# **5.9 Neutronics and Shielding** 

## **5.9.1 Introduction**

The design is in the preconceptual design phase where many different
design options and operating scenarios are being considered. DT pulses
with widths up to 20 s and fusion powers up to 150 MW producing a total
of 5 TJ of fusion energy are planned. In addition, DD pulses with
different widths and fusion powers up to 1 MW yield total fusion energy
of 0.5 TJ. The baseline design has a major radius of 2.14 m and an
aspect ratio of 3.6. The average neutron wall loading during the 150 MW
DT pulses is 1.85 MW/m^2^. The corresponding values at the outboard (OB)
midplane, inboard (IB) midplane, and divertor are 2.22 MW/m^2^, 1.67
MW/m^2^, and 1.11 MW/m^2^, respectively. It utilizes 16 wedged Cu TF
magnets. A double walled vacuum vessel (VV) with integral shielding has
been adopted. The plasma facing components include Be coated Cu first
wall (FW) and divertor plates made of tungsten rods mounted on
water-cooled Cu heat sink.

**5.9.2 Calculation Models**

Nuclear analysis has been performed to evaluate the impact of design
options and assess if the major performance objectives of the project
can be met without jeopardizing performance of the radiation sensitive
components. The neutronics and shielding calculations were performed
using the ONEDANT module of the DANTSYS 3.0 discrete ordinates particle
transport code system \[1\] with the most recent FENDL-2 nuclear
evaluated data \[2\]. Both the IB and OB regions were modeled
simultaneously to account for the toroidal effects. The machine is
assumed to have an operation schedule of four pulses per day with 3
hours between pulses.

It is assumed that 2% of the tritium generated in the DD pulses will
burn in the plasma producing some 14 MeV neutrons. These amount to only
2% of the 2.45 MeV neutrons produced in the DD pulses and are only 0.03%
of the 14 MeV neutrons produced in the 150 MW DT pulse. For the total 5
TJ DT fusion energy 1.77x10^24^ 14 MeV neutrons are produced. On the
other hand, for the total 0.5 TJ DD pulses 4.07x10^23^ 2.45 MeV and
8.15x10^21^ 14 MeV neutrons are produced.

The impact of different FW/tile design options on the nuclear parameters
was assessed \[3\]. The FW/tile design chosen consists of 0.5 cm plasma
facing component (PFC) (90% Be), 1.8 cm Cu tiles (80% CuCrZr) and 0.2 cm
gasket (50% Cu). A 2.5 cm water-cooled CuCrZr (15% water) VV cladding is
employed. A 1.3 cm space is provided between the gasket and VV cladding
and is devoted for diagnostics. It consists of wires, stainless steel
tubes, and MgO electrical insulation. A composition of 20% CuCrZr, 10%
steel, 10% MgO, and 60% void was assumed for that zone.

Detailed neutronics calculations were performed for the outer divertor
that is exposed to the most severe conditions in the divertor region.
The front layer is a 0.5 cm W Brush (90% W) followed by a 0.1 cm region
(84% W, 14% CuCrZr, 2% void) where the W rods are joined to the Cu heat
sink. The 1.9 cm heat sink is made of Cu finger plates (78% CuCrZr, 20%
water, 2% void). A 3 cm region (47% CuCrZr, 48% SS316, 5% void)
represents the mechanical attachment between the Cu finger plates and
the 7 cm thick backing plate (84% SS316, 16% water).

The VV consists of 1.5 cm thick inner and outer facesheets made of
316SS. The space between the VV facesheets (VV shielding zone) includes
60% 304SS and 40% water except in the IB region where 11% 316SS and 89%
water is used because of the small thickness. The thicknesses of the VV
at the IB midplane, OB midplane, and divertor region are 5, 54, and 12
cm, respectively. A 1.5 cm thick layer of thermal insulation (10%
Microtherm insulation) is attached to the back of the coil-side VV
facesheet. The Cu TF coils are included in the model with 90% packing
factor. While beryllium copper is used in the inner legs, OFHC copper in
utilized in the rest of the TF coils. A 316SS coil case is used in the
OB region with 4 cm front thickness and 6 cm back thickness.

3.  **Nuclear Heating**

Nuclear heating deposited in the different components was determined and
used in the thermal analysis. The calculations were performed for the DT
pulses with 150 MW of DT fusion power to determine the largest nuclear
heating generated. Nuclear heating results can be modified for lower
power pulses by scaling linearly with the fusion power. For the DD
pulses with the largest fusion power (1 MW), nuclear heating values are
at least two orders of magnitude lower than the values for the 150 MW DT
pulses.

Table 5.9.1 gives the peak power density values in the different
components at the chamber midplane. The largest power density values in
the magnet occur in the IB region at midplane. Fig. 5.9.1 gives the
nuclear heating distribution in the OB FW/tiles at midplane for the
baseline design. Nuclear heating was calculated also for the components
in the 1.3 cm diagnostics space behind the tiles. The nuclear heating
values in these components at midplane are 25.5 W/cm^3^ in MgO, 23.7
W/cm^3^ in SS and 24.9 W/cm^3^ in Cu. Fig. 5.9.2 gives the radial
variation of nuclear heating in the VV at the OB midplane. Nuclear
heating in the VV drops by an order of magnitude in \~18 cm. Table 5.9.2
lists the peak nuclear heating values at the outer divertor. Relatively
high nuclear heating is deposited in the W PFC. Fig. 5.9.3 shows the
nuclear heating distribution in the outer divertor plate.

**Table 5.9.1.** Peak Nuclear Heating (W/cm^3^) at Midplane

+-----------------------------------+-------------+--------------------+
|                                   | # IB        | OB                 |
+-----------------------------------+-------------+--------------------+
| Be PFC                            | 20.65       | 22.07              |
+-----------------------------------+-------------+--------------------+
| Cu Tiles                          | 29.08       | 28.71              |
+-----------------------------------+-------------+--------------------+
| Gasket                            | 25.17       | 25.17              |
+-----------------------------------+-------------+--------------------+
| ## Cooled Cu Vessel Cladding      | 24.92       | 24.86              |
+-----------------------------------+-------------+--------------------+
| H2O FWCoolant                     | 17.11       | 19.16              |
+-----------------------------------+-------------+--------------------+
| SS Inner VV Wall                  | 20.96       | 19.16              |
+-----------------------------------+-------------+--------------------+
| SS VV Filer                       | 20.40       | 17.67              |
+-----------------------------------+-------------+--------------------+
| H2O VV Coolant                    | 9.24        | 9.61               |
+-----------------------------------+-------------+--------------------+
| SS Outer VV Wall                  | 18.79       | 0.043              |
+-----------------------------------+-------------+--------------------+
| Microtherm Insulation             | 6.08        | 0.012              |
+-----------------------------------+-------------+--------------------+
| SS Inner Coil Case                | NA          | 0.024              |
+-----------------------------------+-------------+--------------------+
| Cu Magnet                         | 12.09       | 0.012              |
+-----------------------------------+-------------+--------------------+
| SS Outer Coil Case                | NA          | 1.73x10^-5^        |
+-----------------------------------+-------------+--------------------+

