Accident consequences analysis of the HYLIFE-II

# Inertial Fusion Energy Power Plant Design

S. Reyes^a,b,\*^, J. F. Latkowski^a^, J. Gomez del Rio^b^,and J. Sanz^b^

*^a^ Lawrence Livermore National Laboratory, P. O. Box 808, L-446,
Livermore, CA 94550,*

*Tel: 925-424-5684  Fax: 925-422-7390*

*^b^ Universidad Nacional de Educacion a Distancia, Escuela Tecnica
Superior Ingenieros Industriales,*

### Departamento Ingenieria Energetica, C/ Ciudad Universitaria, s/n, 28040 Madrid, Also: Instituto Fusion Nuclear, Madrid, Spain

**Abstract**

Previous studies of the safety and environmental (S&E) aspects of the
HYLIFE-II inertial fusion energy (IFE) power plant design have used
simplistic assumptions in order to estimate radioactivity releases under
accident conditions. Conservatisms associated with these traditional
analyses can mask the actual behavior of the plant and have revealed the
need for more accurate modeling and analysis of accident conditions and
radioactivity mobilization mechanisms. In the present work computer
codes traditionally used for magnetic fusion safety analyses (CHEMCON,
MELCOR) have been applied for simulating accident conditions in a simple
model of the HYLIFE-II IFE design. Here we consider a severe loss of
coolant accident (LOCA) in conjunction with simultaneous failures of the
beam tubes (providing a pathway for radioactivity release from the
vacuum vessel towards the confinement) and of the two barriers
surrounding the chamber (inner shielding and confinement building it
self). Even though confinement failure would be a very unlikely event it
would be needed in order to produce significant off-site doses. CHEMCON
code allows calculation of long-term temperature transients in the
fusion reactor first wall, blanket, and shield structures resulting from
decay heating. MELCOR is used to simulate a wide range of physical
phenomena including thermal-hydraulics, heat transfer, aerosol physics
and fusion product transport and release. The results of these
calculations show that the estimated off-site dose is less than 5 mSv
(0.5 rem), which is well below the value of 10 mSv (1 rem) given by the
DOE Fusion Safety Standards for protection of the public from exposure
to radiation during off-normal conditions.

*Keywords:* IFE safety analysis; Accident consequences; Off-site doses

\* Corresponding author.

*E-mail address:* <reyessuarez1@llnl.gov> (S. Reyes)

**1. Introduction**

The HYLIFE-II IFE power plant concept \[1\] sought to maximize its
attractiveness by providing lifetime components, high availability, low
cost of electricity and favorable S&E characteristics. During the
original HYLIFE-II study, S&E goals included:

-   low off-site dose from a severe accident

-   avoidance of nuclear grade (N-stamp) for most components (requiring
    less than 0.25 Sv off-site dose in the event of failure of that
    component)

-   working area dose rates less than 50 μSv/hr for low occupational
    risk

-   dose from routine atmospheric effluents less than 50 μSv/yr.

Since that time, the nuclear community (fission and fusion) has reached
a consensus and established more restrictive safety goals. The DOE
Fusion Safety Standards \[2\] for public protection in case of accident
give a limit off-site dose of 10 mSv (1 rem). This dose is a 50-year
committed effective dose to the most exposed individual (MEI) at the
site boundary (1 km) that only includes contributions from direct
cloudshine and inhalation during plume passage, because they pose an
immediate threat. This value is based upon a 1991 Environmental
Protection Agency document that specifies the level at which sheltering
and evacuation should be undertaken \[3\].

This work concentrates on severe accident analysis in order to compare
off-site obtained doses to the values provided by DOE Fusion Safety
Standards. The present work focuses on what has been considered to be
one of the most severe possible accidents in the HYLIFE-II design. We
assume a LOCA with loss of all the liquid flibe and simultaneous break
of all the beam tubes, which provides a flow path from the target
chamber to the confinement. In addition, breaks in both the inner
shielding wall and the confinement building wall are assumed. These
breaks provide a pathway for release of radioactive material to the
environment.

In order to perform a consequences analysis, first we need to know the
source term of radioactivity that is released to the environment.
Neutron transport and activation, thermal transients, and aerosol
transport calculations have been performed. Given the activity release
and the specific dose of each isotope, which has been evaluated with the
MACCS2 code \[4\], the accident dose may be calculated.

**2. Long-term thermal behavior**

Prior to the calculation of radioactivity release fractions, one needs
to know the time-temperature histories of the different reactor
components during the accident. This is used to determine the activation
products source term available for mobilization, which will later be
used as input data for the MELCOR calculations \[5\].

