NASA/TP—2012–216315 

 

Two Dimensional Heat Transfer around 

Penetrations in Multilayer Insulation 

Final Report 

Wesley L. Johnson and Andrew O. Kelly 

Fluids Test and Technology Development Branch 

Kennedy Space Center, Florida

Kevin M. Jumper 

Sierra Lobo, Kennedy Space Center, FL 

 

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October 2012
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NASA/TP—2012–216315 

 

Two Dimensional Heat Transfer around 

Penetrations in Multilayer Insulation 

Final Report 

Wesley L. Johnson and Andrew O. Kelly 

Fluids Test and Technology Development Branch 

Kennedy Space Center, Florida 

Kevin M. Jumper 

Sierra Lobo, Kennedy Space Center, FL 

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National Aeronautics and 

Space Administration 

Kennedy Space Center 

Kennedy Space Center, FL 32899-0001 

October 2012
Acknowledgments 

 Enter acknowledgments here, if applicable. 

The authors would like to acknowledge Wayne Heckle and Johnny Kerce for performing a 

majority of the testing in this report and James Fesmire for his support and guidance in boil-off 

calorimetery. 

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443-757-5802 
1

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Table of Contents

1 Background...........................................................................................................................................4

2 Introduction...........................................................................................................................................6

3 Experimental.........................................................................................................................................8

3.1 Instrumentation ........................................................................................................................... 10

3.2 MLI Blankets.............................................................................................................................. 13

3.3 Penetrations.................................................................................................................................15

3.4 Buffer Materials.......................................................................................................................... 18

3.5 Test Matrix..................................................................................................................................19

3.6 Experimental Uncertainty ........................................................................................................... 21

3.7 System Data ................................................................................................................................22

4 Thermal Models..................................................................................................................................26

4.1 Model Overview ......................................................................................................................... 26

4.1.1 Cryostat Detailed Model Overview .................................................................................... 26

4.1.2 Scaling Model Overview .................................................................................................... 28

4.1.3 Cryostat Detailed Model Validation ................................................................................... 29

4.1.4 Detailed Model Validation Discussion ............................................................................... 33

4.1.5 Scaling Model Results ........................................................................................................ 34

5 Test Results and Analysis................................................................................................................... 39

5.1 No Penetration – Null testing...................................................................................................... 40

5.2 No Integration ............................................................................................................................. 42

5.3 Buffer Materials.......................................................................................................................... 42

5.4 Temperature Matching................................................................................................................ 47

6 Conclusion ..........................................................................................................................................50

References...................................................................................................................................................52

Appendix A – More data plots.................................................................................................................... 53

Appendix B – Sample Calculation.............................................................................................................. 58

Appendix C – Uncertainty Analysis Derivations........................................................................................ 60
2

List of Figures

Figure 1: Schematic of Insulation Penetration Cryostat, Reference 1: Figure 18......................................... 4

Figure 2: Installation of fiberglass strut, Reference 2: Figure 6. .................................................................. 5

Figure 3: Shield-penetration grid and shield-buffer-penetration grid, Reference 3: Figures 2a and 3a........5

Figure 4: Cryostat-600 cold mass during liquid nitrogen cold shock ........................................................... 8

Figure 5: Passive heater attached to the half inch diameter strut.................................................................. 9

Figure 6: Schematic & controls diagram for Cryostat-600......................................................................... 10

Figure 7: Temperature sensor locations with no strut................................................................................. 12

Figure 8: Temperature sensor location with strut present (assumes a 0.5" thick buffer)............................13

Figure 9: Thermocouples in a semi-circular pattern ................................................................................... 14

Figure 10: Punch used to insert holes into MLI blankets............................................................................ 14

Figure 11: G-10 MLI support shown holding up MLI underneath Cryostat-600 .......................................15

Figure 12: Threaded taps on the bottom of Crysotat-600 ........................................................................... 16

Figure 13: Composite strut for penetration testing. .................................................................................... 17

Figure 14: Aerogel bead package................................................................................................................ 19

Figure 15: General Test Matrix for Penetrations Calorimetry .................................................................... 20

Figure 16: Temperature profile of MLI and penetrations through the thickness of the MLI......................21

Figure 17: Typical mass flow rates and fill level data ................................................................................ 23

Figure 18: Typical system temperatures and preferences........................................................................... 24

Figure 19: Typical MLI temperatures......................................................................................................... 25

Figure 20: Steady state MLI temperatures plotted vs. location .................................................................. 25

Figure 21: Section View of Detailed Model .............................................................................................. 28

Figure 22: Predicted Temperature Accross Cold Mass.............................................................................. 33

Figure 23: Change in Heat Leak with Penetration Diameter..................................................................... 35

Figure 24: Change in Heat Leak with Buffer Thickness............................................................................ 37

Figure 25: Change in Heat Leak with number of MLI layers.................................................................... 38

