for the n_TOF Collaboration

# Design of the third-generation lead-based neutron spallation target for the
neutron time-of-flight facility at CERN

R. Esposito<EMAIL_ADDRESS>European Laboratory for Particle
Physics (CERN), 1211 Geneva 23, Switzerland  Thermomechanical Metallurgy
Laboratory, PX Group Chair, Ecole Polytechnique Fédérale de Lausanne (EPFL),
2002 Neuchâtel, Switzerland  M. Calviani<EMAIL_ADDRESS>O. Aberle  M.
Barbagallo  European Laboratory for Particle Physics (CERN), 1211 Geneva 23,
Switzerland  D. Cano-Ott  Centro de Investigaciones Energéticas
Medioambientales y Tecnológicas, Madrid, Spain  N. Colonna  Istituto Nazionale
di Fisica Nucleare, Bari, Italy  T. Coiffet  European Laboratory for Particle
Physics (CERN), 1211 Geneva 23, Switzerland  C. Domingo-Pardo  Instituto de
Física Corpuscular, CSIC-Universitat de Valencia, Spain  F. Dragoni  R.
Franqueira Ximenes  L. Giordanino  D. Grenier  European Laboratory for
Particle Physics (CERN), 1211 Geneva 23, Switzerland  F. Gunsing  CEA Saclay,
Irfu, Gif-sur-Yvette, France  K. Kershaw  European Laboratory for Particle
Physics (CERN), 1211 Geneva 23, Switzerland  R. Logé  Thermomechanical
Metallurgy Laboratory, PX Group Chair, Ecole Polytechnique Fédérale de
Lausanne (EPFL), 2002 Neuchâtel, Switzerland  V. Maire  P. Moyret  A. Perez
Fontenla  A. Perillo-Marcone  F. Pozzi  S. Sgobba  M. Timmins  V. Vlachoudis
European Laboratory for Particle Physics (CERN), 1211 Geneva 23, Switzerland

###### Abstract

The neutron time-of-flight (n_TOF) facility at the European Laboratory for
Particle Physics (CERN) is a pulsed white-spectrum neutron spallation source
producing neutrons for two experimental areas: the Experimental Area 1 (EAR1),
located 185 m horizontally from the target, and the Experimental Area 2
(EAR2), located 20 m above the target. The target, based on pure lead, is
impacted by a high-intensity 20-GeV/c pulsed proton beam. The facility was
conceived to study neutron-nucleus interactions for neutron kinetic energies
between a few meV to several GeV, with applications of interest for nuclear
astrophysics, nuclear technology, and medical research. After the second-
generation target reached the end of its lifetime, the facility underwent a
major upgrade during CERN’s Long Shutdown 2 (LS2, 2019–2021), which included
the installation of the new third-generation neutron target. The first and
second-generation targets were based on water-cooled massive lead blocks and
were designed focusing on EAR1, since EAR2 was built later. The new target is
cooled by nitrogen gas to avoid erosion-corrosion and contamination of cooling
water with radioactive lead spallation products. Moreover, the new design is
optimized also for the vertical flight path and EAR2. This paper presents an
overview of the target design focused on both physics and thermo-mechanical
performance, and includes a description of the nitrogen cooling circuit and
radiation protection studies.

††preprint: APS/123-QED

##  I Introduction and motivations

The neutron time-of-flight (n_TOF) facility at the European Laboratory for
Particle Physics (CERN) is a neutron source capable of providing high-
intensity pulsed white-spectrum neutrons covering almost eleven orders of
magnitude, from thermal neutrons to several GeV. Neutrons are produced via
spallation mechanism [1] from the interaction between a pure-lead target and
pulsed 20-GeV/c proton bunches from the Proton Synchrotron (PS) accelerator
ring at CERN. The generated neutrons travel inside vacuum tubes along two
flight paths directed to two experimental areas, Experimental Area 1 (EAR1)
and Experimental Area 2 (EAR2), where cutting-edge experimental setups for
neutron-induced reaction studies are in place. EAR1 is located at the end of a
horizontal beamline, 185 m from the spallation target, while EAR2 is located
at the end of a vertical beamline, 20 m above the target (Fig. 1).

Figure 1: The n_TOF Facility. A high-intensity proton beam from the CERN’s
Proton Synchrotron collides with a pure lead target, producing neutrons that
travel along two flight paths toward two experimental areas: EAR1 and EAR2.
EAR1 is located 185 m from the target, while EAR2 is located 20 m above the
target.

The facility was conceived to study neutron-nucleus interactions with
applications of interest for nuclear astrophysics, nuclear technology, and
medical research [2, 3, 4, 5, 6].

The facility operated with EAR1 and a first-generation spallation target
(Target #1) from 2000 to 2004 [7]. A second-generation target (Target #2)
operated from 2008 to the end of 2018. Even though the second target included
many upgrades if compared with the first one, both targets were cooled by
water in direct contact with pure lead [8].

For the first target, the cooling water was used as moderator. An independent
moderator assembly made from aluminum alloy (EN AW-5083 H111) was added to the
second-generation target to separate the cooling water from the liquid used
for moderation. The moderator liquid was then switched to borated water (1.28
wt% with 10B-enriched boron at 99%), which reduced the $\gamma$-ray background
for the neutron capture measurements in EAR1 [9] (the $\gamma$-ray background
in n_TOF is described in detail in Sec. II.3).

The second-generation target was designed considering an estimated lifetime of
ten years. In the last years of target operation, an increase in the cooling
water activity was detected, and preliminary signs of corrosion of the neutron
window were observed by endoscopic inspections [10]. Moreover, during 2014,
the second experimental area (EAR2) was built 20 m above the spallation
target, although the neutron beam characteristics in EAR2 were not optimal
since the target shape was not conceived for this experimental area.

All these aspects triggered the design of a new third-generation spallation
target (Target #3), to be installed during the Long Shutdown 2 (LS2), a three-
year stop of the CERN accelerators, and with the objective of starting
operation in 2021. High-purity lead has been kept as core material owing to
its superior performances in terms of reduced photon background with respect
to other spallation materials, due to its very low neutron capture cross-
sections.

Target #3 was designed with the assumption of a beam momentum of 20 GeV/c,
with a maximum number of protons per pulse equal to 1013 (equivalent to a
pulse kinetic energy of 32 kJ). The minimum pulse period is 1.2 s and the
maximum average intensity allowed is 1.67$\times$1012 p+/s, corresponding to
an average power on target of 5.4 kW. With a pulse duration of 7 ns (RMS),
this yields a peak deposited power of 1.8 TW. The beam size on target is
assumed to be 15 mm (RMS).