**Table 5.9.2.** Peak Nuclear Heating (W/cm^3^) at the Outer Divertor

+----------------------------------------------+-----------------------+
| #### W rods in divertor                      | 30.4                  |
+----------------------------------------------+-----------------------+
| Cu heat sink in divertor                     | 10.7                  |
+----------------------------------------------+-----------------------+
| SS structure in divertor                     | 9.2                   |
+----------------------------------------------+-----------------------+
| SS VV                                        | 4.2                   |
+----------------------------------------------+-----------------------+
| Cu Magnet                                    | 1.1                   |
+----------------------------------------------+-----------------------+

The amount of nuclear heating in the TF magnets strongly influences the
achievable pulse length. The radial variation of nuclear heating in the
IB magnet at midplane is shown in Fig. 5.9.4 for the DT and DD pulses.
Nuclear heating in the IB magnet drops by an order of magnitude in \~ 28
cm. The total nuclear heating in the 16 TF coils for 150 MW DT fusion
power was estimated based on the results of the 1-D calculations taking
into account the poloidal variation of neutron wall loading, shielding
thickness, and magnet toroidal coverage. Table 5.9.3 gives the breakdown
of total magnet nuclear heating. The total nuclear heating in the 16 TF
coils during the DT shots is 15.5 MW with 91% of it deposited in the
lightly shielded IB legs.

**Table 5.9.3.** Total Magnet Nuclear Heating (MW)

  ------------------------------------- ---------------------------------
  IB region                             14.2

  OB region                             0.03

  Divertor region                       1.3

  Total                                 15.5
  ------------------------------------- ---------------------------------

![](media/image1.wmf){width="4.766666666666667in"
height="4.024305555555555in"}

**Fig. 5.9.1.** Nuclear heating distribution in the OB FW/tiles.

![](media/image2.wmf){width="4.416666666666667in"
height="3.7402777777777776in"}

**Fig. 5.9.2.** Nuclear heating distribution in the OB VV.

![](media/image3.wmf){width="4.566666666666666in"
height="4.052777777777778in"}

**Fig 5.9.3.** Nuclear heating in the outer divertor.

![](media/image4.wmf){width="5.091666666666667in"
height="4.188194444444444in"}

**Fig 5.9.4.** Radial variation of nuclear heating in IB magnet.

**5.9.4 Vacuum Vessel Radiation Damage**

Since the VV is protected from the fusion neutrons by the thin FW/tiles,
the issue of reweldability was addressed. The end-of-life helium
production in the VV structure should be limited to 1 appm to allow for
rewelding. This is the limit used in ITER \[4\]. Table 5.9.4 gives the
results at different poloidal locations for the FIRE goal of cumulative
5 TJ DT and 0.5 TJ DD fusion energy. The contribution from DD shots is
very small (\< 0.15%). Lower VV He production occurs in the divertor
region as a result of shielding by the relatively thick divertor plate.
The results imply that reweldability of the VV should not be a concern.

####### **Table 5.9.4**. Peak End-of-life He Production (appm) in VV

  ------------------------------------------ ----------------------------
  IB midplane                                0.091

  OB midplane                                0.125

  Divertor                                   0.013
  ------------------------------------------ ----------------------------

# **5.9.5 Copper Radiation Damage** 

Table 5.9.5 gives the end-of-life peak dpa values in the Cu tiles, VV
cladding, Cu heat sink in outer divertor, and Cu TF coils for the FIRE
baseline design. Although the damage levels are very low, significant
effects on physical and mechanical properties might occur. These effects
are strongly dependent on irradiation temperature and have been the
subject of numerous studies as part of the ITER R&D program \[5\].

Radiation embrittlement at T \< 150°C is a concern for Cu alloys with
reductions in tensile ductility (uniform elongation) below 5% being
observed for damage levels on the order of 0.01 dpa. However, the
fracture toughness is typically maintained at a sufficiently high level,
at least in precipitation hardened alloys such as CuCrZr and CuNiBe. It
is possible to maintain the high tensile ductility by periodically
annealing the Cu at \~ 300°C for \~ 50 hr. Irradiation to ITER doses of
1-10 dpa at higher temperatures showed a pronounced increase in the
uniform elongation of CuCrZr compared with irradiation at lower
temperatures. However, at T \> 300°C this is accompanied by significant
softening. This was demonstrated by about an order of magnitude loss of
yield strength at about 300°C for 2-10 dpa.

Table 5.9.5. Peak End-of-life Cu dpa

  ------------------------------------ ----------------------------------
                                       Total dpa in Cu at end-of-life

  IB Tiles                             0.0271

  OB Tiles                             0.0298

  Divertor                             0.0125

  IB VV Cladding                       0.0178

  OB VV Cladding                       0.0204

  Magnet at IB                         0.0055

  Magnet at OB                         6.26x10^-6^

  Magnet at Divertor                   3.78x10^-4^
  ------------------------------------ ----------------------------------

Void swelling takes place in copper alloys irradiated in the temperature
range of 180 to 530°C. While void swelling is pronounced in copper
containing oxygen impurities (\~ 2.5%/dpa), it is only \~ 0.5%/dpa in
pure Cu and is generally insignificant in Cu alloys up to doses of 60
dpa. Therefore, for the low dose levels in FIRE, void swelling is not a
concern.

The effect of irradiation on the creep of Cu alloys is uncertain due to
limited data. The extremely low doses expected in FIRE reduce the
importance of irradiation creep. The magnitude of the irradiation creep
can be estimated using the creep compliance coefficient B. Using a
conservative B value of \~ 3x10^-6^ MPa^-1^dpa^-1^ and the peak
cumulative dpa value in FIRE (0.03 dpa), the irradiation creep for an
applied stress of 100 MPa would amount to a total deformation of only
10^-5^ (0.001%) at end-of-life in FIRE. It was recommended \[5\] that
the operation temperature of high strength Cu alloys should be limited
to \< 300°C for applied stresses of 100-200 MPa to have tolerable
irradiation and thermal creep at ITER conditions of 1-10 dpa. The
thermal creep strength begins to decrease rapidly for temperatures \>
300°C. This might cause deformation in the Cu during extended operation
(\> 100 hr) at 300°C. Due to the low doses in FIRE significant
deformation from irradiation creep is not anticipated. Some thermal
creep deformation in Cu alloys might occur if operated at elevated
temperatures (\>300°C). There is a lack of detailed studies on fatigue,
fracture toughness and fatigue crack growth rate behavior in
high-strength, high-conductivity copper alloys \[6,7\].