Figure 26: Test Result Summary for Penetration Calorimetry ................................................................... 39

Figure 27: Temperature Profile for P100.................................................................................................... 41

Figure 28: Temperature Profile for P118.................................................................................................... 41

Figure 29: Temperature profiles for No Integration testing on Aluminum strut ........................................ 42

Figure 30: Aerogel blanket buffer - 0.5" (0.25" radius).............................................................................. 43

Figure 31: Aerogel beads packed in nylon and installed on calorimeter .................................................... 44

Figure 32: 1" (0.5" radius) Cryo-Lite buffer............................................................................................... 44

Figure 33: P115, 1 inch buffer, 0.25" strut 2-D thermal projection............................................................ 45

Figure 34: Composite strut installed on calorimeter with a Cryo-Lite buffer............................................. 47

Figure 35: Predicted MLI and penetration temperature gradients.............................................................. 48

Figure 36: Temperature gradients through MLI showing predicted, P100, and P112. Also shown for 

reference are the strut temperatures. ........................................................................................................... 49

Figure 37: Comparison between different integration methods for a half inch aluminum penetration ......50
3

List of Tables

Table 1: Key Performance Parameters for Penetrations Testing .................................................................. 7

Table 2: Pressure transducers and their ranges........................................................................................... 11

Table 3: Description of the four struts tested.............................................................................................. 17

Table 4: Buffer material descriptions.......................................................................................................... 19

Table 5: Uncertainty Analysis Results....................................................................................................... 22

Table 6: Material properties used in the calorimeter thermal model .......................................................... 26

Table 7: Typical Cold Mass Temperatures................................................................................................ 27

Table 8: Penetration Heat Load Predicted From Model Compared With Test Result Calculations..........30

Table 9: Penetration Heat Load - Test vs Model ....................................................................................... 32

Table 10: Test P115 Strut Temperature Data............................................................................................. 32

Table 11: No penetration MLI test results.................................................................................................. 40

Table 12: Results from buffer material testing at different buffer thicknesses........................................... 43

Table 13: Test Results for Cryo-Lite Buffers on various penetration sizes................................................ 46

Table 14: Temperature matching locations along the strut (total length .203 m).......................................48
4

1 Background 

Cryogenic multilayer insulation (MLI) has been studied thoroughly over the last 50 years. 

Numerous tank and calorimeter tests have been performed using many different insulation 

approaches. Many different variables have been tested and documented, mainly within the 

insulation system itself. There are several factors in insulation application that can drive up the 

heat load on the entire system. These include the treatment of insulation seams, 

instrumentation wires running through the insulation, and the integration of the insulation with 

the structural and fluids. 

Several attempts have been made to identify the performance losses due to structural 

integration with a real system. (1-3) In Reference 1, Coston, Murray, and Lambert investigated 

the losses due to penetrations using the device shown in Figure 1. They investigated using an 

intermediary buffer made of Dexiglass (which at the time was often used as the spacer in MLI) 

of three and five inches. In cooperation with thermal models, the heat load was found to 

minimize at a point much less than the three inches thick buffer. They could not measure a 

meaningful added heat load through the insulation except for in the five inch thick Dexiglass 

buffer. In Reference 2, Sumner installed a fiberglass strut onto a liquid hydrogen calorimeter as 

shown in Figure 2. No buffer or intermediary was installed around the strut and the degradation 

area was half of a meter with an excess heat load of just over 0.4 W due to the integration. In 

Reference 3, Johnson and Sprague built a nodal based thermal model which they used to 

explore the relationship between the strut, MLI, and a buffer (see Figure 3). They used 

aluminum foil for the reflection layers and a general isotropic material with thermal conductivity 

of 16 mW/m-K as a buffer material. The end result of their modeling provided equations for 

estimating the net excess heat load into the tank as well as the degradation area of the MLI for 

each configuration. 

Figure 1: Schematic of Insulation Penetration Cryostat, Reference 1: Figure 18
5

Figure 2: Installation of fiberglass strut, Reference 2: Figure 6.

Figure 3: Shield-penetration grid and shield-buffer-penetration grid, Reference 3: Figures 2a and 3a.

Due to the nature of MLI, these were tied to specific programs and configuration dependent. In 

a review of the state of the art for multilayer insulation applications, it was determined that 

uncertainty related to the integration of penetrations through the MLI was a risk to 

demonstrating a high performance system in low earth orbit. 
6

2 Introduction 

The objective of this task was to quantify thermal losses involving integrating MLI into real life 

situations. Testing specifically focused on the effects of penetrations (including structural 

attachments, electrical conduit/feedthroughs, and fluid lines) through MLI. While there have 

been attempts at quantifying these losses both analytically and experimentally, none have 

included a thorough investigation of the methods and materials that could be used in such 

applications. 