###  I.1 Pulsed white neutron sources

Several white neutron sources are operating worldwide, each of them with
specific characteristics and features, depending on the specific user requests
[11]. However, only a few of them are pulsed and thus optimized for time-of-
flight measurements. They are generally characterized by relatively high
instantaneous neutron flux with very reduced photon background. The GELINA
facility at the Joint Research Centre in Geel (Belgium) is based on a linear
electron accelerator impinging on a mercury-cooled uranium target [12, 13].
The Gaerttner LINAC Center at the Rensselaer Polytechnic Institute in the
United States employs a water-cooled tantalum target impacted by an electron
beam [14, 15]. Tantalum-clad tungsten targets are employed in the Lujan Center
at the Los Alamos Neutron Source Center (LANSCE) in the United States [16].

Neutron production targets based on lead are employed at the Paul Scherrer
Institute (PSI) in Switzerland [17] and in the neutron time-of-flight setup
nELBE at the Helmholtz-Zentrum Dresden-Rossendorf in Germany [18]. In the case
of the SINQ source, the lead is solid without beam but operates in liquid
state inside a Zircalloy cladding during operation [19, 20, 21], while in case
of nELBE Pb is in liquid state at all times. Solid lead targets for neutron
production are utilized for the Lead Slowing Down Spectrometer (LSDS) at the
LANSCE accelerator [22].

##  II Physics reach

The n_TOF facility has been at the forefront of neutron physics since 2001
[23]. The design of Target #3 aims to further expand the measuring
capabilities of the facility [8]. The combination of a high-energy and high-
intensity proton beam, long flight paths, specifically designed optical
elements along the beamlines, and an innovative spallation target results in
high quality neutron beams making n_TOF a unique facility. The main features
of the facility are summarized in Table 1.

Table 1: Key features of the n_TOF facility [9, 24]. Quantity |  EAR1 |  EAR2
---|---|---
Neutron flux (n/bunch) | 106 | 108
Energy range | min. | subthermal | subthermal
max. | 1 GeV | 100 MeV
Best resolution (E/E) | 10-4 | 10-3

Extensive Monte Carlo studies, performed by means of the FLUKA simulation
package [25, 26], were conducted in parallel with the engineering design of
Target #3 to help solving the issues of Target #2 mentioned in Sec. I and
further improve physics performance. A detailed and comprehensive FLUKA model
of the target assembly was developed, based on the engineering design
described in the next sections. The guiding goals of the Monte Carlo studies
were, compared to Target #2: (1) Preserve intensity and shape of the neutron
flux in the direction of EAR1. (2) Improve the resolution function for EAR2.
(3) Reduce the $\gamma$-ray background in EAR1.

The next subsections provide an account of these optimization studies and
their results.

###  II.1 Neutron fluence

The key feature for both experimental areas is the high instantaneous neutron
flux [27, 9, 24], of fundamental importance for measuring cross-sections of
radioactive samples or samples available in small quantities. The integrated
flux of 106 and 108 neutrons/bunch in EAR1 and EAR2, respectively, is due to
the high intensity and high energy of the primary proton beam and the overall
mass of the target (especially its lead component).

As shown in Fig. 2, Target #3 produces values of the neutron flux similar to
Target #2.

Figure 2: FLUKA simulation of the neutron flux in the two experimental areas
for Target #2 and Target #3\. The neutron flux in the energy range of interest
is similar in EAR1 and much better for Target #3 in EAR2, especially
considering the absence of absorption dips in the tens of keV range.

The shape of the neutron beam and, to some extent, its intensity, also depend
on the moderator systems. When compared to Target #2, Target #3 has a slightly
thinner layer of moderator liquid (water or borated water) crossed by neutrons
directed towards EAR1. This results in an additional contribution of non-
moderated neutrons at the evaporation peak (hundreds keV to several MeV) and,
consequently, in a slightly lower neutron flux in the thermal and epithermal
regions ($\leq$ 100 eV). Overall, the neutron flux reaching EAR1 with Target
#3 is 10% lower than with Target #2\. Regarding EAR2, Target #3 leaves the
neutron flux in the range 25 meV–100 keV essentially unchanged, while a factor
2-3 higher is expected at the evaporation peak. The higher flux is only
indirectly related to the new target design, as it depends mainly on the new
design of the vacuum window coupling the target to the beamline.

Moreover, the smaller amount of aluminum present in the structure of the
Target #3 ensures a substantial reduction of the related absorption dips in
the flux, as indicated by the black arrows in Fig. 2.

###  II.2 Resolution function

The neutron-induced reaction cross-sections under investigation at n_TOF
exhibit narrow resonances in the eV–keV region. The energy resolution of the
facility is of utmost importance to accurately determine the cross-sections in
this energy range. Energy resolution essentially depends on the flight path of
the two beamlines and on the neutron moderation. The resolution function can
be expressed as the uncertainties d$\lambda$ in the equivalent flight path
distance, and it is purely dictated by the geometry of the moderator. In the
energy range of interest it is characterized by a dominant peak having a full
width at half maximum (FWHM) of 2 cm, due to the pronounced slowdown in water,
and a relatively small but long tail (up to several meters), due to the
moderation in the lead core.

FLUKA simulations showed that the optimum, in terms of FWHM, can be reached
with water layers of 4 cm for both EAR1 and EAR2 moderators. While the new
design does not show any difference in the EAR1 resolution function when
compared to Target #2 [27], it significantly improves the scenario for EAR2.
In the Target #3 design, particular attention was paid in making the neutrons
traveling towards EAR2 pass through a uniform layer of moderator water. This
was not the case for Target #2, built before the construction of the second
beamline and therefore not optimized for it. Figure 3 shows the comparison of
the $\lambda$ distribution of neutrons in EAR2 in the energy range 1–10 eV for
the two targets, revealing a double-peak structure for Target #2 that led to
substantial limitations on the measurements in EAR2. The double-peak structure
is not present in the $\lambda$ distribution of Target #3, characterized by a
much better FWHM around the peak.

Figure 3: Distribution of the moderation length $\lambda$, in EAR2, for
neutrons in the energy range 1–10 eV. Target #2 exhibits a double-peak
structure leading to substantial limitations on the measurements in EAR2,
whereas Target #3 exhibits a single peak with narrower FWHM, which is better
suited for high-resolution measurements.

###  II.3 Background

A low background is a fundamental prerequisite to carry out high-quality
cross-section measurements. At n_TOF, the two main components of background
related to the spallation target are neutrons and $\gamma$-rays.

The neutron contribution is proportional to the neutron flux and also depends
on the optical elements placed in the downstream beamlines (i.e.,
collimators), so it is essentially unchanged in the new design. The
contribution of $\gamma$-rays is strongly dependent on the target assembly, in
particular on the moderator system and on the structural materials of the
target. The $\gamma$-rays reaching the experimental areas can be divided into
two contributions: prompt and delayed. The threshold between the two has been
set at 900 ns for EAR1 and 200 ns for EAR2.