The Cu alloys operate at different temperatures in the FIRE components.
The tiles can get to temperatures over 400°C. The tiles carry no primary
stresses and should be basically unloaded except for thermal stresses
and disruptions. Therefore, problems with high-temperature softening and
creep should not be of concern. In addition, the tiles can be easily
replaced if needed. The temperature of the VV Cu cladding is lower than
250°C. At this peak temperature, occurring at midplane, the
low-temperature embrittlement for CuCrZr is not an issue. That will be a
concern only for the lower temperature parts of the cladding at the top
and bottom of the chamber. However, the dpa level will also be lower at
these locations resulting in alleviating the embrittlement concern. We
also have the option of annealing out the copper damage if we bake the
vessel to \> 300°C. The Cu in the divertor will have peak temperatures
close to 500°C. The peak damage level is only 0.015 dpa. The issue here
will be mainly thermal creep. The temperature of the TF coils rises from
80 to 373 K during each pulse. The main issue here is the
low-temperature embrittlement. The low temperature embrittlement data on
CuNiBe and OFHC Cu are limited to tensile tests between room temperature
and 100°C \[8,9\]. The concern is primarily at the IB midplane where the
peak damage rate is \~ 0.007 dpa, which is at the lower range of damage
for the occurrence of radiation embrittlement. Much lower damage levels
occur at other locations of the TF coil.

Based on the irradiation levels and operation conditions in FIRE and the
available data on Cu alloys, we can identify the R&D needs as follows:

-   Data on loss of ductility of BeCu (or OFHC) at temperatures between
    80 and 373 K with doses \< 0.01 dpa.

-   Data on fatigue, fracture toughness and fatigue crack growth rate
    behavior in high-strength, high-conductivity copper alloys.

-   Thermal creep data for CuCrZr at temperatures up to 500°C. There is
    no need to perform irradiation creep measurements on Cu alloys for
    the low doses proposed in FIRE.

**5.9.6 Radiation Induced Resistivity in the Cu Conductors of the TF
Coils**

The 17510 BeCu alloy is used in the inner legs of the TF coils with
10200 OFHC copper being utilized in the rest of the TF coils. A concern
with Cu magnet conductors is the increased electrical resistivity that
impacts the performance of the TF coils through increasing the I^2^R
heating and re-distributing the current across the coil. The temperature
at the end of the pulse will increase with possible impact on the
achievable pulse length. The increase in electrical resistivity results
from solute transmutation products and displacement damage. In a low
fluence machine like FIRE resistivity increase is dominated by point
defects and defect clusters produced by displacement damage.

The solute transmutation component of resistivity increase is directly
proportional to the solute content. We calculated the concentration of
transmutation products accumulating at the end-of-life as a function of
position in the TF coils. The BeCu alloy includes 1.8% Ni and 0.4% Be
and the transmutation products are dominated by Ni, Zn, Co, Fe, and H.
The worst conditions for the BeCu occur in the IB leg at midplane where
the peak cumulative dpa is 0.0055. For the OFHC copper, the worst
conditions are behind the divertor at the top/bottom of the machine
where the peak cumulative dpa is 3.78x10^-4^. Table 5.9.6 gives the peak
concentration of the transmutation products at these poloidal locations.
The Ni included in the unirradiated BeCu transmutes at a rate comparable
to the production rate of Ni from Cu resulting in negligible net
production of Ni. Fe is produced from the transmutation of Ni. The
resistivity increases associated with these concentrations were
calculated using the solute resistivities of 1.12, 0.3, 6.4, 9.3, and
1.5 pΩ-m/appm for Ni, Zn, Co, Fe, and H, respectively \[8\]. The results
are given in Table 5.9.6. Due to the high mobility of hydrogen in copper
at room temperature, it diffuses out of the material as the magnets heat
up at the end of each pulse. Thermal annealing is not expected to cause
any changes in the resistivity increase contributed by the other solute
transmutation products. Therefore, the largest resistivity increase from
transmutation effects is 0.783 pΩ-m in the BeCu alloy and 0.124 pΩ-m in
the OFHC copper. These values are much smaller than the unirradiated
resistivities in the temperature range of the FIRE magnets. In addition,
the neutron flux drops as one moves deeper in the magnet resulting in
decreasing transmutations and lower resistivity increase.

**Table 5.9.6.** Peak Resistivity Increase from Solute Transmutation
Products at Different Locations of TF Coils at End-of-life

<table style="width:100%;">
<colgroup>
<col style="width: 10%" />
<col style="width: 0%" />
<col style="width: 0%" />
<col style="width: 20%" />
<col style="width: 0%" />
<col style="width: 21%" />
<col style="width: 0%" />
<col style="width: 0%" />
<col style="width: 22%" />
<col style="width: 0%" />
<col style="width: 25%" />
</colgroup>
<tbody>
<tr class="odd">
<td colspan="3">Solute</td>
<td colspan="5">Peak solute concentration (appm)</td>
<td colspan="3">Peak resistivity increase (pΩ-m)</td>
</tr>
<tr class="even">
<td colspan="2"></td>
<td colspan="3"><p>BeCu alloy</p>
<p>IB midplane</p></td>
<td colspan="2"><p>OFHC copper</p>
<p>Behind divertor</p></td>
<td colspan="3"><p>BeCu alloy</p>
<p>IB midplane</p></td>
<td><p>OFHC copper</p>
<p>Behind divertor</p></td>
</tr>
<tr class="odd">
<td>Ni</td>
<td colspan="3">0</td>
<td colspan="2">0.0867</td>
<td colspan="3">0</td>
<td colspan="2">0.0970</td>
</tr>
<tr class="even">
<td>Zn</td>
<td colspan="3">1.602</td>
<td colspan="2">0.0591</td>
<td colspan="3">0.481</td>
<td colspan="2">0.0178</td>
</tr>
<tr class="odd">
<td>Co</td>
<td colspan="3">0.034</td>
<td colspan="2">0.0014</td>
<td colspan="3">0.217</td>
<td colspan="2">0.0090</td>
</tr>
<tr class="even">
<td>Fe</td>
<td colspan="3">0.009</td>
<td colspan="2">0</td>
<td colspan="3">0.085</td>
<td colspan="2">0</td>
</tr>
<tr class="odd">
<td>H</td>
<td colspan="3">0.261</td>
<td colspan="2">0.0147</td>
<td colspan="3">0.393</td>
<td colspan="2">0.0221</td>
</tr>
</tbody>
</table>

An analytical formula given by Eq. 1 \[10\] can be used to estimate the
resistivity increase due to displacement damage for copper and Cu
alloys. At high doses the displacement damage component approaches a
constant saturation value due to displacement cascade overlap effects.
The recommended value of the parameter B is 100 for both BeCu and OFHC
copper under FIRE operating conditions \[11\]. The parameter A
represents the saturation resistivity. Based on electrical resistivity
measurements, the expected saturation resistivity increase for pure Cu
irradiated near room temperature is 1.2 nΩ-m \[12\]. Higher resistivity
increases have been measured for Hycon CuNiBe \[6\] due to partial
dissolution of precipitates by displacement cascades, which injects
solute into the copper matrix. Based on available experimental data, we
used saturation values (parameter A in Eq. 1) of 1.2 and 4.2 nΩ-m for
OFHC and BeCu, respectively. These include the effect of annealing
during the pulse as the magnet warms up.