To attempt to quantify the excess heat load coming into the system due to the integration 

losses, a calorimeter was designed to study two dimensional heat transfer through penetrated 

MLI. The test matrix was designed to take as many variables into account as was possible with 

the limited test duration and system size. The parameters varied were the attachment 

mechanism, the buffer material (for buffer attachment mechanisms only), the thickness of the 

buffer, and the penetration material. 

The work done under this task is an attempt to measure the parasitic heat loads and affected 

insulation areas produced by system integration, to model the parasitic loads, and from the 

model produce engineering equations to allow for the determination of parasitic heat loads in 

future applications. The methods of integration investigated were no integration, using a buffer 

to thermally isolate the strut from the MLI, and temperature matching the MLI on the strut. 

Several materials were investigated as a buffer material including aerogel blankets, aerogel 

bead packages, cryolite, and even an evacuated vacuum space (in essence a no buffer 

condition). 

A very specific method was used to quantify the integration dependent heat loads. First the MLI 

blanket was tested with no penetrations to provide a baseline heat load case. The test 

penetrations were made out of known materials, either aluminum or carbon fiber composites, 

with known geometries to allow for the heat load down the penetration to be easily calculated 

using Fourier’s Law and two temperatures measured along the penetration: 

The parasitic integration heat load could then be calculated by subtracting the MLI and 

penetration heat loads from the measured heat loads. Ideally, if there were no integration 

losses, this would be zero. 

To measure the affected insulation zone, temperature sensors were placed within the MLI 

blanket on layers 3, 8, 16, and 25 at predetermined spacing. Thus by starting at the center of 

the blanket and working out radially, temperature gradients could be determined and if they 

were within the uncertainty of the thermocouple (roughly 1 K), then it was assumed that the 

affected area had ended. All degradation radii are measured from the center of the penetration 

and are determined to be where the temperature changed by 1 K or less. 
7

The Key Performance Parameters for the testing are shown in Table 1. 

Table 1: Key Performance Parameters for Penetrations Testing

KPP Minimum 

Success 

Full Success 

Integration Heat Load (W) < 0.2 < 0.1 

Temperature affected zone (m) 0.05 0.02 

Two different types of thermal models have been developed for the system. The first is a 

Thermal Desktop model, where the whole calorimeter system was modeled. The actual 

materials used in building the calorimeter along with their appropriate mechanical, thermal, and 

optical properties were input into the Thermal Desktop model. The Thermal Desktop model was 

used to better understand the testing and how to interpret the testing. Additionally, once 

validated with test data, the model was scaled to allow for extrapolation on several parameters. 
8

3 Experimental 

To perform the desired testing, a guarded flat plate calorimeter was designed and fabricated for 

this special application. The outer diameter of the guard chamber is 12.000 inches and the 

height is 7.50 inches. The test chamber is 8.230 inches in diameter and 2.50 inches tall. Four 

threaded taps were put into the bottom surface of the calorimeter to allow for the attachment of 

the various penetrations. Both the test and guard chambers were leak checked to 1 x 10-8 sccs 

following cold shock with liquid nitrogen (see Figure 4). 

Figure 4: Cryostat-600 cold mass during liquid nitrogen cold shock 

After cold shock, the cold mass instrumentation was then placed on the vessel to measure the 

temperature on the bottom of the cold mass, on the side a few inches from the top and on the 

top of the cold mass. A half inch versify strip was attached to the cold mass to provide an edge 

guard for the test samples. Then the top and sides of the guard chamber were wrapped with 10 

layers of multilayer insulation. 

Heaters were wrapped around the vacuum chamber and installed on the outer surface of a 

cylinder made from aluminum sheet which was placed inside of the vacuum chamber. The 

inside of the cylinder was painted with Aeroglaze Z306 Flat Black Absorptive Polyurethane (4) 

to provide a black environment to simulate space and to allow for baking out of samples prior to 

testing. In addition to heaters internal and external to the vacuum chamber, a “passive heater” 

was hung on the end of the penetration to maintain the warm boundary on the end of the 

penetration and extend the absorptive view to the bottom of the test area (see Figure 5). The 

passive heater was also painted with the Aeroglaze Z306. Using an A-Z Technology TESA 
9

2000 reflectometer, the solar absorptivity of the Aeroglaze Z306 was measured at 0.945 and the 

infrared emissivity was measure at 0.89. 

At the beginning of each test, the back pressure control system that is connected to both 

Cryostat-100 and Cryostat-600 was set to control the pressure at 3 Torr above the current local 

ambient pressure at +/- 0.1 Torr. This helped to dampen oscillations in the liquid nitrogen heat 

of vaporization due to changes in local atmospheric pressure. During testing, the guard and test 

chambers are periodically topped off to ensure that neither one runs out of fluid. 