The prompt $\gamma$-rays reach EAR1 within 900 ns and EAR2 within 200ns. Their
amount is directly proportional to the length of lead crossed by the primary
proton beam, so no sizable differences are expected in the new target design
with regard to EAR1. For EAR2, on the other hand, the coupling between target
and beamline has been improved with a new optimized shape of the vacuum
chamber, so a larger portion of the spallation target is seen by the
experimental area and the $\gamma$-ray contribution to the background
increases by a factor of 6. To compensate for the higher background, a 5-cm
thick lead plate has been integrated just above the target core and below the
EAR2 moderator. FLUKA simulations indicate that the addition of this lead
plate is enough to keep the prompt $\gamma$-ray contribution to the background
in EAR2 as low as for Target #2, with no significant reduction in the neutron
flux reaching the area.

The delayed $\gamma$-rays reach EAR1 and EAR2 at a time-of-flight longer than
900 ns and 200 ns, respectively. They are generated from radiative capture
reactions of neutrons in the hydrogen of the moderator water and in the
structural materials of the target. The former can be inhibited using borated
water (due to thermal neutron capture in 10B). For this reason, Target #3
includes two decoupled moderator circuits, with the possibility of changing
independently the circulating moderator liquid. In addition, considerable
effort was put into minimizing the structural materials of the target that are
sources of delayed $\gamma$-rays. Figure 4 shows the FLUKA results of the
delayed $\gamma$-ray background expected in EAR1 in comparison to Target #2,
revealing a background lower by a factor 4-5, in particular at the 2.2-MeV
energy peak and at the 7-MeV transition lines.

Regarding EAR2, $\gamma$-ray background arises mainly from the interaction of
neutrons with the collimator at the entrance of the experimental area and with
the beam dump in the roof of the building. Therefore, the design of the target
does not impact the delayed $\gamma$-ray background in EAR2.

Figure 4: Delayed $\gamma$-ray distribution in EAR1 for Target #2 (blue) and
Target #3 (red). A reduction by a factor 4-5 is visible in Target #3 at the
2.2-MeV energy peak and at the 7-MeV transition lines.

##  III Mechanical design

The new Target #3 assembly is shown, in exploded view, in Fig. 5: housed
inside a stainless steel vessel, six lead slices are cooled by gaseous
nitrogen and are supported by precisely machined anti-creep plates. These are
made from aluminum alloy and include the channels through which the cooling
fluid flows. The cooling gas is distributed through two main arteries inside a
cradle made from aluminum alloy, which supports the lead core from below.
Connected to the vessel, two moderator containers, made from aluminum alloy,
are positioned on the path of the neutrons directed to the two experimental
areas. The bond between the stainless steel vessel and the aluminum moderator
is obtained by an explosive-bonded joint. The lead slices are 5-cm thick, with
the exception of the slice close to the EAR1 moderator, which is 15-cm thick.
These values were selected to obtain a reasonable compromise between the total
target length and lead surface exposed to the coolant. FLUKA simulations
showed a conspicuous reduction in $\gamma$-ray background and an increase in
neutron flux by merging the 15 cm of lead closest to the EAR1 moderator into a
single slice. Figure 6 shows a photo of the target core inside the stainless
steel vessel before the proton side of the vessel was welded on to close it.

Figure 5: Exploded model of the target. The cooling nitrogen gas flows through
the channels machined into the anti-creep plates. The inner core is enclosed
in an AISI 316L stainless steel vessel. The two moderators are in aluminum EN
AW-5083 H112. The EAR1 moderator is bonded to the stainless steel vessel by a
bimetallic transition obtained by explosive bonding. Figure 6: Photo of the
n_TOF Target #3 core inside the stainless steel vessel before closing the
proton side of the vessel by TIG welding.

###  III.1 Lead core and anti-creep structure

The main target core is composed of six slices of high-purity lead supported
by an aluminum (EN AW-6082 T6) structure that includes anti-creep plates
between the slcies (Fig. 5). The lead grade is UNS L50006, with a minimum
purity of 99.98 wt%. The anti-creep aluminum structure also plays a
determinant role in the gas distribution thanks to channels machined to
distribute optimally the nitrogen gas between the lead slices.

To ensure a correct assembly of the anti-creep plates between the lead slices,
the gap between them must be within 45 µm and 195 µm. The minimum value takes
the thermal expansion of the lead slices during operation into account, while
the maximum value is required to avoid any negative impact on the nitrogen
flow. This requirement entails the anti-creep plates to be manufactured with
tight dimensional tolerances with respect to thickness (9.85 $\pm$ 0.05 mm)
and tight geometrical tolerances (flatness lower than 1 mm). The large size of
the plates (0.6 m $\times$ 0.6 m) made this requirement particularly difficult
to keep while milling the channels, since each milling step can bring
distortions to the plate. The accumulation of deformations along the same
direction was avoided by changing the side of the plate after each milling
step. Furthermore, a dedicated vacuum table holding the plates during the
milling procedure was manufactured, and the whole procedure was tuned and
tested by first building a prototype of the plate.

###  III.2 Vessel and Moderators design

The target core is contained in a stainless steel (AISI 316L) vessel designed
to hold a nominal nitrogen pressure of 0.5 bar. Made by first welding two
u-shaped parts, the upstream side is closed by a plate equipped with the
nitrogen inlet pipes, welded once the target core was inserted (Fig. 5). A
beam window was machined at the beam impact area by locally reducing the
vessel thickness down to 3 mm to avoid significant heat dissipation due to the
interaction between the proton beam and the vessel. An additional stainless
steel neutron window (4-mm thick) was welded by electron beam to the top of
the vessel and supports the lead plate that reduces the $\gamma$-ray
background in EAR2 mentioned in Sec. II.3. The two moderators are connected to
two independent water circuits at a nominal pressure of 2.5 bar. They are made
from forged aluminum EN AW-5083 H112. This alloy was chosen for its good
mechanical properties, their stability after welding, good general corrosion
resistance, and the ability to withstand the stresses induced by nitrogen
pressure and thermal gradients. The EAR2 moderator is located outside the
vessel, placed on top of the 5-cm lead plate above the target core. The
moderator is bolted to the vessel and sealed at the top edge to prevent
moisture building up between it and the lead plate.

