**SPALLATION NEUTRON SOURCE Cryomodule Heat Loads and Thermal Design**

> E. F. Daly, V. Ganni, C. H. Rode, W. J. Schneider, K. M. Wilson and M.
> A. Wiseman

Thomas Jefferson National Accelerator Facility (Jefferson Lab)

> Newport News, VA, 23606, USA

**ABSTRACT**

When complete, the Spallation Neutron Source (SNS) will provide a 1 GeV,
2 MW beam for experiments. One portion of the machine's linac consists
of over 80 Superconducting Radio Frequency (SRF) 805 MHz cavities housed
in a minimum of 23 cryomodules operating at a saturation temperature of
2.1 K. Minimization of the total heat load is critical to machine
performance and for efficient operation of the system. The total heat
load of the cryomodules consists of the fixed static load and the
dynamic load, which is proportional to the cavity performance. The
helium refrigerator supports mainly the cryomodule loads and to a lesser
extent the distribution system loads. The estimated heat loads and
calculated thermal performance are discussed along with two unique
features of this design: the helium heat exchanger housed in the
cryomodule return end can and the helium gas cooled fundamental power
coupler.

**INTRODUCTION**

The cryomodules supplied to the SNS\[1\] by Jefferson Lab must
facilitate reliable 2.1 K operation of the superconducting cavities and
minimize the heat load to the refrigeration system. The design of the
SNS cryomodule (CM), described in \[2\], is largely based on the proven
CEBAF CM design \[3\] but also integrates technology from the 12 GeV
Upgraded CM as well as proven designs from machines at DESY, CERN and
KEK. There are two distinct cavity designs \[4\] housed within the
cryomodules, a β = 0.61 (medium β) cryomodule (MBCM) and a β = 0.81
(high β) cryomodule (HBCM). Where possible, common designs have been
developed for some of the CM subsystems such as end cans, helium
vessels, and fundamental power couplers in an effort to minimize
manufacturing costs by obtaining economies of scale.

Two key elements of the SNS cryogenic system design were adapted from
existing machines. The more efficient concept of providing 5 K
supercritical helium gas and then producing superfluid helium in the CM
rather than supplying 2.2 K supercritical helium from the CHL as was
done in the CEBAF machine was taken from the LHC cryogenic design. The
fundamental power coupler (FPC) design \[5\] is scaled from an existing
KEK-B design, but outer conductor requires forced-flow cooling rather
than pool boiling heat transfer to minimize static and dynamic heat
inputs to the cavity. The CM heat loads are absorbed by three helium
streams flowing within the CM (FIG. 1). The primary circuit (3 atm, 5 K)
employs a heat exchanger to pre-cool the fluid and a Joule-Thomson (J-T)
valve to produce 2.1 K superfluid helium within the CM. There is a
secondary circuit with J-T valve control that provides cooling to the
FPC outer conductors, and a third circuit (4 atm, 35 K) that cools the
radiation shield. The secondary circuit, available for use because
superfluid is produced in the CM rather than at the CHL, includes heat
intercepts and a surge tank to minimize the potential for flow
instabilities. After exiting the CM through the side of the vacuum
vessel, the exhaust is routed back to the CHL via a warm helium gas
return pipe.

The design parameters for the CM were developed in conjunction with the
design parameters for both the cryogenic system and the SC portion of
the linac and are given in TABLE 1. Specifically, the heat load budgets
were developed considering operational experience with the CEBAF
cryogenic system, where applicable. The heat load estimates and the
budget for the CM are given in TABLE 2. This paper contains a summary of
the heat loads as well as detailed descriptions of the contributions of
each subsystem.

**cavity, Helium Vessel and tuner**

The SNS helium vessel design, described in \[6\], supports the cavity
during all phases of operation, facilitates cavity tuning and contains
the cryogens with appropriate plumbing. Each titanium helium vessel, 
0.61m, contains a single niobium cavity whose six cells are immersed in
\~150 liters of He-II during normal operation. The ends of the cavities
protrude through the helium vessel heads into the insulating vacuum to
allow beamline assembly as well as attachment of FPCs, Higher Order Mode
(HOM) filters and field probes. These end groups are subject to
radiation heat transfer from the 50 K shield and the FPC, solid heat
conduction from instrumentation leads and the FPC as well as power
generated within the HOMs. Extraction of HOM signals to room temperature
is planned, requiring  3.6 mm semi-rigid coaxial cable running from the
cavity to the vacuum tank exterior. Based on finite element analysis,
the heat loads incident on the HOMs must not exceed 0.25 W in order to
maintain the HOMs well below the niobium critical temperature.
Additionally, instrument wiring for temperature diodes, liquid level
sensors, and heaters are routed from the helium vessel to the vacuum
tank exterior. The helium vessel heads possess attachment points for
Nitronic-50™ stainless steel support rods and the cavity tuner. The
calculated heat load intercepted by the cavity and helium vessel
assembly is given in TABLE 2.