∆ρ \~ A\[1-e^-B\*dpa^ \] (1)

Periodic annealing of the copper components at temperatures well above
recovery Stage V (\~ 425 K, corresponding to thermal dissociation of
vacancy clusters \[13\]) would cause a significant additional reduction
in the displacement damage component of the resistivity increase beyond
that already achieved at room temperature. It is speculated that up to
\~ 90% of the displacement damage component of the resistivity increase
present in Cu specimens irradiated near room temperature could be
recovered by annealing near 573 K \[11\]. Hence, if the magnets of FIRE
can be baked-out to temperatures above 200°C (preferably 300°C), we can
significantly reduce the resistivity increase. We made the conservative
assumption that such bake-out is not employed.

Using Eq. 1 along with the cumulative atomic displacement values, we
determined the largest increase in BeCu resistivity to be 1.79 nΩ-m in
the inner leg at midplane. The largest increase in OFHC copper
resistivity is only 0.044 nΩ-m behind the divertor at the top/bottom of
the machine. Notice that these resistivity increases occur during pulses
near the end of the machine's life. Early in the life of the machine,
the accumulated dpa is very low and the resistivity increase is much
lower. In addition, the resistivity increase drops as one moves from the
plasma side of the coil deeper in the magnet.

Fig. 5.9.5 shows the resistivity increases at end-of-life from solute
transmutation products and displacement damage in the BeCu alloy as a
function of depth in the inner leg of the TF magnet at midplane. It is
clear that the total resistivity increase is dominated by displacement
damage with the resistivity increase from solute transmutation products
contributing less than 0.05% of the total resistivity increase. The
resistivity increase drops by a factor of \~ 30 across the magnet
thickness. Resistivity increase from solute transmutation products
contributes less than 0.3% of the total resistivity increase of the OFHC
copper conductor. The spatial distribution of the resistivity results in
re-distributing the current across the coil.

![](media/image5.wmf){width="4.641666666666667in"
height="3.9180555555555556in"}

**Fig. 5.9.5.** Resistivity change in BeCu as a function of depth in
magnet at IB midplane.

The unirradiated resistivity of the 68% IACS BeCu used in FIRE varies
from \~ 10 nΩ-m at 80 K to \~ 30 nΩ-m at room temperature \[14,15\].
This implies that the maximum increase in resistivity of the BeCu at
end-of-life varies from \~ 18% at the start of the pulse to \~ 6% at the
end of the pulse. The unirradiated resistivity of the 10200 OFHC copper
used in FIRE ranges from \~ 2 nΩ-m at 80 K to \~ 16 nΩ-m at room
temperature \[14,16\]. Hence, the maximum increase in resistivity of the
OFHC copper at end-of-life varies from \~ 2% at the start of the pulse
to \~ 0.25% at the end of the pulse. The unirradiated resistivities
given above are without applied magnetic field. Magnetoresistance
effects can lead to significant resistivity increase at cryogenic
temperatures of up to a factor of 10 depending on the magnetic field and
Cu purity \[17,18\]. However, much smaller magnetoresistivity increases
occur in copper components operating at the higher temperatures of FIRE.
The largest magnetoresistivity effects occur in the inner leg. For the
peak magnetic field of 15 tesla the peak magnetoresistivity increase in
the BeCu is \~ 7% at the start of the pulse and drops to \~ 2% at the
end of the pulse when the magnet heats up. The magnetoresistivity
effects also decrease as one moves deeper in the inner leg of the coil
due to the decreasing magnetic field. These increases are the same
during all pulses over the life of the machine. The magnetoresistivity
effects are lower than those resulting from irradiation effects during
pulses at the end of the machine life. However, during pulses early in
the life of the machine magnetoresistivity effects yield resistivity
increases larger than those resulting from irradiation effects.

**5.9.7 Magnet Insulator Dose**

The insulator dose rate in the TF magnet was calculated at the front
layer of the magnet winding pack. For 5 TJ of DT fusion energy and 0.5
TJ of DD fusion energy, Table 5.9.7 provides the peak cumulative magnet
insulator dose in the IB, OB, and divertor regions. The peak cumulative
magnet insulator dose is 1.05x10^10^ in the lightly shielded IB leg at
midplane. At this location the fast neutron fluence (E \>0.1 MeV) is
8.1x10^18^ n/cm^2^ and the total neutron fluence is 1.5x10^19^ n/cm^2^.
About 55% of the neutron fluence is above 0.1 MeV at the front of the
magnet and drops to 35% at the back of the magnet. The DD shots
contribute 13% of this value. The dose rate decreases by three orders of
magnitude as one moves poloidally to the OB midplane. The relative DD
contribution decreases as one moves poloidally from the IB midplane to
only 1.6% at the OB midplane due to increased attenuation of low energy
DD neutrons. The insulator dose decreases as one moves radially from the
front to the back of the winding pack as shown in Fig. 5.9.6. The gamma
contribution to the total dose is 35-50%. The dose decreases by an order
of magnitude in \~ 22 cm of the IB magnet. The radial variation of he
neutron and gamma fluence in the IB leg of the TF coils is shown in Fig.
5.9.7.

**Table 5.9.7.** Cumulative Peak Magnet Insulator Dose (Rads)

  --------------------------- ---------------------- ---------------------
                              Total Dose (Rads)      \% from DD Shots

  IB midplane                 1.05x10^10^            13%

  OB midplane                 1.05x10^7^             1.6%

  Divertor                    8.10x10^8^             10%
  --------------------------- ---------------------- ---------------------

![](media/image6.wmf){width="4.166666666666667in"
height="3.7333333333333334in"}

**Fig. 5.9.6.** Radial variation of insulator dose in the IB magnet.

![](media/image7.wmf){width="4.166666666666667in" height="3.66875in"}

**Fig. 5.9.7.** Radial variation of neutron and gamma fluence in the IB
magnet.

The mechanical strength, dielectric strength, and electric resistivity
are the important properties that could be affected by irradiation. The
shear strength is the property most sensitive to irradiation. The
commonly accepted dose limit for epoxies is 10^9^ Rads. This is the
limit used in ITER \[4\]. Polyimides and bismaleimides are more
radiation resistant with experimental data showing only a small
degradation in shear strength at dose levels in excess of 10^10^ Rads.
However, they are difficult and expensive to process due to their high
viscosity and requirement for high temperatures to fully cure. In
addition, they have initial mechanical properties lower than those
achievable with epoxies. Hybrids of epoxies, polyimides, bismaleimides,
and cyanate esters are being developed to both improve the insulation
system's ability to withstand high levels of radiation and to improve
the resin system's overall processibility \[19\]. That includes
irradiation measurements at dose levels close to those expected in FIRE
\[20\]. The information provided here gives guidance to the insulator
development R&D program regarding the dose levels and the proper mix
between neutron and gamma irradiation to insure relevance to the FIRE
conditions.