The EAR1 moderator, the larger of the two, was manufactured by deep-milling
technology, which reduced the number of welds needed but required a rigorous
production-process qualification to ensure the necessary dimensional
tolerances. This moderator is an integral part of the vessel. The transition
between the aluminum alloy of the moderator and the stainless steel vessel is
a rectangular-shaped strip cut out from an explosive-bonded plate. The
challenge of such a solution lies in the qualification of the bimetallic
plates available from industry and on the design of the interfaces between the
different parts to ensure the desired gas tightness for the entire lifetime of
the target. The transition from aluminum alloy (EN AW-5083 H112) to stainless
steel (AISI 316L) includes two additional ductile layers of pure aluminum (EN
AW-1050A H14) and pure titanium (Grade 1) to ensure the good quality of the
explosive bonding process (Fig. 7). To validate the quality of the bond in
terms of impurities, voids, and mechanical resistance, CERN established strict
specifications and quality controls: the bond was subjected to helium leak
tests (max. leak rate accepted: 10-7 mbar L/s), and a dye penetrant inspection
was performed in accordance with the EN 10228-2 standard on 100% of the
machined surfaces. Ultrasonic tests were carried out for each explosive-bonded
interface according to EN 10288-4, along with metallographic inspections.

Figure 7: (a) Cross-cut of the explosive-bonded bimetallic transition between
the stainless steel target vessel and the aluminum EN AW-5083 H112 EAR1
moderator. (b) Schematic view of the explosive-bonded vessel-moderator
connection.

In addition to the nondestructive examinations, tensile tests and post-weld
metallographic investigations were carried out on representative samples
extracted from the original explosive-bonded plate. The tensile tests
confirmed a bond strength of 105 MPa, higher than minimum acceptable value of
80 MPa and limited by the strength of the weakest base material, aluminum EN
AW-1050A H14 (Fig. 8). The post-weld metallographic investigations confirmed
that, after electron-beam welding the explosive-bonded joint to the moderator,
the Heat-Affected-Zone (HAZ) is confined far from the interfaces of the
bimetallic transition within a safe distance of 3 mm, and thus it does not
affect them.

Figure 8: (a) Bimetallic-transition specimens tested in tension: for all of
them, the interfaces between the different materials of the transition
(stainless steel AISI 316L, titanium Grade 1, aluminum EN AW-1050A H14,
aluminum EN AW-5083 H112) were stronger than the weakest base material (pure
aluminum EN AW-1050A H14). (b) Stress–strain curves resulting from the tensile
tests: the ultimate bond strength is 105 MPa.

A first prototype of the explosive-bonded joint, 3-mm thick, failed the leak
validation. A second prototype was built to find a safe thickness value for
the joint: the prototype was designed to easily perform leak tests and reduce
thickness by milling. After each leak test, the thickness of the prototype was
reduced in steps of 5 mm, starting from 28 mm and down to 8 mm: no leak was
detected at any iteration. It was concluded that, rather than its thickness,
the main factor determining the soundness of a leak-tight joint is the choice
of the original region of the plate from where the final joint is extracted.
Regions that appear sound after preliminary dye-penetrant and ultrasonic
inspections are the best candidates. The final product was selected on this
basis, with a final joint thickness of 25 mm, the maximum permitted by
assembly and space constraints.

An innovative manufacturing process was developed to accommodate both
mechanical and physics requirements of the EAR1 moderator. Shaped like a short
cylindrical vessel, the two flat end surfaces, each 3 mm thick, would exceed
the maximum acceptable deformations and stresses induced by the 2.5 bar
pressure without additional stiffening elements. The use of a precise deep-
milling manufacturing technique offered the possibility to add such stiffeners
while preserving the monolithic nature of the part and avoiding welds, which
would increase failure risks. Up to 80% of the material was removed from a
solid aluminum disk, with the remaining 20% shaped to obtain deformations and
stresses compliant with the specification. The resulting monolithic center
part of the moderator was then welded to the outer housing by electron-beam
welding. The outer housing includes the interface with the explosive-bonded
joint (Fig. 9).

Figure 9: Schematic view of the deep milling procedure for the EAR1 moderator
manufacturing. The central part is entirely machined from a single solid
aluminum block (EN AW-5083 H112). It is then bonded to the housing via
electron-beam welding.

Special effort was put into the selection process of the bulk material for the
disk in terms of homogeneity and grain size, low number of defects, and
impurity. The choice fell on 3D forged aluminum EN AW-5083 H112 blocks for its
good compliance with the specifications and stability during machining.
Compliance with the specifications was checked before machining by dye-
penetrant and ultrasonic inspections. In addition, the qualification of the
electron-beam welding process was carried out on representative samples. To
complete the qualification process of the EAR1 moderator, vacuum and pressure
tests were successfully carried out to ensure the compliance with pressure
vessel safety standards.

##  IV Nitrogen cooling system: Computational Fluid Dynamics

Computational Fluid Dynamics (CFD) studies were carried out to support and
optimize the design of the target. The optimization consisted of maximizing
the cooling efficiency (minimizing the maximum steady-state temperature) while
minimizing the pressure drop and the locations of high flow velocity inside
the target. The CFD calculations also provided input for the thermo-mechanical
analyses (see Sec. V) and for the design of the cooling station, (see Sec.
VI).

The commercial CFD software Ansys CFX [28] has been used throughout the
studies to model simultaneously the solid (lead slices) and the fluid bodies
(nitrogen), and perform conjugate heat transfer calculations. The map of
average beam power deposition, calculated with FLUKA Monte Carlo simulations,
was imported as a power-generation boundary condition. The Shear Stress
Transport (SST) turbulence model [29] was employed to accurately compute the
momentum and the thermal boundary layers.

###  IV.1 Design and performance optimization

The CFD design optimization encompassed three main components: inlet and
outlet ducts of the target vessel, cradle, and anti-creep plates.

The optimization was initially focused on tuning the number of inlets and
outlets, their position, and dimensions. The final configuration presents the
best compromise between the target pressure drop and a balanced distribution
of nitrogen, while complying with the space constraints of the assembly.

The cradle was designed to evenly distribute the nitrogen flow from the inlets
to the channels between the lead slices. Large plenum volumes, as shown in
Fig. 10, induce a deceleration of the fast flowing nitrogen, while their
curved geometry inhibits flow separation at the walls due to the adverse
pressure gradient, typical of a sudden expansion. As shown in Fig. 10, the
incident beam axis forms a 10∘ angle with the target to reduce the EAR1
background caused by $\gamma$-rays and high-energy charged particles [8]. The
beam is sufficiently diverging so that a dedicated dump is not required
downstream the target in addition to the iron and marble already present
around the target area (see Sec. VII). The cradle design is complemented with
flow deflectors aligned with the beam axis to focus the cooling flow on the
hottest regions of the lead surface.

Figure 10: (a) Top view of the main core assembly (see also Fig. 5). Nitrogen
flows through the channels machined in the anti-creep plates, perpendicular to
the view and toward the observer. The beam does not impact the lead surface
perpendicularly, but at a 10∘ angle. (b) Cross-cut of the aluminum cradle
through the second row of channels, showing the plenum volume and the geometry
used to guide the flow into the channels.