Figure 5: Passive heater attached to the half inch diameter strut
10

3.1 Instrumentation 

Instrumentation on Cryostat-600 includes temperature sensors, pressure transducers, and flow 

meters. The pressure transducers and flow meters are shown in the schematic below (see 

Figure 6). MKS 10MB series of mass flow meters were used for all testing. These flow meters 

use thermal anemometry to measure the mass flow rate of the gas through an orifice. The 

Cryogenics Test Laboratory has approximately 40 flow meters, all of which are calibrated using 

in-house NASA standards and are accurate to +/- 2% or less over the range of the flow meter. 

Figure 6: Schematic & controls diagram for Cryostat-600 

The pressure transducers cover the whole range of vacuum pressure between atmospheric 

pressure and 1 x 10-7 Torr. Additionally, a pressure transducer is placed on the back pressure 

control system to allow for control of the system. The pressure control systems consist of the 

MKS Baratron capacitance manometers (and readout), a switching box to pick a signal from the 

readout to send to the control module, and a MKS 250E pressure/flow control module that 

regulates an MKS 248A flow control valve. The desired pressure setting (by means of voltage) 

is set on the user panel of the MKS 250E control module, which contains proportional, integral, 

and derivate based control logic. The details for each pressure transducer are shown in Table 

2. During checkout testing, the Granville-Phillips Micro-Ion gauge was not turned off when the 

vacuum chamber was repressurized. This caused the gauge to overheat and fail. The 

remainder of the testing was done with the Pfeiffer PBR 260 gauge substituting for the Micro-Ion 

gauge below 1x10-5 torr. 
11

To measure the temperatures within the test apparatus, 21 type-E 36 gauge thermocouples 

were attached in various locations. Prior to installation, the thermocouples were submersed in a 

liquid nitrogen bath and then an ice water bath and matched within a few tenths of a Kelvin. 

Three of the thermocouples were on the cold mass: one on the top of the guard (T1), one on the 

side of the guard chamber approximately 12.7mm from the top (T2), and one on the bottom of 

the test chamber (T3), roughly one inch from the center (off center due to strut presence). Five 

thermocouples were reserved for the penetration (T4-7, 13). Two thermocouples were reserved 

for the heater (T12) and heater controller (T12H). The remaining 15 thermocouples were 

interspersed within the MLI on layers 3, 8, 16, and 25. Figure 7 and Figure 8 show the 

temperature sensor placement within the MLI blanket for the no-penetration and penetration 

cases. As the penetration and buffer grew larger, the displaced thermocouples were generally 

set to the edge of the hole in the MLI and then 10 mm outside of that thermocouple. 

Table 2: Pressure transducers and their ranges 

Instrument Tag Description Range 

P1 Pfeiffer PBR 260 10-10 – 103 Torr 

P2 Pfeiffer CMR 361 10-2 – 103 Torr 

P3 MKS Baratron capacitance manometer 10-4 – 10-1 Torr 

P4 MKS Baratron capacitance manometer 0.1 – 100 Torr 

P5 Granville-Phillips Micro-Ion gauge 10-9 – 10-2 Torr 

P6 MKS Baratron capacitance manometer 1 – 1000 Torr 
12

Figure 7: Temperature sensor locations with no strut 
13

Figure 8: Temperature sensor location with strut present (assumes a 0.5" thick buffer)

3.2 MLI Blankets 

A total of 8 MLI blankets were cut out for testing. The first four were cut on August 15, 2011 and 

the second four were cut out on September 30, 2011. All the blankets were 25 layers (with one 

layer consists of both a reflector and spacer material). The reflector material is Dunmore DE 

028 dated from 9/09 (0.25 mil unperforated double aluminized mylar). The emissivity of the 

mylar was measured to be less than 0.03. The spacer material is Dacron B4A netting acquired 

from Aerospace Fabrication and Materials. 

All thermocouples were routed in a semi-circular pattern around the MLI blanket as shown in 

Figure 9. This lengthened the thermocouple leads considerable and lowered the thermal 

conductance of those leads into the blanket. Overall, the thermocouples added on the order of 

5 mW to the total heat load of the blanket, well under the uncertainty of the entire system. 
14

Figure 9: Thermocouples in a semi-circular pattern 

In order to put the penetration through the MLI, holes were punched in the MLI. A standard 

tooling punch (see Figure 10) was used that has punches from 1/8th of an inch all the way to 2” 

at approximately 1/16” intervals. Generally, the punched hole was sized to be 1/16th inch larger 

on the diameter than the strut plus the buffer size. From best practices, thermocouples were 

always moved prior to punching the hole and replaced if necessary after the hole was punched. 

This prevented the accidental destruction of a thermocouple bead. 

Figure 10: Punch used to insert holes into MLI blankets 
15

A G-10 ring with Kevlar string cross thread was used to hold the MLI into place underneath the 

cold mass (see Figure 11). The G-10 ring was held in place one half of an inch below the cold 

mass with help from the Versify edge guard. 

Figure 11: G-10 MLI support shown holding up MLI underneath Cryostat-600

3.3 Penetrations 

Four different penetrations were built. Three were Aluminum 6061-T6 tubes (0.25” OD by 

0.035” wall, 0.5” x 0.032”, and 1.0” x 0.049”) made from standard tubes with a welded end cap. 