In the FIRE design with wedged coils and added compression ring, the TF
inner leg insulation does not have to have significant bond shear
strength, which is most sensitive to radiation. The peak torsional shear
stresses occur at the top and bottom of the IB leg behind the divertor.
The end-of-life insulator dose at these locations is reduced to \~ 10^9^
Rads due to the additional shielding provided by the divertor. Based on
that, it is expected that insulation materials will be identified that
can last for the whole device lifetime with the proposed operation
scenario and load conditions.

# 

# **5.9.8 Source Geometrical Effects on Peak Nuclear Parameters in FIRE TF Coils**

Neutronics calculations performed using a one-dimensional cylindrical
model that includes both the IB and OB regions account for the toroidal
effect and the impact of each region on the nuclear parameters in the
other. However, important three-dimensional effects are not included. In
the 1-D model the volumetric neutron source in the plasma zone is
assumed to be uniform and extended infinitely in the vertical direction.
In the real 3-D geometry, the neutron source has finite vertical
extension with a spatial profile that peaks near the plasma major radius
at the reactor midplane. As a result, the angular distribution of source
neutrons incident on the first wall will be quite different. In the 1-D
model most source neutrons are incident at a glancing angle compared to
the mostly perpendicular distribution in the exact 3-D geometry.
Consequently, the 1-D model is expected to yield higher nuclear
parameters near the first wall. The 3-D and 1-D results for ARIES-ST and
ARIES-AT were compared and the 1-D approximation was found to
overestimate the peak damage parameters at the first wall by up to a
factor of \~1.5 depending on the aspect ratio \[21\]. The IB legs of the
TF coils in FIRE are lightly shielded with relatively large nuclear
heating and insulator dose that impose restrictions on the achievable
pulse length and device lifetime. We performed calculations to assess
the impact of the source geometrical effects on the peak nuclear
parameters in the FIRE TF coils occurring at the IB midplane.

A two-dimensional model was developed for FIRE using the TWODANT module
of the DANTSYS discrete ordinates particle transport code system \[1\].
The 2-D model is adequate because of the axi-symmetric configuration
around the major axis. In addition, only the upper half of the machine
is modeled because of symmetry around the midplane. The plasma boundary
used is shown in Fig. 5.9.8. The neutron source density profile was
approximated by the parabolic formula given in Eq. 2.

S(r) = S(0) \[1-(r/a)^2^\]^4.75^ (2)

For the DT pulses with fusion power of 150 MW, the total neutron source
generated in the upper half of the plasma is normalized to 2.66x10^19^
14.1 MeV neutrons per second. The calculations were performed also for
the DD shots with 1 MW fusion power. In this case, in addition to the
2.45 MeV DD neutrons produced some 14.1 MeV DT neutrons will be produced
due to the 2% burnup of produced tritium. For the DD shots, the neutron
source in the upper half of the plasma is normalized to 4.08x10^17^ 2.45
MeV neutrons and 8.15x10^15^ 14.1 MeV neutrons per second. Since, the
primary interest is in the peak nuclear parameters at the IB midplane,
the detailed radial build in the IB region was modeled. On the other
hand, only an averaged homogenized material composition was used in the
OB and top regions.

![](media/image8.wmf){width="4.552083333333333in"
height="4.2868055555555555in"}

Fig. 5.9.8. Plasma boundary in FIRE.

Fig. 5.9.9 shows the radial variation of nuclear heating in the IB Cu
magnet at midplane during a 150 MW DT shot. The results are compared in
Fig. 5.9.10 to the results obtained from the 1-D calculation with an IB
neutron wall loading of 1.67 MW/m^2^ at midplane. The peak value is 9.25
W/cm^3^ which is \~23% lower than that obtained from the 1-D
calculations. Nuclear heating drops also as one moves away from the
midplane as shown in Fig. 5.9.11. The total end-of-life cumulative
insulator dose was calculated for a total of 5 TJ DT and 0.5 TJ DD
fusion energies. The radial variation is given in Fig. 5.9.12 at the IB
midplane. The contributions from the DT and DD shots are shown
separately. The results are compared to the 1-D results in Fig. 5.9.13.
The peak insulator dose is 8.71x10^9^ Rads which is \~17% lower than
that predicted by the 1-D model. This value drops by a factor of \~4 as
one moves to the top and bottom of the IB leg of the TF coil as
illustrated in Fig. 5.9.14.

Notice that the peak insulator dose of 8.71x10^9^ Rads is based on a
total of 5 TJ DT and 0.5 TJ DD fusion energies. We have 1.57x10^9^ Rads
per TJ of DT fusion energy and 1.7x10^9^ Rads per TJ of DD fusion
energy. These values can be used to determine the peak insulator dose
for any combination of DT and DD pulses. For the current baseline
experimental plan with 6.336 TJ DT and 0.049 TJ DD the peak insulator
dose is 10^10^ Rads. For the proposed extended operation schedule with
12.771 TJ DT and 0.185 TJ DD the peak dose will be 2.04x10^10^ Rads.

![](media/image9.wmf){width="4.4319444444444445in"
height="4.272222222222222in"}

Fig. 5.9.9. Radial variation of nuclear heating in TF coil at IB
midplane.

![](media/image10.wmf){width="4.319444444444445in"
height="4.138194444444444in"}

Fig. 5.9.10. Comparison between 1-D and 2-D magnet nuclear heating
results.

![](media/image11.wmf){width="4.372222222222222in"
height="4.242361111111111in"}

Fig. 5.9.11. Vertical variation of nuclear heating at front of IB TF
leg.

![](media/image12.wmf){width="4.042361111111111in"
height="3.9229166666666666in"}

Fig. 5.9.12. Radial variation of insulator dose at IB midplane.

![](media/image13.wmf){width="3.975in" height="3.9972222222222222in"}

Fig. 5.9.13. Comparison between 1-D and 2-D insulator dose results.

![](media/image14.wmf){width="4.304861111111111in" height="4.13125in"}

Fig. 5.9.14. Vertical variation of insulator dose at front of IB TF leg.

**5.9.9 Radiation Environment at FIRE Midplane Diagnostic Ports**

Neutron and gamma fluxes can affect plasma diagnostic performance
through enhanced conductivity of electrical insulation and scintillation
and absorption in optical components close to the tokamak. Determination
of the radiation environment is essential for estimating shielding
requirements for diagnostic components such as insulated cables,
windows, fiberoptics and transducers, as well as detectors and their
associated electronics. In addition, streaming through diagnostics
penetrations could lead to excessive doses outside the machine.

In FIRE, a few diagnostics, such as the neutral particle analyzer (NPA)
and impurity pellet guide tubes require straight holes through the port
shielding plugs. Other diagnostics, notably optical systems such as
Thomson scattering, will make use of labyrinthine penetrations to
curtail the streaming. Such penetrations will include four bends with
mirrors at the corners. A schematic of the diagnostics and penetrations
in the midplane diagnostics port plug J is shown in Fig. 5.9.15. This
port was identified as the most critical diagnostic port that requires
special attention regarding radiation shielding.