The anti-creep plates drive the nitrogen over the lead surface while
containing the creep deformation of the material. The distribution of flow
rate among the channels is optimized by adding wedge-shaped obstructions in
the channels that are located farther from the beam, as schematically shown in
Fig. 11 and Fig. 12. The geometry of these obstructing wedges was carefully
designed to minimize pressure losses and flow disturbances.

Figure 11: Schematic view of the main core assembly (see also Fig. 5) showing,
highlighted in yellow, the channels with wedge-shaped obstructions. The
obstructions partially restrict the flow in the channels located farther from
the beam axis, increasing the flow rate on the regions of lead surface
directly impacted by the beam. Figure 12: Example of an anti-creep plate and
detail of the wedge-shaped obstructions. The obstructing wedges increase the
pressure drop but help deliver higher flow rate on the regions of lead surface
directly impacted by the beam.

###  IV.2 Cooling performance

The cooling performance is presented for four operational cases, summarized in
Table 2. Cases A and B reflect the two normal operational limits of the
cooling station, while cases C and D assess the cooling performance when some
of the cooling channels are obstructed by lead due to creep deformation. The
cases C and D are defined by considering no flow nor heat transfer in the
channels that are adjacent to the lead surfaces where the highest temperature
values are expected. The channels simulated as obstructed are the ones closest
to the beamline, on both faces of the 2nd slice and on the downstream face of
the 6th slice (according to the nomenclature in Fig. 11). For each of these
faces, three channels are considered obstructed in scenario C, while five
channels are considered obstructed in scenario D. Given the expected creep
deformations discussed in Sec. V, these two cases represent rather
conservative degraded scenarios.

Table 2: Normal and degraded scenarios simulated for cooling performance assessment. Case |  n∘ clogged channels |  Volume flow (Nm3/h)
---|---|---
A | 0 | $800$
B | 0 | $1000$
C | 3 | $800$
D | 5 | $800$

Considering a beam average power of 5.4 kW (2.7 kW absorbed by the lead
slices) and a beam size of 15 mm (RMS), the maximum steady-state temperature
in each slice is reported in Fig. 13, and it appears correlated to the heat
absorbed by each slice.

Figure 13: Maximum steady-state temperature in each lead slice and scenario
described in Table 2, considering an average beam power of 5.4 kW and a beam
size of 15 mm (RMS). The highest temperature is reached in the 2nd slice,
followed by the 6th slice.

The 1st and 6th slices are an exception: the 1st slice absorbs the lowest
amount of heat without being the coldest one, while the 6th slice absorbs the
highest amount of heat without being the hottest one. This is due to the fact
that they are cooled only on one of the two faces.

The maximum steady-state temperature values occur in the 2nd and 6th slices,
ranging between 85–89∘C and 83–90∘C, respectively, and depending on the
nitrogen flow. These values are likely to increase if substantial creep occurs
where lead is subjected to higher temperature and stress values, with the 2nd
slice experiencing the highest increase (25% for case D).

Within each slice, the temperature is highest close to the beam axis, as shown
in Fig. 14.

Figure 14: Steady-state temperature distribution in the lead slices, cross-
sections intersecting the beam axis. The beam axis is not orthogonal to the
lead surface, but tilted by 10∘ in the horizontal plane to reduce the EAR1
background caused by $\gamma$-rays and high-energy charged particles.

For the downstream and upstream face of the 2nd and 6th slices, respectively,
the temperature profile along the horizontal at the height of the beam axis is
shown in Fig. 15. The average velocity of the nitrogen flow in the
corresponding channels is also represented: the speed is higher in the
channels without obstructing wedges, which translates into a higher heat
transfer coefficient (HTC) where most of the beam power is absorbed.

Figure 15: Steady-state temperature profile along the horizontal at the height
of the beam axis, for the downstream and upstream face of the 2nd and 6th
slices, and average nitrogen velocity in the corresponding cooling channels.

The HTC values were estimated using the Gnielinski formulation with entrance
length correction [30], using the velocity and surface temperature values
obtained from the CFD simulations, and the fluid properties at the bulk
temperature of nitrogen. The HTC values computed analytically at beam height
exceed the ones from the CFD simulations by 10–30%, which is a reasonable
match given the complexity of the geometry.

The pressure drop estimated by simulations and the peak value of velocity in
the target for the two operational cases are summarized in Table 3.

Table 3: Nitrogen pressure drop in the target and peak velocity values for the four scenarios described in Table 2. |  A |  B |  C |  D
---|---|---|---|---
Pressure drop (kPa) | $4.7$ | $7.0$ | $4.8$ | $5.0$
Maximum velocity (m/s) | $77.2$ | $89.5$ |  |

The pressure drop in the target is expected to be between 4 and 7 kPa
depending on the flow. A detailed analytical estimation of the pressure drop
in the target vessel was calculated via the loss coefficient method using
coefficients from [31, 32, 33], and a good agreement (difference lower than
10%) was found when compared to simulations. The blockage of cooling channels
due to creep, as in case D, results in an increase of 6% in pressure drop,
which is well within the range in which the cooling station can operate.

The analysis also shows that the nitrogen velocities in the target are below
Mach 0.3, thus the flow can be considered incompressible everywhere, with the
highest velocities occurring at the bend of the vessel inlets. Inside the
target vessel, specifically in the cradle ducts, the nitrogen flows relatively
fast, decelerating as it flows into the plenums along the ducts. The analysis
results reveal that velocities up to 65 m/s are expected to occur at the
entrance of a few channels due to the apparent cross section reduction and
flow separation at the entrance of the channels. The nitrogen flowing through
the narrow sections beside the obstruction wedges can reach velocities between
70 m/s and 87 m/s for cases A and B, respectively; however, the velocities at
the beam impinging spots are below 40 m/s. The highest velocity values,
besides taking place in limited and localized areas over colder lead surfaces,
are within the acceptable threshold value to avoid erosion of 110 m/s
(according to [34]).

##  V Thermo-mechanical assessment of the lead core

The time evolution of temperature, stress, and strain in the target were
estimated by means of CFD and FEM analyses with Ansys and LS-DYNA [28].

###  V.1 Thermal analyses

The temperature dependence of every material property has been included in the
analyses. The reference sources for the thermal properties are [35, 36, 37].

The most conservative assumption that can be made regarding the beam load is
that all of the six pulses hit the target consecutively with the minimum
period of 1.2 s. The six pulses are then followed by a cool-down of 30 s,
totaling a cycle of 36 s (hereinafter referred to as supercycle). The load
boundary condition for the thermal analyses is a power generation field in the
target, resulting from the interaction between beam and target and obtained by
means of FLUKA Monte Carlo simulations.