The fourth was a composite strut made from two plies of carbon fiber tube (1.060” OD by 0.032” 

wall thickness) and Araldite LY8604/Aradur Epoxy Resin of which the fabrication process is 

further described below. An overview of the penetrations is given in Table 3. In order to 

minimize internal heat transfer, the tubes were filled with glass bubbles and capped with Lydall 

Cryo-Lite disks to prevent the bubbles from falling out. The tubes were attached to the 

calorimeter via threaded holes in the bottom of the calorimeter as shown in Figure 12. The ends 

of the penetrations were also lightly coated with Apeizon N type cryogenic thermal grease. 

The Fesmire number is a number that relates the thermal and structural properties of a material. 

It is defined by the ratio of the compressive strength to the mass times the thermal conductance 

of the material. The Fesmire number for a material is defined as: 

For a specific part or application, it can be specialized as: 

ൌ ݁ܨ

݈݅݁݀ݕߪ

ݐߣߩ
16

Where: 

Fe – Fesmire Number 

Fes – Special Fesmire Number 

m – mass (kg) 

kt – thermal conductance (W/m) 

tthermal conductivity (W/m/K)

 ‐ density (kg/m3

)

yield – tensile yield strength (MPa) 

Fyield- force at which the part yields (kN) 

Figure 12: Threaded taps on the bottom of Crysotat-600 
17

Table 3: Description of the four struts tested 

Strut 

Name 

Strut Size 

(OD x 

wall 

thickness

) 

Strut Material Strut 

Mass 

(g) 

Bucklin

g Load 

(kN) 

Conductivit

y (W/K) 

Special 

Fesmire 

Number 

0.25” 

strut 

0.25” x 

.035” 

6061-T6 Aluminum 10.2 108 4.5e-3 2.4e6 

0.5” strut 0.5” x 

0.032” 

6061-T6 Aluminum 27.1 89.3 2.2e-2 1.5e5 

1” strut 1.0” x 

0.049” 

6061-T6 Aluminum 77.6 209 7.0e-2 3.9e4 

Composit

e Strut 

1.060” x 

0.032” 

Carbon 

sleeve/Araldite 

LY8604/Aradur Epoxy 

20.4 10.4 7.5e-4 6.8e5 

Figure 13: Composite strut for penetration testing.

The composite strut was made using braided carbon biaxial sleeve and the Araldite LY 

8604/Aradur Epoxy Infusion System. A 1” round aluminum mandrel was prepared first by 

cleaning with Zyvax Fresh Start, treating the tool surface with Zyvax PreFlight, then applying the 

release agent Zyvax Departure to the tool surface. Two layers of carbon sleeve were then fitted 

and taped to the tool surface, one layer over the other. The carbon was then wrapped with a 

perforated release barrier with a layer of 1K 5 harness satin fiberglass on top of that as a 

bleeder material. On each end of the fiberglass bleeder a full wrap of spiral tubing was wound 

around the tool with each wrap of spiral tubing having its own flow tube attached. The tool was 

then bagged using a vacuum bagging film and sealant tape, allowing the flow tubes to pass 

through the sealant tape seals. Vacuum was pulled through one flow tube while the other tube 

was crimped closed. After checking for leaks the crimped tube was opened in a reservoir of 

epoxy. The epoxy was allowed to flow across the part then both feed tubes were closed. The 

strut was allowed to fully cure over 24 hours and then was removed from the bagging materials 

and mandrel and was trimmed to size. 

The composite plug was made from approximately 80 1.5” x 1.5” plies of Hexcel 282 3K plain 

weave carbon fabric and the Araldite LY 8604/Aradur Epoxy Infusion System. The plies were 

pierced through their centers by and stacked on a perforated metal infusion tube .31” in 

diameter. End caps were then fitted onto the metal infusion tube’s ends to compress the plies 

to approximately .75” thick. A flow tube was installed in the end of the aluminum tube for the 

infusion resin. The whole setup is then wrapped in a layer of 1K 5 harness satin fiber glass as a 
18

bleeder with spiral tubing on the inside of this bleeder. The spiral tubing is formed in a ring 

approximately 2” larger in diameter than the setup containing the carbon plies and is situated 

concentrically with the metal infusion tube, halfway between the end caps. A flow tube is also 

attached to the spiral tube. This final setup is bagged in vacuum film and sealant tape with the 

two flow lines passing through the sealant tape seals. The flow line attached to the metal 

infusion tube was crimped while vacuum was pulled through the other. After checking for leaks 

the crimped tube was opened in a reservoir of epoxy. The epoxy was allowed to flow until it 

filled the spiral tubing attached to the vacuum line. The plug was allowed to fully cure over 24 

hours and then was removed from the bagging materials and mandrel and was trimmed to size. 