In this case the only energy source is the radioactive decay heat from
the activated material in the reactor components. Once the accident
begins, it is conceivable that fusion reactions could continue to occur
for a short period of time. Several factors must be considered:
accelerator operation, target injection, and beam propagation through a
varying environment. If accelerator operation and target injection both
continue, then beam propagation ultimately will determine when ignition
is no longer possible. We have assumed a complete loss of coolant
accident combined with multiple breaks in the beam tubes, shielding, and
confinement building. These breaks will allow air to enter the beam
tubes and chamber, and this air will result in the inability to
propagate the heavy ion beams to the target.

Langdon discusses the pressure of gas that is allowable for successful
beam propagation \[6\]. For an unneutralized beam, a background pressure
of more than 13 Pa (conservatively assuming lithium gas) guarantees that
stripping and plasma instabilities will destroy the beam. For a fully
neutralized beam, a pressure of about 2700 Pa will produce multiple
scattering, which will destroy propagation. If we conservatively assume
that the beams are fully neutralized, then fusion yield might continue
until enough air enters the beam tube to raise the pressure to 2700 Pa.

To determine the time required to reach 2700 Pa within the beam tubes, a
simple MELCOR calculation was performed. We see that 2700 Pa is reached
within only 50 ms, and thus, a plant operating at 6.4 Hz will not be
able to ignite the next target before beam propagation is halted due to
air pressure.

Considering then that decay heat from activated materials is the only
existing source available for heat transfer, a CHEMCON \[7\] calculation
was developed using a simple 1-D cylindrical model. This model consists
of four stainless steel shells that represent the tubes containing the
flibe, first structural wall, blanket structure and vacuum vessel, a
inner concrete shield and the confinement building wall. Conduction,
convection and radiation heat transfer was calculated between these
structures. The activation and decay heat results were obtained from the
ACAB \[8\] activation code following 30 years of plant operation. Fluxes
in structures were calculated with TART \[9\] Monte Carlo neutron
transport code. Figure 1 shows the thermal transient in the different
reactor components due to the decay heat from the activation products.

The first wall temperature, initially being 675 °C, displays a mild peak
at 679 °C at about 15 minutes, after which it gradually falls. This
indicates that the radioactive afterheat is low enough to avoid melting
of the stainless steel structure (T~melt~ \~ 1400 ºC) or any over-stress
damage. This result eliminates the possibility of volatilization of the
stainless steel structures during the accident, which would result in a
significant increase in the radioactive mobilized mass.

In order to protect the first structural wall from direct neutron
fluxes, the updated HYLIFE-II design considers a flibe pocket 60 cm
thick, which is the value used in our final accident analysis. A
parametric study was developed in order to evaluate the benefit of a
thick liquid wall on the activation of the structures and accident
temperatures. ACAB calculations were performed for different flibe
pocket thickness and irradiation times of 3 and 30 years irradiation.
Figure 2 shows the maximum accident temperatures obtained with CHEMCON
for the different cases.

It can be observed that an unprotected wall would not reach the melting
point either. This result is very important when accounting for the
stainless steel mobilized masses. Figure 2 also shows the benefit of
introducing a thick liquid protection. For example, note that for
thickness \> 15 cm, a 30-yr protected wall is better than a 3-yr
unprotected wall in terms of maximum accident temperature.

**3. Activation products source term**

There are four main sources of radioactivity that must be considered in
HYLIFE-II. First, the x-rays from each target vaporize a maximum of 10
kg of flibe. Although we assume a total LOCA, our analyses
conservatively include this flibe aerosol with its activation products.

Second, it is estimated that approximately 140 g of tritium would be
trapped within the chamber, blanket, and piping \[1\]. We assume that
entire tritium inventory is converted to the more radiotoxic HTO form.
This yields a mass of HTO aerosol of 930 g, which we round up to 1 kg.

Third, we assume that the corrosion of type 304 stainless steel (SS304)
by flibe within the chamber and blanket can be limited to 1 μm/y via
corrosion control methods. Additionally, we assume that the flibe
clean-up system can maintain the mobilizable inventory of corrosion
products to a 1-y supply. Given a total surface area of 1040 m^2^, we
obtain a corrosion product inventory of 8.3 kg in the total flibe
volume.

Finally, we use data from oxidation-driven mobilization experiments on
PCA performed at the Idaho National Engineering and Environmental
Laboratory (INEEL) to calculate an additional 0.5 kg of SS304 (we assume
that SS304 mobilization will be the same as that from PCA) that is
mobilized under exposure to steam at the accident temperatures
previously calculated \[10,11\]. Adding this 0.5 kg to the 8.3 kg of
corrosion products, we have \~ 9 kg, which we round up to 10 kg of
SS304. Considering that only about 5% of the flibe is present in the
chamber at any given time, we end up getting 0.5 kg of SS304 as
corrosion and oxidation products.