The cooling effect of nitrogen is modeled by a convective boundary condition,
where the heat transfer coefficient (HTC) field was imported from the CFD
simulations described in Sec. IV. The average HTC is 63.8 W m-2 K-1, but the
flow is optimized to be faster at the beam impinging spot, where the HTC
reaches a value of 130 W m-2 K-1.

The 2nd lead slice crossed by the beam is subject to the highest energy
deposition and temperature. The results of thermal simulations for this slice
are shown in Fig. 16 and 17: Fig. 16 shows the temperature field at the
instant of peak temperature, while Fig. 17 shows the plot of maximum
temperature vs. time during three 36-s supercycles, once periodic steady state
is reached. The peak temperature is 135∘C, safely lower than the melting point
of lead (327∘C) but high enough to significantly decrease strength and promote
creep. Figure 17 also shows the benefits of a change of optics, if feasible: a
larger beam size (25 mm RMS) lowers the peak temperature to 96∘C.

A series of six radiation-hard thermocouples have been placed inside the
target, directly in contact with the lead slices. Although they can provide
benchmarks for the thermal simulations and feedback on the status of the
target core, they are not essential for the target operation in case they fail
prematurely.

Figure 16: Temperature map, in the 2nd lead slice crossed by the beam, at the
instant of peak temperature. The peak temperature is equal to 135∘C. Figure
17: Maximum temperature in the target Pb core during three 36-s supercycles
once periodic steady state is reached, with beam size of 15 mm and 25 mm
(RMS).

###  V.2 Stress analyses

The target is subject to mechanical stress due to two effects, both included
in the analyses: quasi-static stresses and dynamic stresses. Quasi-static
stresses are induced by heat from the absorbed beam power in the target, which
provokes temperature and thermal expansion gradients. Furthermore, each beam
pulse is absorbed in 7 ns (RMS), provoking an impulsive temperature rise and
consequent thermal expansion, which causes the propagation of stress waves and
induces dynamic stresses.

The properties used for the structural analyses are taken from [38, 39, 40].
Two constitutive models were used to model plasticity: pure isotropic
hardening (without Bauschinger effect), and pure kinematic hardening (with
Bauschinger effect) [41]. Material characterization activities were carried
out to develop a nonlinear kinematic hardening model for plasticity of pure
lead. The details of this research will be reported in a dedicated
publication.

Figures 18 and 19 show some of the results of the structural simulations
featuring the pure kinematic hardening model. Figure 18 shows the von Mises
equivalent stress map during the stress-wave propagation stage. Figure 19
shows the oscillations, due to dynamic effects, of the von Mises equivalent
stress after the sixth (and last) pulse in a supercycle. The peak value of von
Mises stress is 2.6 MPa, while pressure and maximum principal stress reach 80
MPa: the hydrostatic component of stress is dominant. The stress level is
quite high if compared to the strength of pure lead (plastic flow begins below
1 MPa); thus, a target prototype was tested under beam irradiation in the
HiRadMat facility at CERN [42]. The outcome of the beam irradiation test was
positive as no defects and internal voids were observed inside the material by
neutron tomography inspections [8]. Further details of this beam irradiation
test will be presented in a separate publication.

Figure 18: Peak von Mises equivalent stress in the target lead core during
propagation of stress waves. The stress waves start propagating radially from
the area of impact on the proton beam axis. Figure 19: Von Mises stress vs.
time in the target Pb core after the sixth (and last) pulse of a supercycle.
The oscillations are provoked by the instantaneous temperature spike due to
heat dissipation of the beam pulse.

###  V.3 Creep analyses

Pure lead is strongly affected by creep, i.e. plastic flow of material with
time, especially at the temperature reached in the n_TOF target (135∘C max.).
Creep phenomena could induce lead to flow into some of the cooling channels,
obstructing them and reducing the cooling efficiency. The creep strain rate
for each level of temperature and stress is plotted in the deformation-
mechanism maps [40]. A constitutive model able to reproduce lead creep over
all the stress and temperature ranges of interest for the n_TOF target is not
readily available in Ansys; thus, a new model was coded as User Programmable
Feature (UPF). The new model consists of a series of piecewise functions, each
composed of multiple power laws (of the kind $\dot{\epsilon}=c\,\sigma^{n}$)
with different coefficients: each piecewise function reproduces the
deformation-mechanism maps for a single temperature; the creep strain rates
for intermediate temperatures are obtained by logarithmic interpolation.

The sixth (and last) lead slice intersected by the beam is the thickest (15
cm) and is subjected to the largest creep deformations. A simulation,
featuring the new creep constitutive model, was performed to estimate the
maximum penetration, due to creep, of the lead slice into the cooling
channels. The loads considered in the analysis are the weight of the slices
and the temperature field at the instant of peak temperature. The temperature
field is conservatively assumed to be constantly applied for 2$\times$108 s
(approximately 6 years and 4 months of continuous operation, twice the target
lifetime). The resulting displacements are shown in Fig. 20: the maximum
penetration is 0.64 mm, whereas the channel depth is 3 mm.

Figure 20: Penetration, due to creep, of the 15-cm thick lead slice into the
cooling channels (view from the top, cross-section cut through the point of
maximum penetration). The gray parts are the aluminum (EN AW-6082 T6) anti-
creep plates with the cooling channels. The color map shows the maximum
projected displacements after 2$\times$108 s (approximately 6 years and 4
months of continuous operation, twice the target lifetime).

As described in Sec. IV, degraded scenarios including the possibility of
obstructed channels were considered in CFD simulations. The results of this
analysis shows that the target operation would not be compromised: the lead
slice would be subject to a maximum peak temperature of 160∘C, which is
considered acceptable for the degraded scenario under examination.

##  VI Design of the nitrogen cooling station

The nitrogen gas for the target cooling system is produced by a nitrogen
generator, recirculated in the cooling circuit, and cooled by chilled water.
The design requirements of the nitrogen cooling system are summarized in Table
4. The nitrogen flow is filtered by High Efficiency Particulate Air (HEPA) and
active carbon filters to capture particles and volatile compounds released in
the nitrogen stream. The part of the cooling system between the target vessel
and the filter casing containing the filters is enclosed in a sealed room
under dynamic confinement: an extractor ensures that the pressure in the
sealed room is lower than the surrounding environment. In case of a leak of
activated nitrogen from the filter casing, the radioactive gas is therefore
contained in the sealed room and extracted to a dedicated ventilation system
in the target tunnel area. The maximum leak rate in the unconfined part of the
circuit that has been considered acceptable after radiation protection studies
(see Sec. VII) is 5 L/h.