Two-dimensional neutronics calculations have been performed to determine
the nuclear radiation environment at selected locations in the
diagnostics penetrations and to assess the impact of streaming on the
average flux outside the port flange. The neutronics calculations have
been performed using the two-dimensional module of the DANTSYS neutral
particle transport code \[1\]. A simplified geometry was modeled in r-z
cylindrical geometry. In the model both the inboard and outboard regions
are modeled simultaneously to properly account for the toroidal geometry
effects. The calculations were performed for the DT pulses with 150 MW
fusion power using the midplane radial build for the FIRE machine with
2.14 m major radius. The front of the 110 cm thick port plug facing the
plasma is at a radius of 282.2 cm. The radial distance between the front
of the port plug and the 2.5 cm thick port flange is 339 cm. The port
plug and flange are assumed to consist of 80% steel and 20% water. Three
different models were used in the calculations. The first one assumes no
penetrations in the plug and flange. The results from this calculation
are used as a reference to quantify the impact of streaming. The second
case considered is for the worst case streaming with a 10 cm straight
penetration through the plug and flange. This is representative of the
NPA tube. The third case includes a penetration with four bends and
represents the Thomson scattering laser well. Due to the limitation of
two-dimensional modeling, all bends were modeled in the same plane as
shown in Fig. 5.9.16. Notice that conservative estimates are obtained
using these two-dimensional models since the penetrations are modeled as
slots that extend toroidally and attenuation by components in the
penetrations is not included. The total neutron flux (integrated over
all energies), the fast neutron flux (E \> 0.1 MeV) and the total gamma
flux were calculated at the front of the plug, along the penetrations,
at the back of the plug and at the back of the port flange. In addition,
the absorbed dose rates in silica (SiO~2~) and alumina (Al~2~O~3~), were
calculated. This includes the contributions from both neutrons and
gammas. The gamma contribution to the dose varies from \~30% at the
front of the port plug to \~80% at the back of the port flange.

![](media/image15.wmf){width="5.314583333333333in"
height="2.954861111111111in"}

Fig. 5.9.15. Schematic of diagnostics and penetrations in port plug J.

Fig. 5.9.16. Two-dimensional model used for the Thmpson scattering
penetration.

Table 5.9.8 gives the neutron and gamma fluxes for the case without
penetrations at the front of the port plug, at the back of the plug and
at the back of the port flange. The fluxes attenuate by about seven
orders of magnitudes in the shield plug. For the case with a straight
penetration the neutron and gamma fluxes along the penetration are given
in table 5.9.9. In addition, the average values at the three radial
locations are provided. At the back of the port plug, a flux peaking
factor of \~7 occurs due to the penetration. At the back of the flange
the peaking factor is only \~4 following neutron and gamma transport in
the large void space between the plug and flange. Notice that the
average flux values behind the plug and flange are about four orders of
magnitude higher than those in the case without the penetration. Hence,
using a straight penetration in the port plug increases the radiation
environment in the test cell area behind the flange by about four orders
of magnitude. This implies also that the biological dose rate will
increase accordingly.

For the Thompson scattering penetration with four bends, the fluxes and
absorbed dose rates were calculated at the mirrors located at each of
the bends. In addition, results are given at the front of the plug, back
of the plug, and back of the flange. Table 5.9.10 lists the neutron and
gamma fluxes along the penetration. The average values at the back of
the plug and at the back of the flange are given in table 5.9.11. At the
back of the plug, the peak values occur at the exit of the penetration
with a peaking factor of \~4. Since the penetration does not go through
the port flange, the peaking factor behind the flange is only \~2. The
average flux values behind the plug and the flange are about a factor of
200 higher than those without penetrations. Hence, the presence of the
4-bend penetration results in enhancing the radiation fluxes and dose
rates in the test cell area by about two orders of magnitude. Diagnostic
equipment in the penetration will reduce this enhancement. Table 5.9.12
gives the absorbed dose rates in silica and alumina along the 4-bend
penetration and at the back of the port flange. The dose rates in the
4-bend penetration at the back of the port plug are about two orders of
magnitude less than in the case of a straight penetration.

Table 5.9.8. Neutron and Gamma Fluxes during 150 MW DT Fusion Power
Pulses for Port Plug without Penetrations

+------------------------+-------------+----------------+-------------+
|                        | Total       | Fast Neutron   | Total Gamma |
|                        | Neutron     | Flux (E\>0.1   | Flux        |
|                        | Flux        | MeV)           |             |
|                        |             |                | (γ/cm^2^s)  |
|                        | (n/cm^2^s)  | (n/cm^2^s)     |             |
+------------------------+-------------+----------------+-------------+
| Front of port plug     | 8.38x10^14^ | 5.85x10^14^    | 4.94x10^14^ |
+------------------------+-------------+----------------+-------------+
| Back of port plug      | 3.51x10^7^  | 1.64x10^7^     | 3.04x10^7^  |
+------------------------+-------------+----------------+-------------+
| Back of port flange    | 1.04x10^7^  | 6.23x10^6^     | 1.04x10^7^  |
+------------------------+-------------+----------------+-------------+

Table 5.9.9. Neutron and Gamma Fluxes for the Case of Straight 10 cm
Penetration

+--------------+---------+-------+--------+-------+--------+-------+
|              | Total   |       | Fast   |       | Total  |       |
|              | Neutron |       | N      |       | Gamma  |       |
|              | Flux    |       | eutron |       | Flux   |       |
|              |         |       | Flux   |       |        |       |
|              | (n/     |       | (      |       | (γ/c   |       |
|              | cm^2^s) |       | E\>0.1 |       | m^2^s) |       |
|              |         |       | MeV)   |       |        |       |
|              |         |       |        |       |        |       |
|              |         |       | (n/c   |       |        |       |
|              |         |       | m^2^s) |       |        |       |
+--------------+---------+-------+--------+-------+--------+-------+
|              | Along   | Av    | Along  | Av    | Along  | Av    |
|              | Pene    | erage | Penet  | erage | Penet  | erage |
|              | tration |       | ration |       | ration |       |
+--------------+---------+-------+--------+-------+--------+-------+
| Front of     | 7.46    | 8     | 5.34x  | 5     | 4.14x  | 4     |
| port plug    | x10^14^ | .38x1 | 10^14^ | .85x1 | 10^14^ | .94x1 |
|              |         | 0^14^ |        | 0^14^ |        | 0^14^ |
+--------------+---------+-------+--------+-------+--------+-------+
| Back of port | 2.87    | 4     | 1.28x  | 1     | 2.08x  | 2     |
| plug         | x10^12^ | .06x1 | 10^12^ | .57x1 | 10^12^ | .79x1 |
|              |         | 0^11^ |        | 0^11^ |        | 0^11^ |
+--------------+---------+-------+--------+-------+--------+-------+
| Back of port | 5.14    | 1     | 2.75x  | 6     | 3.96x  | 1     |
| flange       | x10^11^ | .27x1 | 10^11^ | .38x1 | 10^11^ | .03x1 |
|              |         | 0^11^ |        | 0^10^ |        | 0^11^ |
+--------------+---------+-------+--------+-------+--------+-------+