The dynamic heat load for each cavity design has been calculated for
E~peak~ = 27.5 MV/m. It is desirable to increase the accelerating
gradient such that E~peak~ = 35 MV/m, thereby reducing the number of
HBCMs required to reach a given linac output energy. The dynamic heat
load increases at least quadratically with increasing peak gradients or
higher if operated in the field emission region. These increased losses
would be handled by some portion of the available margin in the primary
circuit of the refrigeration system.

**Vacuum Vessel and Space Frame**

The vacuum vessel and space frame, described in \[7\], serve to locate
the cavities accurately within the CM via the Nitronic-50™ support rods,
provide the structural links to the external supports and remain at room
temperature (300 K) during normal operation (FIG. 2). The vacuum vessel,
 0.99 m, contains the insulating vacuum and provides pressure
containment in the unlikely event of cryogenic piping failure. The
insulating vacuum is cryo-pumped during accelerator operation to less
than 1.3 x 10^-9^ atm, an adequate vacuum level to minimize residual gas
conduction. If a cryogenic piping failure occurs, two spring-loaded
parallel plate pressure reliefs, identical to those used in the CEBAF
CM, open at 1.2 atm on the vacuum vessel. In addition, there are many
flanged penetrations in the vacuum vessel: ports in the bottom of the
vacuum vessel through which the FPC provides RF power to the cavities,
instrumentation ports containing electrical feed-throughs for diodes,
cavity heaters, tuner power and limit switches, field probes, and
connections for the warm helium coupler
exhaust.![](media/image3.png){width="2.51875in"
height="2.745833333333333in"}

**Thermal Radiation Shield**

![](media/image4.png){width="2.332638888888889in"
height="1.7451388888888888in"}![](media/image5.png){width="2.4590277777777776in"
height="1.7506944444444446in"}

As in CEBAF, a copper shield 2.37 mm thick, operating between 35 and 50
K is used to intercept thermal radiation (FIG. 3). The shield is cooled
by supercritical helium gas (4 atm, 35 K) through a single pipe, 22.2 mm
outer diameter by 1 mm thick wall. Thermal analysis of both single- and
three-pass shield cooling showed \~2% increase in the total static heat
load to 2 K with a single-pass cooling scheme. Cost ultimately dictated
the choice of single pass vs. three-pass piping.

The shield is divided into segments that cover the helium vessel
assemblies and bridges that cover both the cavity-cavity and
cavity-end-can interconnect regions. The bridge sections allow access to
the cavity string for assembly, alignment, instrumentation wire routing
and tuner maintenance. There are three segments in a MBCM and four
segments in a HBCM. A hydro-formed stainless steel bellows, soldered
with EasyFlo-45™, connects the copper piping between segments. In
addition to the bellows, a series of tabs 76 mm long x 76 mm wide, where
the pipe is soldered with BCuP to the shield, provides strain relief
from differential thermal contraction of the piping and shield during
cool down and operation.

The horizontal and vertical helium vessel Nitronic-50™ support rods are
heat-stationed to the shield segments by copper straps, 1 mm x 25 mm x
125 mm long. All wiring and cabling routed to 2 K is heat stationed at
50 K on the nearest shield segment as well. Due to the long length
(\>600 mm), small cross-sectional area (20 mm^2^) and difficulty during
assembly, the two axial support rods are not heat stationed. The
additional heat load to 2 K is negligible (\< 1 mW/support).

Each bridge is fixed at only one end to an adjoining segment with PEM
fasteners and allowed to slide over the other neighboring segment. Each
segment-bridge assembly weighing \~45 kg is supported by eight G10
straps, 38 mm x 4.8 mm x 508 mm long, terminated with stainless steel
pipe clamps affixed to the space frame support tubes. The four
horizontally oriented supports react lateral shipping and handling
loads, the four vertically oriented supports react both gravity and
transportation loads and all eight react any axial loads.