To create the mounting hole in the plug, the hole left in the center of the plug from the metal 

infusion tube was used as a pilot hole for installing a UNC 5/16-18 stainless steel helicoil. 

To bond the plug in the strut a fixture was used to hold the plug’s mounting hole concentric to 

the strut. Hysol EA 9394 adhesive was used for the bonding. In trimming the plug, care was 

taken to insure the outer diameter of the plug was approximately .020” smaller than the inner 

diameter of the strut to ensure an adhesive bond line thickness of .010.” After sanding the 

mating surfaces with 120 grit abrasive and cleaning, the mating surfaces were generously 

coated with the adhesive and the plug was slowly inserted while rotating back and forth. The 

strut was then put under vacuum to help remove air bubbles from the bond line. Afterward, the 

joint was packed with more adhesive and left to cure for 24 hours in the fixture mentioned 

above. When installing the strut into the fixture, the plug was pressed inward slightly so its outer 

surface was approximately .030” below the edge of the strut. After curing the plug end of the 

strut was sanded flush with the plug and any porosity in the bond line from packing after 

vacuuming was filled with more adhesive and left to cure for another 24 hours. After the final 

cure, excess adhesive was sanded away and the strut was trimmed to 8” in length. 

3.4 Buffer Materials 

Three different types of buffer materials were tested. Aerogel was used in two separate forms: 

blanket (5) and a loose fill bead pack (6), along with Cryo-Lite (7). Detail about each material is 

given in Table 4. The aerogel blanket material is a fiberglass batting that is fully embedded with 

aerogel material. The blanket is made via a proprietary supercritical drying process of the 

material that leaves a nanoporous aerogel based material. Since the aerogel blanket is 

encased in the fiberglass batting, no special preparation was required to install it as a buffer. 

The aerogel beads are made via a similar process to the aerogel blanket, however, without the 

batting. Since the beads are a loose fill material, they must be contained to act as a buffer. For 

this testing, a thin nylon mesh was used to contain the beads and was tied off at each end. 

Figure 14 shows the bead package just prior to being installed in the calorimeter. Cryo-Lite is a 

special fiberglass based material designed specifically for cryogenic tanks. Testing at the 

Cryogenics Test Laboratory has shown the parallel direction thermal conductivity to be 3.25 

mW/m/K, thus the fiberglass is slightly anisotropic (i.e. the fiber direction conductivity is 70% 

higher than the normal direction). 
19

Table 4: Buffer material descriptions

Buffer Material Trade Name Manufacturer Avg Thermal 

Conductivity 

(mW/m/K) 

Aerogel blanket Cryogel Z Aspen Aerogels 1.55 

Aerogel bead 

pack 

Nanogel 

granules 

Cabot 1.70 

Cryo-Lite Cryo-Lite Lydall/Manning 1.95 

Note: Average thermal conductivity based on KSC testing between 293 K and 77 

K boundary conditions using Cryostat-100, and was measured in the normal 

direction. 

Figure 14: Aerogel bead package 

3.5 Test Matrix 

The test plan was developed to investigate several of the major variables associated with MLI 

and the integration with various penetrations. The general test matrix is given in Figure 15. The 

worst case scenario was assumed to be the no integration case, where a hole is left in the MLI 

for the penetration, but no attempt is made to prevent conduction from the penetration to the 

MLI. Both buffers and temperature matching were tested to determine the flexibility of the two 

methods (i.e. different environmental temperatures) in multiple applications. 

Multiple materials were tested as a buffer: aerogel blankets, aerogel beads, and Cryo-Lite 

fiberglass blankets. The aerogel blankets and Cryo-Lite fiberglass blankets were tested at two 

different diameters, 0.5” (0.25” on the radius) and 1.0” (0.5” on the radius). The aerogel beads 

were install in a nylon sleeve and tested at a thickness of approximately 0.5” (0.25” on the 

radius). 
20

Figure 15: General Test Matrix for Penetrations Calorimetry 

To determine the temperature matching, two different scenarios were used. The first was to use 

the Lockheed Equation for double aluminized Mylar and silk net (treated) to determine the 

temperature gradient through the MLI (8). The National Institute of Standards and Technology’s 

Cryogenic Material Property database was used to determine the temperature gradient through 

penetrations of aluminum and fiberglass epoxy (G-10, normal) using a 20 node model (9). 

The location where layers 1, 3, 8, 16, and 25 should meet the penetration was determined 

based on these calculations. Also, in the first few tests, the actual temperature profiles of the 

MLI and the penetration were measured. These measurements were used to calculate the 

location where layers 3, 8, 16, and 25 should meet the penetration for the second “Test #1” test. 

Layers 3, 8, and 16 were chosen because they are approximately equidistant along the 

temperature profile (~50 K apart) as can be seen in Figure 16. 
21

Figure 16: Temperature profile of MLI and penetrations through the thickness of the MLI 

3.6 Experimental Uncertainty 

The uncertainty was calculated using root sum squared method using the allowable calibration 

and other instrumentation uncertainty and equations for heat transfer and thermodynamics. The 

analysis assumes the measurement is representative of the heat transfer from the test article. 