Table 5.9.10. Neutron and Gamma Fluxes during 150 MW DT Fusion Power
Pulses for the 4-Bend Penetration

+------------------------+-------------+----------------+-------------+
|                        | Total       | Fast Neutron   | Total Gamma |
|                        | Neutron     | Flux (E\>0.1   | Flux        |
|                        | Flux        | MeV)           |             |
|                        |             |                | (γ/cm^2^s)  |
|                        | (n/cm^2^s)  | (n/cm^2^s)     |             |
+------------------------+-------------+----------------+-------------+
| Entrance at front of   | 7.48x10^14^ | 5.34x10^14^    | 4.15x10^14^ |
| port plug              |             |                |             |
+------------------------+-------------+----------------+-------------+
| First bend             | 1.20x10^14^ | 6.24x10^13^    | 7.68x10^13^ |
+------------------------+-------------+----------------+-------------+
| Second bend            | 2.06x10^13^ | 7.95x10^12^    | 1.37x10^13^ |
+------------------------+-------------+----------------+-------------+
| Third bend             | 2.57x10^12^ | 9.19x10^11^    | 1.79x10^12^ |
+------------------------+-------------+----------------+-------------+
| Fourth bend            | 2.87x10^11^ | 6.75x10^10^    | 2.15x10^11^ |
+------------------------+-------------+----------------+-------------+
| Exit at back of port   | 2.92x10^10^ | 6.82x10^9^     | 2.46x10^10^ |
| plug                   |             |                |             |
+------------------------+-------------+----------------+-------------+
| Peak at back of port   | 5.00x10^9^  | 1.50x10^9^     | 5.00x10^9^  |
| flange                 |             |                |             |
+------------------------+-------------+----------------+-------------+

Table 5.9.11. Average Neutron and Gamma Fluxes at Backs of Plug and
Flange with a Penetration with 4 Bends

+------------------------+-------------+----------------+-------------+
|                        | Total       | Fast Neutron   | Total Gamma |
|                        | Neutron     | Flux (E\>0.1   | Flux        |
|                        | Flux        | MeV)           |             |
|                        |             |                | (γ/cm^2^s)  |
|                        | (n/cm^2^s)  | (n/cm^2^s)     |             |
+------------------------+-------------+----------------+-------------+
| Back of port plug      | 6.98x10^9^  | 2.26x10^9^     | 5.97x10^9^  |
+------------------------+-------------+----------------+-------------+
| Back of port flange    | 1.99x10^9^  | 8.84x10^8^     | 2.10x10^9^  |
+------------------------+-------------+----------------+-------------+

Table 5.9.12. Absorbed Radiation Dose Rate in Silica and Alumina along
the 4-Bend Penetration

+----------------------------+-------------------+---------------------+
|                            | Dose Rate During  |                     |
|                            | 150 MW DT Pulses  |                     |
|                            |                   |                     |
|                            | (Rad/s)           |                     |
+----------------------------+-------------------+---------------------+
|                            | Silica            | Alumina             |
+----------------------------+-------------------+---------------------+
| Front of port plug         | 7.90x10^5^        | 8.48x10^5^          |
+----------------------------+-------------------+---------------------+
| First bend                 | 7.69x10^4^        | 9.85x10^4^          |
+----------------------------+-------------------+---------------------+
| Second bend                | 1.01x10^4^        | 1.76x10^4^          |
+----------------------------+-------------------+---------------------+
| Third bend                 | 1.24x10^3^        | 2.25x10^3^          |
+----------------------------+-------------------+---------------------+
| Fourth bend                | 1.41x10^2^        | 2.67x10^2^          |
+----------------------------+-------------------+---------------------+
| Back of port plug          | 1.65x10^1^        | 3.14x10^1^          |
+----------------------------+-------------------+---------------------+
| Back of port flange        | 3.80              | 7.07                |
+----------------------------+-------------------+---------------------+

Fig. 5.9.17 shows the vertical distribution of the fast neutron flux at
the back of the port plug for the three cases considered. While the flux
is nearly uniform for a plug without any penetrations, using a straight
penetration results in significant peaking at the exit of the
penetration. For the penetration with 4 bends, the flux peaks at the
exit of the penetration with a small local peaking behind the third
bend. The figure indicates also the relative enhancement in the average
flux level behind the port plug resulting from these penetrations.

Since the flux values in the test cell area behind the port flange are
increased significantly due to streaming, the biological dose rate will
be enhanced affecting the accessibility for hands-on clearing services
prior to removal/installation of port assemblies by remote handling
means. Acceptable biological dose rates behind the port flange were
obtained from activation calculations performed with a 1.1 m port plug
without any penetrations \[3\]. These dose rates are \~10 mrem/hr at
shutdown, \~3 mrem/hr after 1 hr, \~0.1 mrem/hr after 1 day, and about
0.05 mrem/hr after 1 week. These acceptable dose rates can be maintained
in the cases where penetrations are employed in the shield plug by
increasing the plug thickness. Past activation calculations indicated
that the dose rate scales linearly with the total neutron flux provided
that the material used in the area where the dose is calculated is not
changed. We performed two-dimensional neutronics calculations with
different plug thicknesses for the cases with straight and 4-bend
penetrations. Table 5.9.13 gives the results for 1.1 m plug and for the
case where all the port is filled by the plug. The peak and average flux
levels are shown behind the port flange. The results for the 1.1 m plug
without penetrations are included for comparison. For the straight
penetration, the flux peaking factor increases from \~4 with 1.1 m plug
to \~15 with 3.4 m plug. On the other hand, for the 4-bend penetration,
the flux peaking factor increases from \~2 with 1.1 m plug to \~13 with
3.4 m plug. The average neutron flux behind the flange decreases by an
order of magnitude by making the shield plug with straight penetration
thicker by \~0.5 m. The same drop in flux is obtained by \~0.42 m
thicker plug for the case with 4-bend penetration. The results imply
that if a straight 10 cm penetration is employed in the port plug, the
plug thickness needs to be increased to \~3.1 m to ensure that the dose
rates after shutdown in the area behind the port flange are similar to
the acceptable levels obtained without penetrations. For the case with
the 4-bend penetration employed, the plug thickness should be increased
to \~2.1 m. Since both penetrations exist in the diagnostic port J, it
is recommended that the port plug thickness should be increased to 3.4
m, the location of the cryostat interface.

![](media/image16.wmf){width="4.8069444444444445in"
height="4.584027777777778in"}

Fig. 5.9.17. Spatial distribution of fast neutron flux behind the port
plug.