**MultiLayer Insulation (MLI)**

The insulation system design is identical in principle to that employed
in the original CEBAF CM and is expected to achieve the same thermal
performance. The insulation scheme reduces the radiative heat load to
the cold surfaces, provides ample mass and heat capacity to mitigate
thermal transients and is comprised of materials suitable for use in a
high radiation environment.

Blankets are constructed of alternating layers of double-aluminized
Mylar™ (DAM), 25 m thick, and Reemay™ #2250, a spun-bonded polyester 75
μm thick. The maximum emissivity allowed by emissometer measurement is
0.035 at room temperature. The surfaces of the helium vessel are covered
with two 12-layer blankets, the beamline components and piping are
spiral-wrapped with 15 layers at a 50% overlap, and the 50 K surfaces
are covered with four 15-layer blankets. The joints and seams, staggered
both axially and around the circumference by 25 to 50 mm, are closed
with 25.4 mm wide aluminized Mylar™ tape.

Effective thermal conductance through the blanket layers is calculated
considering three mechanisms: radiative heat transfer -- governed by
emissivity (ε) and geometry, solid conduction -- governed by effective
thermal conductivity (k~eff~) and geometry, and residual gas conduction
-- governed by average local pressure. A Fortran program, TRANSAM,
written originally for transfer line design was used to estimate the
expected heat flux to the 50 K and 2 K surfaces. Using a worst-case
scenario with average pressure of 10^-4^ torr, ε = 0.2 and k~eff~ = 1.5
mW/cm·K, the heat flux to the 50 K thermal shield was calculated
conservatively as 2.5 W/m^2^. Using the same average pressure, ε = 0.06
(lowered since emissivity typically decreases with decreasing
temperature) and k~eff~ = 1.5 mW/cm·K, the heat flux to the 2 K surfaces
was calculated as 94 mW/m^2^. While these values are conservative,
proper assembly and installation of the blankets is required for ideal
performance. Taped joints, insufficient overlap and compaction of
blanket layers can reduce the effectiveness of the MLI thereby
increasing significantly the incident heat load. Combining the realities
of assembly with the fact that nine of twenty-two measured CEBAF CMs, as
reported in \[3\], exceeded the 50 K static heat load design criteria,
it was decided to increase the estimated radiation heat load to 200% of
the calculated values.

**End Cans, heat exchanger and Cryogenic Piping**

The end cans route helium to and from the CM, provide controls for the
primary and secondary circuits, provide beamline vacuum connections and
contain pressure relief valves used in the event of loss of vacuum (FIG.
4). The end cans use bayonet connections that are identical to those
used in the CEBAF CM \[8\]. In addition to the primary and shield supply
bayonet connections, two J-T valves similar to those used in CEBAF are
installed in the supply end can: the primary J-T has a C~v~ = 0.3 and
the secondary J-T has a C~v~ = 0.05. The return end can contains a
sub-atmospheric primary return bayonet, the shield return bayonet, a
Circle Seal™ pressure relief valve plumbed in parallel with a parallel
plate relief valve, a cool-down valve with a C~v~ = 3.0 and a
helium-helium heat exchanger (HX).![](media/image6.png){width="2.5in"
height="1.0486111111111112in"}![](media/image7.png){width="2.997916666666667in"
height="1.2048611111111112in"}![](media/image8.jpeg){width="2.9993055555555554in"
height="1.9534722222222223in"}

The counter flow HX (FIG. 5) is a plate-fin type core constructed of
aluminum with a pressure rating of 12 atm. Stainless-to-aluminum joints
transition from the aluminum body to the stainless steel process piping.
The design parameters are given in Table 3. Silicon diodes are installed
in the assembly to measure HX terminal temperatures in order to assess
the HX effectiveness.

The stainless piping within the CM is strain-relieved from the end can
connections with flexible metal hose. All of the shield circuit, the
fill-lines between helium vessels and the primary circuit plumbing
through the HX high-pressure side are 26.7 mm OD x 2.1 mm wall. The
primary return piping through the HX and in the relief stack is 60.3 mm
OD x 2.7 mm wall, sized for a catastrophic loss of beamline vacuum. The
helium vessel return headers are 88.9 mm OD x 3.0 mm wall, and along
with the helium vessel provide 10% ullage. The secondary circuit
consists of 6.35 mm outer diameter x 1.3 mm wall stainless steel tubing,
sized to reduce helium inventory in that circuit, and a 1 liter surge
tank which functions to damp potential flow oscillations that may arise
during operation.