That means 1) that the cryostat’s design eliminates heat transfer to the test chamber from 

sources other than the penetration and test section of MLI, and 2) that nothing but heat input is 

affecting the boiloff rate of nitrogen such as impurities, changes in atmospheric pressure, or 

other factors that are unknown. Because the system is dependent on nonlinear formulas, the 

uncertainty changes from test to test. Table 5 shows the uncertainty for a representative test. 

The full uncertainty analysis is contained Appendix C. 
22

Table 5: Uncertainty Analysis Results

Test Test P101 

MLI Serial Number SN 002 

Description 1/2" strut, no buffer 

Related 

Parameter Units 

Relative 

Uncertainty 

Ux/x 

Nominal 

Value 

Thermal Conductivity k W/m-K 0.005 138.3 

Volumetric Flow Rate Vmeas sccm 0.05 744.1 

Volumetric Flow Rate VMLI sccm 0.05 31 

Density ρ Kg/m3 0.0223 1.167 

Heat of Vaporization hfg kJ/kg 0.02 199.2 

Length of Heat Transfer x m 0.0019 0.18603 

Area of Strut A m2 2.75E-04 3.04E-05 

Temperature Difference ∆T K 0.0261 108.5 

Total Heat Leak - W - 3.079 

MLI Heat Leak - W Test P100 0.13 

Strut Heat Leak - W - 2.45 

Total Uncertainty 0.0101 

Percent Uncertainty 2.022% 

3.7 System Data 

This section gives a broad overview of all the data collected on each test in order to allow for a 

better understanding of the more brief data presented later as well as the operations of the 

testing. 

One of the main measurements taken is the boil-off (or mass) flow coming out of each chamber 

of the cold mass. Figure 17 shows the flow meter data for both the guard and test chambers for 

a typical test. Additionally, the liquid fill level is shown. The liquid level is calculated by 

integrating the flow over time and is solely for the purpose of knowing when the chambers need 

to be refilled. For Cryostat-600, the guard chamber needs to be kept at roughly 20% full 

minimum to ensure that the test chamber is always submerged in liquid nitrogen. The flows 

were generally very steady and had a slight variance with fill level. In order to calculate the heat 

load, the test flow rate is averaged over a long period of time (typically at least several hours, 

often approaching a day) and is multiplied by the density of the vapor at standard temperature 

and pressure and the heat of vaporization of the liquid at the saturation pressure. 
23

Figure 17: Typical mass flow rates and fill level data 

The other main data measured is the temperature data and vacuum chamber pressure. Figure 

18 shows the system temperatures on the cold mass and penetration as well as the vacuum 

chamber pressure. T1 – T3 are on the cold mass as shown in Figure 7. At about the 5.5 hour 

mark in Figure 18, the liquid nitrogen is refilled. This rechills the top of the cold mass and also 

causes some disturbances in the vacuum pressure due to the cryopumping and subsequent off 

gassing of the system associated with pouring the cold liquid through the vent/fill lines. 

However, after several hours the pressure levels off well below 10-6 torr. The strut temperatures 

(T4-T7) were slightly affected by the refill but not greatly so. 
24

Figure 18: Typical system temperatures and preferences 

Figure 19 shows a representation of typical MLI temperatures as a function of time. It can be 

seen that after roughly 5 hours, the temperatures leveled off and the temperatures are steady 

with a few exceptions during which refills occur. Figure 20 shows the same data, except instead 

of being plotted versus time, the averaged measurement of each sensor during steady state is 

plotted versus location within the MLI. Thus the dark blue line is the temperatures on the warm 

layer of MLI, the light blue line is the temperatures on the 16th layer from the cold side, and so 

on. For which temperature represents which data point position, please refer to Figure 8. Layer 

one is the closest layer to the cold mass. This is the main data representation that is used to 

show the degradation radius. 
25

Figure 19: Typical MLI temperatures 

Figure 20: Steady state MLI temperatures plotted vs. location
26

4 Thermal Models 

Thermal models were developed in Thermal Desktop and run on Sinda/Fluint to allow for 

scaling of the test results to larger struts and different boundary conditions. The models were 

anchored to the testing performed in this report. This section describes the thermal models. 

4.1 Model Overview 

Two models were created, a detailed model of the C-600 cold mass that was used for validation 

of approach and a model that was used to scale those results to flight applications. The 

detailed model included both of the liquid chambers, the edge guard, penetration, MLI test 

sections, and other detailed components of the testing. Its purpose was to develop and validate 

a method for modeling the penetration issue. The flight scaling model was a much more basic 

investigation that used the same method developed in the detailed model but didn’t include the 

details of the calorimeter, instead looked at a much more basic and open configuration where 

edge effects were basically ignored. 