It is worth noting that we are being quite conservative in assuming that
the corrosion products will be in an aerosol form and available for
mobilization (one could make an argument that the corrosion products
would leave with the flibe during the LOCA). Additionally, our
assumption of steam oxidation driven mobilization is conservative --
because the secondary system uses water, it is credible that some water
could collect in low points of the plant, but the layout will be
designed to avoid water anywhere near the flibe ducts. At temperatures
of interest, oxidation-driven mobilization from PCA under exposure to
air occurs at a rate that is \~ 900× lower than that under exposure to
steam.

**4. Radioactivity release and off-site doses**

To estimate release fractions, a MELCOR model of HYLIFE-II was
developed. This model consists of four stainless steel shells, which
represent the SS304 tubes, first wall, blanket structure and vacuum
vessel, a concrete inner shield and the containment building. The beam
tubes were also modeled, considering an updated HYLIFE-II version that
includes 96 beams per side. Beam tubes and volumes within the vacuum
vessel are assumed to be at vacuum. The SS304 shells are initially at
their operating temperature of 675 ºC.

The radioactive source term described in the previous section is used as
input data for the masses of activated material that is available for
mobilization during the transient.

We have considered a severe accident consisting of a total LOCA, with
loss of all the liquid flibe, with simultaneous failure of the beam
tubes, the concrete inner shielding and the confinement. The total
failure of the beam tubes (the area of the break is modeled as the
calculated area of the total number of beams) means that the inner
volumes are going to have a flow path towards the outer volumes, which
are filled with air at atmospheric pressure. Both the inner shield and
confinement walls are initially at an estimated temperature of 32 ºC.
Additionally, we assume a double failure of the inner shielding and
concrete building (each of these breaks is 1 m^2^ in area) that provides
a pathway to the environment. This model is used by MELCOR to simulate
the progression of the accident.

The MELCOR heat transfer package considers conduction, convection and
radiation between the structures. The aerosol transport module treats
SS304 and flibe aerosol nucleation and agglomeration, vapor
condensation, gravity settling and gaseous/liquid transport. A new
module introduced by INEEL allows simulation of HTO transport and
condensation. Calculations of tritium migration from the stainless steel
shells were also developed using TMAP (Tritium Migration Analysis
Program) rather than assuming an instantaneous liberation from the
structures. These calculations showed that at the operating temperature
of 675 ºC, the tritium migration from the steel is fast enough (\>90% in
only 1.5 hour) that there is no difference when comparing with the
results from the instantaneous migration assumption \[12\].

The transient is considered to be one month long. Table 1 shows the
fraction of the source term masses that are released to the environment
during the accident, so as the final doses. It can be noticed that from
the initial 0.5 kg of SS304 that is mobilized, only 63 g are released to
the environment. Approximately 1.1 kg (out of 10 kg) of flibe is
released. Finally, due to partial condensation of the HTO on the colder
walls of the building, we get a release fraction of 86% of the initial 1
kg mobilized.

With these data about the activity released in units of Bq, and the
adequate dose conversion factors (DCF) in units of Sv/Bq, off-site doses
can be obtained. Of the total inventory in the reactor at the moment of
shutdown (just before the accident happening), all the radionuclides
whose contribution to the radioactivity release was above 3.7 x 10^7^ Bq
(1 mCi) were considered for dose calculation purposes. The DCF library
has been updated to include radionuclides that were missing from
traditional (fission) libraries as some of these radionuclides maybe
important for fusion doses. Data on DCF were obtained using the MACCS2
code in order to calculate early doses (defined as the equivalent
effective doses with 50 years commitment, resulting from the first 7
days of exposure during the plume passage). In addition to direct
cloudshine and inhalation during plume passage, our doses also consider
contributions from groundshine, and inhalation of resuspended material.
Typical weather conditions and ground level release were assumed as
recommended by the Fusion Safety Standards.

With these assumptions, the results from Table 1 show that the dominant
dose is from the tritium in HTO form, which gives a total of 4.6 mSv
(460 mrem). Both the SS304 and the flibe make small contributions to the
total dose, being 3.3 μSv (0.33 mrem) and 38 μSv (3.8 mrem),
respectively. If conservative weather conditions are assumed instead of
typical, then the calculated doses would be about 10× higher. On the
other hand, assuming an elevated release in typical weather conditions
would result in doses an order of magnitude lower than those calculated
in the case of ground level release.