Table 4: Nitrogen cooling station: cooling requirements. Quantity |  Value
---|---
Cooling power | 2.6 | kW
Nominal flow rate | 780 | Nm3/h
Pressure drop | 30 | mbar
Gas supply temperature | 20 $\pm$ 2 |  ∘C
Maximum pressure | 500 | mbar

The cooling station supplies nitrogen gas to cool the target lead slices via
AISI 304L stainless steel pipes. The main elements of the cooling station are
shown in Fig. 21.

Figure 21: (a) Main components of the cooling and moderator station. The
cooling nitrogen is compressed in the blower, cooled down in the heat
exchanger, and filtered upstream and downstream the target. The station also
includes the two seperate moderator circuits for EAR1 and EAR2, which can work
independently with demineralized water or borated water. The filters are
located inside an airtight room in dynamic confinement. (b) Photo inside the
airtight room in dynamic confinement. The HEPA and active-carbon filter
capture particles and volatile compounds in the nitrogen stream.

The gas compressor (blower) is a multistage centrifugal compressor with a flow
rate of 1000 Nm3/h and a pressure rise of 285 mbar. It integrates features to
minimize gas leaks such as magnetic coupling between compressor shaft and
motor, silicone groove O-rings between casing discs, and EPDM seals at all
bolted flanges. The 29-kW heat exchanger is installed downstream of the blower
to cool the hot nitrogen gas down to 20∘C via chilled water. It is of tubular
type to minimize the pressure drop and has tubes arranged in horizontal
configuration. The nitrogen stream is filtered by a HEPA filter upstream of
the target and by a double-stage nuclear-grade HEPA plus active-carbon filter
downstream of the target. Class 3 (ISO 10648-2) gas-tight filter casings host
all of the filters. The casing design allows the filter status to be monitored
and the filters to be replaced without contact with nitrogen. The HEPA filters
capture particles and aerosols with an efficiency greater than 99.98% for a
minimum particle size of 0.15 µm. The active carbon filter captures any
volatile compound released in the nitrogen gas stream. A gas-analysis system
based on quadrupole mass spectrometry analyzes continuously the gas
composition upstream and downstream of the double-stage filter casing with the
HEPA and the active carbon filters. As previously mentioned, the filters are
installed inside a sealed room kept in dynamic confinement via an extraction
fan.

##  VII Radiation protection considerations

The radiation protection (RP) implications of the target were thoroughly
analyzed during its design phase; the outcomes of this analysis, which also
took into account the return on experience from ten years of operation of
Target #2, contributed to the optimization of the target design. The main RP
aspects addressed in the target design study are: stray radiation in
accessible areas and compliance with CERN’s radiological area classification;
critical aspects of the new cooling system; air activation in the target area
and atmospheric releases; radioactive waste considerations in view of the
target final disposal at the end of its lifetime.

Along with the target, the target shielding was also upgraded. The Target #2
shielding, composed of fixed concrete blocks with a maximum thickness of 200
cm, has been replaced with a new mobile version. The new shielding is composed
of a first layer of 40-cm thick iron followed by 80 cm of concrete and finally
a 20-cm layer of marble. The shielding wall has two main goals: to limit air
activation in the target area during beam operation and subsequent releases
into the environment (see Sec. VII.3) and to reduce the residual activation
levels in the area during access. The new mobile shielding provides direct
access to the target and enhances the potential applications of the facility
(e.g., to perform irradiation of material samples in the proximity of the
target [43]). Several openings are present in the shielding to permit the
future installation of a pneumatic system to transport radioactive samples and
to access the target with endoscopes and radiation probes. The shielding plays
a major role in ensuring adequate radiation protection and, therefore, it was
studied and optimized by means of FLUKA simulations (Fig. 22).

Figure 22: (a) 3D view of the FLUKA geometry of the new shielding. (b) 3D
cross-sectional view inside the shielding wall. (c) Photo of the shielding in
’open’ position: it can be moved on rails to access the target area.

###  VII.1 Stray radiation

As already mentioned in Sec. II.1, Target #3 provides, when compared to Target
#2, a neutron fluence towards EAR2 twice as high in the energy range 100
keV–200 MeV; the locations of the n_TOF facility that could be impacted by the
resulting stray radiation are the ones surrounding the 20-m vertical beam
line, i.e. EAR2 and ISR8 (the eighth octant of the former CERN Intersecting
Storage Ring, Fig. 23).

Figure 23: FLUKA geometry: side view of of the vertical beam line. The most
important locations impacted by stray radiation are the experimental area EAR2
and the ISR8 area.

The ISR8 area is shielded with concrete blocks and houses the cooling station
and a space for storage.

FLUKA simulations were performed to assess the ambient dose equivalent,
H*(10). Figure 24 shows the prompt H*(10) rate in the ISR8 area normalized to
the average beam intensity of 1.67$\times$1012 p+/s: the worst location in
terms of radiation exposure is inside the n_TOF storage area close to the
shielding wall. The H*(10) is, however, always lower than 0.3 µSv/h, well
below the applicable limit. In addition, the simulated H*(10) and the
integrated number of protons delivered to the target in a year were used to
estimate the exposure of personnel working in ISR8 (400 hours per year
considering ISR8 a low-occupancy area). Figure 25 shows that the annual
exposure of personnel in ISR8 is always below the CERN optimization objective
of 100 µSv. At the time of the construction of the vertical beam line in 2014,
the shielding of EAR2 was conservatively designed to take into account a
possible increase of the neutron fluence up to a factor of three. Calculations
showed that the new neutron fluence is within this threshold.

Figure 24: Prompt H*(10) rate in ISR8 for the maximum average beam intensity
of 1.67$\times$1012 p+/s. The highest radiation exposure is in the n_TOF
storage area, but still below the applicable limit. Figure 25: Annual ambient
radiation dose equivalent H*(10) in ISR8 for 400 hours of exposure and
2.5$\times$1019 protons on target. The annual exposure is always below the
limit of 100 µSv.

###  VII.2 Cooling circuit

From an RP standpoint, one of the critical aspects of the operation of Target
#2 was the direct contact between the cooling water and the lead core; due to
erosion-corrosion and out-diffusion effects, lead spallation products migrated
from the target into the cooling water. These radionuclides include not only
$\gamma$-emitters like Hg-194, Au-195, and Bi-207, but also $\alpha$-emitters
like Gd-148, Po-208, Po-209, and Po-210. The cooling circuit is equipped with
cartridges to trap some of these components and reduce the contamination in
the circuit. As a consequence, the cartridges represented radiation hot spots
and were an important source of external exposure during inspections or
maintenance activities in the cooling station; in addition, the presence of
contamination in the circuit increased the radiological risk during
intervention.

