>75\%$. Given the current state of knowledge, there is no evidence to believe
that this target can be met.

The RAMI data base for fusion is very limited, especially the data which is
applicable to full-scale fusion power plants. There are several reasons for
this situation. There are no operating fusion power plants or demos from which
to obtain data. The data which is available from operating R\&D facilities and
experiments have not been systematically collected, organized, and analyzed. In
any case, this data would have to be significantly extrapolated from present
R\&D experiments (which are scientific experiments and, by their nature, low
availability) to operating power plants.

Some efforts have been undertaken previously to predict failure rates of
individual components based on data from related technologies (see for example
[9]). A more broad-based fusion-relevant RAMI data base is being initiated here
by attempting to systematically collect data from operating fusion R\&D
facilities. By its nature, this task is long-term; continuous effort is needed
to maintain and update the database and interpret the results. Engineering
guidelines for improving RAMI numbers are an essential element and must be
integrated into the design process.

SUMMARY

The Starlite Project was initiated to investigate the mission, requirements and
goals, features, and R\&D needs of the Fusion Demonstration Power Plant. The
Fusion Demo should demonstrate to utility customers and the general public that
fusion power is a secure, safe, licensable, and environmentally attractive
power source that is ready for commercialization at an economically superior
total cost. To provide a basis for the design development, inputs from the
end-users were solicited and analyzed to yield a Demo mission statement ans a
set of top-level requirements. If this mission statement is achieved and the
requirements fulfilled, the fusion Demo power plant should exhibit all of the
attributes required for fusion to be a commercial success as a power plant heat
source.

\end{document}