It should be pointed out that about 75% of the SS304 dose comes from
^60^Co, and for the flibe, ^18^F is the only contributor to the dose.
From the total activity due to ^18^F at the moment of the reactor
shutdown (1.26 x 10^19^ Bq = 340 MCi), only 8.6 x 10^14^ Bq (0.023 MCi)
are vaporized and available for mobilization during the accident. The
final quantity of activity from the ^18^F that is released to the
environment is 10^14^ Bq (2.7 x 10^3^ Ci).

The final result is that even with the various conservatisms introduced
in the calculation, the off-site accident dose is under 5 mSv (0.5 rem).
This is a relevant result given that it is below the limit of 10 mSv (1
rem) given in the DOE Fusion Safety Standards, which means no sheltering
and evacuation are needed.

**5. Conclusions and future work**

The primary goal of accident consequence analyses is to integrate the
findings of the accident scenarios with the materials mobilization data
and dust/aerosol transport modeling to estimate radiation doses under
accident conditions. Doses and the consequences of those doses must be
calculated to enable direct comparison of IFE hazards with those
associated with fission and non-nuclear systems.

In the present work, accident doses have been analyzed for the HYLIFE-II
IFE design. A severe LOCA with failure of the beam tubes, and double
failure in the inner concrete shielding and confinement building has
been simulated using fusion safety codes.

The long-term thermal study of the various reactor components has shown
that the decay heat from activation products is low enough so that there
is no melting of structures during the accident and temperatures reach a
mild peak over the operational values before they start dropping.

The off-site dose released in this accident has been calculated to be
less than 5 mSv (0.5 rem). This result is well below the limit given by
the Fusion Safety Standards of 10 mSv (1 rem) during an off-normal event
in order to avoid public sheltering and evacuation. We can conclude that
the HYLIFE-II design has very favorable safety characteristics according
to this analysis.

Additional work is needed to identify and simulate other accident
scenarios in order to make a complete safety analysis for HYLIFE-II.
Accidents that do not involve a total LOCA will have to make use of the
equation of state of flibe, which will have to be introduced in MELCOR
code. Also other IFE designs must be considered and analyzed in a
similar way to have a wider perspective of the attractiveness of IFE
designs from a safety and environment point of view.

## Acknowledgements

The authors would like to thank their colleagues at the INEEL for their
assistance and recommendations that have been essential to the
completion of the above work.

Work performed under the auspices of the U. S. Dept. of Energy by
University of California Lawrence Livermore National Laboratory under
Contract W-7405-Eng-48 and under support of scholarship from UNED,
Madrid, Spain.

## References

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\[7\] M. J. Gaeta and B. J. Merrill, Idaho National Engineering and
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\[8\] J. SANZ, Universidad Nacional de Educacion a Distancia (UNED),
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\[9\] D. E. CULLEN, Lawrence Livermore National Laboratory,
UCRL-ID-126455, Rev. 2 (1998).

\[10\] K. A. McCarthy, G. R. Smolik, D. L. Hagrman, and D. A. Petti, *J.
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\[12\] Rion Causey and Thomas Venhaus, personal communication (2000).

Figure 1

![](media/image1.wmf){width="6.0in" height="5.2in"}

Figure 2

![](media/image2.wmf){width="6.020138888888889in"
height="4.307638888888889in"}

Fig. 1. Temperature evolution in the various reactor components during
the first day following the accident.

Fig. 2. Maximum accident temperatures reached at the first structural
wall for different flibe pocket thickness.

Table 1. Mobilized masses/activities, release fractions and off-site
doses for the different considered radioactivity source terms.

+----------------+-----------------------+------------+---------------+
| *              | **Mobilized**         | *          | **Off-site    |
| *Radioactive** | **mass** **/          | *Release** | dose**        |
|                | activity**            |            |               |
| **source**     |                       | **         |               |
|                |                       | fraction** |               |
+----------------+-----------------------+------------+---------------+
| **SS304        | 0.5 kg /              | 13%        | 3.3 μSv /     |
| corrosion /    |                       |            |               |
| oxidation**    | 1.31 · 10^12^ Bq      |            | 0.33 mrem     |
+----------------+-----------------------+------------+---------------+
| **Vaporized    | 10 kg /               | 11%        | 38 μSv /      |
| flibe**        |                       |            |               |
|                | 7.06 · 10^15^ Bq      |            | 3.8 mrem      |
+----------------+-----------------------+------------+---------------+
| **HTO trapped  | 1 kg /                | 86%        | 4.6 mSv /     |
| in steel**     |                       |            |               |
|                | 4.99 · 10^16^ Bq      |            | 460 mrem      |
+----------------+-----------------------+------------+---------------+