Table 5.9.13. Neutron and Gamma Fluxes at the Back of the Port Flange
During 150 MW DT Fusion Power Pulses with Different Plug Thicknesses

+--------------+--------+--------+--------+--------+--------+--------+
|              | Total  |        | Fast   |        | Total  |        |
|              | N      |        | N      |        | Gamma  |        |
|              | eutron |        | eutron |        | Flux   |        |
|              | Flux   |        | Flux   |        |        |        |
|              |        |        |        |        | (γ/c   |        |
|              | (n/c   |        | (      |        | m^2^s) |        |
|              | m^2^s) |        | E\>0.1 |        |        |        |
|              |        |        | MeV)   |        |        |        |
|              |        |        |        |        |        |        |
|              |        |        | (n/c   |        |        |        |
|              |        |        | m^2^s) |        |        |        |
+--------------+--------+--------+--------+--------+--------+--------+
|              | Peak   | A      | Peak   | A      | Peak   | A      |
|              |        | verage |        | verage |        | verage |
+--------------+--------+--------+--------+--------+--------+--------+
| 1.1 m plug   | 1.04   | 1.04   | 6.23   | 6.23   | 1.04   | 1.04   |
| without      | x10^7^ | x10^7^ | x10^6^ | x10^6^ | x10^7^ | x10^7^ |
| penetrations |        |        |        |        |        |        |
+--------------+--------+--------+--------+--------+--------+--------+
| 1.1 m plug   | 5.14x  | 1.27x  | 2.75x  | 6.38x  | 3.96x  | 1.03x  |
| with 10 cm   | 10^11^ | 10^11^ | 10^11^ | 10^10^ | 10^11^ | 10^11^ |
| straight     |        |        |        |        |        |        |
| penetration  |        |        |        |        |        |        |
+--------------+--------+--------+--------+--------+--------+--------+
| 3.4 m plug   | 4.18   | 2.64   | 1.08   | 7.51   | 4.28   | 2.96   |
| with 10 cm   | x10^7^ | x10^6^ | x10^7^ | x10^5^ | x10^7^ | x10^6^ |
| straight     |        |        |        |        |        |        |
| penetration  |        |        |        |        |        |        |
+--------------+--------+--------+--------+--------+--------+--------+
| 1.1 m plug   | 5.00   | 1.99   | 1.50   | 8.84   | 5.00   | 2.10   |
| with 10 cm   | x10^9^ | x10^9^ | x10^9^ | x10^8^ | x10^9^ | x10^9^ |
| 4-bend       |        |        |        |        |        |        |
| penetration  |        |        |        |        |        |        |
+--------------+--------+--------+--------+--------+--------+--------+
| 3.4 m plug   | 9.64   | 7.04   | 1.97   | 1.63   | 1.59   | 1.33   |
| with 10 cm   | x10^4^ | x10^3^ | x10^4^ | x10^3^ | x10^5^ | x10^4^ |
| 4-bend       |        |        |        |        |        |        |
| penetration  |        |        |        |        |        |        |
+--------------+--------+--------+--------+--------+--------+--------+

**5.9.10 Future Work**

As FIRE proceeds beyond the current pre-conceptual design phase, more
sophisticated neutronics analysis will be needed. Detailed
three-dimensional modeling of the whole machine that includes material
heterogeneity, penetrations and gaps is required to the level of detail
used during the ITER engineering design activities. The model should
also account for accurate plasma shape and source profile. The machine
3-D model will be the basis for performing detailed 3-D neutronics
calculations to determine the different neutronics parameters (e.g.,
nuclear heating, radiation damage, insulator dose, etc.)

3-D modeling of ITER required a large amount of manpower because of the
need to manually generate the parameters for the surfaces and elements
in the geometrical input for the neutronics MCNP code. Significant
savings in the modeling effort can be achieved if direct interface
between the CAD drawings and the MCNP geometry module is developed.
Development of codes that facilitate such an interface will be very
valuable and will facilitate performing 3-D neutronics calculations with
the ability to quickly accommodate design changes. Activities developing
these interface codes were initiated at the University of Wisconsin
during the pre-conceptual design phase of FIRE and need to expand as
FIRE design proceeds to the conceptual design phase.

The 2-D analysis performed for the diagnostics port plug J showed that
streaming is be an important issue even with the relatively small 10 cm
penetration. 3-D analysis is essential to estimate the nuclear
environment and radiation damage to the sensitive diagnostic system
components (mirrors, fiber optics, critical electronics, etc.) in the
diagnostic ports of FIRE. The impact of streaming through the
diagnostics penetrations will have also to be evaluated due to its
impact on accessibility

for maintenance behind the cryostat.

The divertor region has a geometrically complex configuration that
mandates 3-D calculations. Important issues that need to be addressed
for that region are the amount of cryopump heating, nuclear heating and
radiation damage in the divertor components, and neutron streaming
through the divertor port.

The magnet insulator in FIRE is expected to experience significantly
higher dose than that in ITER. That puts premier on the need for
insulator R&D program to qualify insulators for the high dose level
expected in FIRE. In addition, we need to investigate the possibility of
using better shielding material (e.g., W, WC, B~4~C) to reduce the
insulator dose particularly if extended FIRE operation is feasible.

***[References for Section 5.9:]{.underline}***

\[1\] R.E. Alcouffe et al., "DANTSYS 3.0, One-, Two-, and
Three-Dimensional Multigroup Discrete Ordinates Transport Code System,"
RSICC Computer Code Collection CCC-547, Contributed by Los Alamos
National Lab, August 1995.

\[2\] M. Herman and H. Wienke, \"FENDL/MG-2.0 and FENDL/MC-2.0, The
Processed Cross-Section Libraries For Neutron-Photon Transport
Calculations,\" Report IAEA-NDS-176, International Atomic Energy Agency
(1997).

\[3\] M. Sawan and H. Khater, "Nuclear analysis of the FIRE ignition
device," Fusion Technolog*y*, vol. 39, p. 393, 2001.

\[4\] Technical Basis for the ITER Final Design Report, Cost Review and
Safety Analysis, ITER EDA Documentation Series, IAEA, Vienna, December
1997.

\[5\] S. Fabritsiev, S. Zinkle, and B. Singh, "Evaluation of Copper
Alloys for Fusion Divertor and First Wall Components," Journal of
Nuclear Materials, [233-237]{.underline}, 127-137 (1996).

\[6\] S.J. Zinkle and S.A. Fabritsiev, "Copper Alloys for High Heat Flux
Structure Applications," Atomic and Plasma-Material Interaction Data for
Fusion, Supplement to Nuclear Fusion, 5 (1994) 163-192.

\[7\] D.J. Alexander, S.J. Zinkle and A.F. Rowcliffe, "Fracture
Toughness of Copper-Base Alloys for Fusion Energy Applications," J.
Nucl. Mater., [271&272]{.underline} (1999) 429-434.

\[8\] S.J. Zinkle and W.S. Eatherly, "Tensile and Electrical Properties
of Unirradiated and Irradiated Hycon 3HP CuNiBe," Fusion Materials
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