The new nitrogen cooling circuit coupled with HEPA and active carbon filters
reduces the presence of lead spallation products and increases the
radiological safety during maintenance activities. As described in Sec. VI,
the room housing the filters is under dynamic confinement and the extracted
air is directed into the target area from where it is released outside, thanks
to the preexisting ventilation and radiation monitoring system (see also Sec.
VII.3). Activation and contamination levels of the nitrogen circuit are
monitored by a dedicated beta-radiation detector.

###  VII.3 Air activation in the target area and environmental releases

The air surrounding the target area becomes activated due to high-energy
hadron interactions. Short-lived positron emitters, like 11C, 13N, 14,15O, are
produced in oxygen and nitrogen by spallation reactions. In addition, 41Ar is
generated by thermal neutron capture in the natural argon in air. These
radionuclides are the main concern when evaluating the environmental impact of
radioactive releases.

The n_TOF target area is dynamically confined from the adjacent zones; the
dynamic confinement is guaranteed by extracting air from the target area at a
rate of about 300 m3/h during beam operation, and up to 500 m3/h in case of
access. After a filtering stage and radiation monitoring, the air is released
into the environment. During the operation of Target #2, the effective dose
due to environmental releases from the n_TOF facility was, on average, about
0.6 µSv per year, well below the CERN optimization objective of 10 µSv per
year.

Following the recent modifications implemented in the target area, i.e. the
design and installation of the new target and the new mobile shielding, air
activation as well as the impact of environmental releases to the reference
population group were assessed and compared with the ones of Target #2\. The
particle spectral fluence for neutrons, charged pions, protons, and photons
were simulated by FLUKA simulation package. Then, using ActiWiz (an analytical
code developed at CERN [44]), the particle spectral fluences were folded with
production cross-sections for air to obtain the radionuclide production yields
per primary proton impinging on the target. The activity released into the
atmosphere all along the irradiation, $A_{rel,i}$, of a given radionuclide $i$
can be calculated using Eq. 1 [45]:

$A_{rel,i}=Y_{i}\ I\ \frac{\lambda_{i}}{\lambda_{i,\mathit{eff}}}\
\frac{Q}{V}\left(t_{\mathit{irr}}-\frac{1-e^{-\lambda_{i,\mathit{eff}}t_{\mathit{irr}}}}{\lambda_{i,\mathit{eff}}}\right)$
(1)

where $Y_{i}$ is the production yield for the radionuclide $i$, $\lambda_{i}$
its decay constant, $\lambda_{i,\mathit{eff}}$ its effective decay constant
(corresponding to the sum of $\lambda_{i}$ and the air changes provided by the
ventilation system per unit time, expressed in the same unit of
$\lambda_{i}$), $I$ the beam intensity, $Q$ the ventilation rate, $V$ the
target area volume (about 1200 m3), and $t_{\mathit{irr}}$ the irradiation
time (assumed to be 200 days per year). In case of operation with the new
target, the total released activity is about 3.9 TBq per year, which is about
20% lower than the one of Target #2 (i.e. 4.6 TBq per year). In both cases, as
shown in Table 5, the released activity is dominated by 41Ar (45% for Target
#3 and 62% for Target #2). For Target #3, the production of air spallation
products is about 20% higher; however, the 41Ar production decreases by more
than 70% because of a lower production of thermal neutrons.

Table 5:  Annual released activity, $A_{rel}$, for the radionuclides mostly contributing to the effective dose for radioactive emission. Nuclide |  t1/2 |  $A_{rel}$ (Bq/year) |  $\displaystyle\frac{\textrm{Targ.~{}{\\#}3}}{\textrm{Targ.~{}{\\#}2}}$
---|---|---|---
| Targ. #2 | Targ. #3
11C | 20.4 | min | 3.69 |  $\times$1011 | 4.31 |  $\times$1011 | 1.17
13N | 10.0 | min | 9.40 |  $\times$1011 | 1.15 |  $\times$1012 | 1.22
15O | 2.0 | min | 4.13 |  $\times$1011 | 4.96 |  $\times$1011 | 1.20
41Ar | 1.8 | h | 2.83 |  $\times$1012 | 1.77 |  $\times$1012 | 0.63

In conclusion, the new target and its shielding reduce the radioactive
emissions due to air activation from the n_TOF facility to the environment.

###  VII.4 Radioactive waste aspects

In view of the final target disposal, the use of aluminum over steel for the
target vessel would offer reduced activation and residual dose rate, but lead
to additional constraints that would cancel these benefits. Chemical reactions
may happen between aluminum and the concrete-based mortar used to prepare the
waste, producing hydrogen gas that may damage the package in the long-term
[46, 47]. For this reason, it was decided to limit the amount of aluminum in
the new target and to build the new vessel in stainless steel. The main cause
of the higher residual dose rate of stainless steel is its cobalt content,
with the production of 60Co from thermal neutron capture on 59Co. To reduce
this contribution, the target vessel is built using stainless steel with
cobalt content lower than 0.1%. FLUKA simulations were performed to compare
the residual dose rate after 10 years of irradiation (target lifetime) at
2.5$\times$1019 protons on target per year and 3 years of cool-down time (time
elapsed between the end of target operation and its shipping to the repository
for long term storage). Figure 26 shows that the new target induces a higher
maximum residual dose rate by a factor of 10 on the side of the proton beam
window of the vessel. These results must be taken into account to plan and
optimize the target final disposal.

Figure 26: Comparison of the residual dose rate after 10 years of irradiation
at 2.5$\times$1019 protons on target per year and 3 years of cool-down time
for Target #2 (red) and Target #3 (blue).

##  VIII Conclusions

An upgrade of the n_TOF facility took place during the CERN Long Shutdown 2
(2019–2021). A major part of the upgrade was the installation of the third-
generation neutron spallation target, along with a new nitrogen cooling
station and movable target shielding. The new target is based on pure lead
cooled by nitrogen gas, a low-cobalt stainless steel vessel, and two water
moderators for the two experimental areas of the facility. The nitrogen
cooling system replaced the water cooling system of the previous targets,
which was the cause of water corrosion and contamination with lead spallation
products. This paper presented an overview of the studies leading to the final
target design, with details on the physics performance, the mechanical design,
the thermo-mechanical aspects, the cooling station, and the radiation
protection aspects.

## Acknowledgments

The authors would like to acknowledge the support and inspiration of the n_TOF
Collaboration in order to reach the current design of this third-generation
neutron spallation target. The Project warmly recognizes the financial support
provided by the CERN Accelerators Consolidation Project and specifically by M.
Benedikt, R. Trant and M. Lamont. R. Logé wishes to thank PX Group for the
generous support to the PX Group Chair. M. Calviani acknowledges the support
and suggestions from the review panel members that helped the project advance
and deliver the final product.

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