**Fundamental Power Coupler**

Due to operating frequency and space constraints, a coaxial coupler with
inner and outer conductors was selected instead of a waveguide coupler
to provide fundamental RF power to the cavities. Besides delivering RF
power, the coupler must not adversely affect the electromagnetic
performance of the cavity or the thermal performance of the cryomodule.
Cavity cleanliness procedures require the coupler to be inserted into
the cavity at the six o'clock position.

To handle static and dynamic heat loads in the coupler and cavity, the
outer conductor is cooled by a nominal stream of 3 atm, 5 K
supercritical helium flowing at 0.038 g/s per coupler and the inner
conductor is conduction cooled by a 1 liter per minute water flow. The
ceramic RF window is maintained at 300 K during operation by a
heater-thermocouple control loop.

The outer conductor is a machined and welded stainless steel assembly
with flanged connections to the cavity and the warm window. The internal
vacuum surface is copper-plated with a nominal thickness of 15 μm and an
estimated RRR = 10 to reduce resistive wall losses induced by the RF
surface currents. A single helical, square-grooved flow passage, 2.3 mm
wide x 1.6 mm deep (FIG. 6), with a pitch of 2.6 turns/cm is machined
into a thick-walled stainless steel tube, 75.75 mm internal diameter. A
thin-walled stainless steel tube (1.6 mm O.D.) is then shrunk-fit over
the outer diameter of the flow passages and welded leak tight onto the
ends nearest the flanges. Due to the small hydraulic diameter of
2.16 mm, flow velocities and Reynolds number are kept high to mitigate
potential deleterious effects of buoyancy, re-circulation and poor heat
transfer in the helium stream.

Conduction heat load estimates (TABLE 2) were calculated for the outer
conductor using an Excel model that included combined conduction and
convection heat transfer as well as temperature-dependent thermal
properties for stainless, copper and supercritical helium gas (FIG. 6).
A separate estimate of the static radiation heat load emitted from the
FPC to the beam pipe was calculated. The values were used as input to a
finite element model of the cavity ends to verify that they remain below
the niobium critical temperature.

In an effort to determine if thermo-acoustic oscillations or flow
instabilities exist in the secondary circuit, a test rig has been
designed and built. Experiments to evaluate the flow circuit are planned
for this summer.

**CONCLUSION**![](media/image9.jpeg){width="1.5in"
height="1.7868055555555555in"}![](media/image10.jpeg){width="1.5in"
height="1.5472222222222223in"}![](media/image11.wmf){width="2.7284722222222224in"
height="2.245138888888889in"}

The estimated heat loads and calculated thermal performance of the SNS
medium and high β cryomodules do not exceed the specified budget and are
consistent with the refrigeration system design capacity.

**ACKNOWLEDGEMENTS**

This work is supported by US DOE under Contract No. DE-AC05-00OR22725.
Thanks to the Jefferson Lab Cryomodule design group and the SNS cavity
prototyping team for their efforts.

**REFERENCES**

1.  T. E. Mason, "The Spallation Neutron Source: A Powerful Tool for
    Materials Research", PAC2001, Chicago, IL, June 2001.

2.  W. J. Schneider et al., "Design of the Spallation Neutron Source
    (SNS) Cryomodule", PAC2001, Chicago, IL, June 2001.

3.  W. Schneider et al., "Thermal Performance of the CEBAF
    Superconducting Linac Cryomodule" in *Advances in Cryogenic
    Engineering*, Vol. 39, Plenum Press, New York (1994), p 589-597.

4.  G. Ciovati et al., "Superconducting Prototype Cavities for the
    Spallation Neutron Source (SNS) Project", PAC2001, Chicago, IL, June
    2001.

5.  I. E. Campisi et al., "The Fundamental Power Coupler Prototype for
    the Spallation Neutron Source (SNS) Superconducting Cavities",
    PAC2001, Chicago, IL, June 2001.

6.  J. Hogan et al., "Design of the SNS Cavity Support Structure",
    PAC2001, Chicago, IL, June 2001.

7.  T. Whitlatch et al., "Shipping and Alignment for the SNS
    Cryomodule", PAC2001, Chicago, IL, June 2001.

8.  D. Kashy et al., "CEBAF Transfer Line Systems" in *Advances in
    Cryogenic Engineering*, Vol. 37, Plenum Press, New York (1992), p
    577-586.