Material properties for most of the components used in these models are included in the table 

below. A thermal conductivity and emissivity as a function of temperature was used for most of 

these materials, the table gives values for these at one temperature for reference. 

Table 6: Material properties used in the calorimeter thermal model

Component Material Property 

Thermal conductivity 

(W/m-K) 

Emissivity 

Aluminum Strut Aluminum 

6061 

k = F(T), 86.6W/m-K at 88.7K ε= F(T), 0.0275 at 80K 

file: Unpolished Aluminum

Composite 

Strut 

Carbon 

Fiber Epoxy 

k = F(T), 1.1W/m-K at 120K 

file: IM2/977-2

ε= 0.8 

file: IM2/977-2 Graphite Epoxy

Cold Mass Stainless 

Steel 304L 

k = F(T), 8.11W/m-K at 80K 

file: Stainless Steel 304L

ε= F(T), 0.128 at 80K 

file: Stainless Steel 304L 

Andy

Buffer Cryolite k = F(T), 0.00123W/m-K at 

150K 

ε= 0.953 

MLI Mylar k = F(T), 0.157W/m-K at 80K 

file: Mylar

ε= F(T), 0.0246 at 80K 

file: MLI-Andy2

Most of the material properties were taken from the NIST website; however, for the composite 

material conductivity the following formula was used: 

The emissivity used for the carbon fiber epoxy was based on testing at ambient temperature. 

4.1.1 Cryostat Detailed Model Overview 

The detailed model includes the components of the cold mass including: 

 Test and guard chambers

 Insulation surrounding the guard chamber

 Fill tubes
27

 Penetration

 Penetration insulation

 Passive Heater

 Edge Guard

 Test section of MLI

 G10 MLI support ring

The dimensions of these features were taken from the design drawings. The edge guard is a 

plastic band, shown in Figure 7, attached to the bottom of the cold mass extending the full 

thickness of the test MLI. The physical significance of this band is to expose the outer edge of 

the test MLI to a thermal profile that closely matches its own, limiting heat transfer between the 

test MLI and other portions of the Cryostat. The G10 ring supports the test MLI and holds it 

near the cold mass. 

Each layer of MLI was modeled as an individual surface. Conduction and convective heat 

transfer between layers of MLI was modeled using the Lockheed equation (8) to calculate heat 

transfer between individual nodes on each layer. Temperature dependent emissivity values 

were used with a multiplier to account for degradation of the MLI. The degradation factor was 

based on testing at ambient temperature with a sample of MLI that had been used during initial 

testing. 

The test and guard chambers were modeled as boundary nodes connected to the wetted 

surfaces of the cold mass by LN2 convection. Both chambers were assumed to be 90% full of 

liquid nitrogen. The Thermal Desktop convection model used assumes no interaction between 

convective surfaces and a constant temperature across the surface. This is an 

oversimplification in this case. The model shows a 16K change in temperature across the face 

of the cryostat for a typical test with a rapid drop moving away from the strut. Table 7 below 

shows the temperature for a point on the cold mass approximately 25.4mm (1 inch) from the 

strut centerline. It shows the test side of the cold mass is significantly higher than the LN2 

boiling point at atmospheric pressure. This temperature gradient could cause changes in the 

convective currents that would result in a significant deviation from the constant temperature 

model used. 

Table 7: Typical Cold Mass Temperatures

Test Total Heat 

Load (W) 

Strut Heat 

Load (W) 

T3 Temperature 

(K) 

P112 0.294 NA 79.984 

P115 1.227 0.645 82.716 

P116 1.529 0.56 82.445 

P117 2.906 2.38 89.047 

P106 3.298 2.906 82.075 

P103 3.92 3.04 81.383 

The guard chamber’s fill and vent tubes were modeled with a constant heat load on the top 

(non-test) surface of the guard chamber. The geometry of the test chamber fill tube was 

included in the model. 
28

Figure 21 is a cutaway view of the detailed model showing the temperature profile of the 

area around the strut. The hottest area is the protruding section of the buffer material shown in 

pink. Not shown in this view is the MLI insulation that surrounds the strut and keeps it at a fairly 

uniform 150K in the area of the MLI. 

Figure 21: Section View of Detailed Model

4.1.2 Scaling Model Overview 

A second model was created which did not include the geometry which would not be present in 

a flight application such as the guard chamber. The model’s cold mass was decreased to 2mm 

Buffer

MLI

Strut

Cold Mass
29

to better simulate a flight application. To keep the model grounded in test data other 

parameters such as the heat conduction scaling factors between layers of MLI, were retained. 

The following cases were considered. 

Case Parameter to 

Vary 

Values Notes 

1 Strut Diameter 6.35mm, 

12.7mm, 

25.4mm, 

50.8mm, 

76.2mm, 

101.6mm 

127mm, 

152.4mm (0.25”, 

0.5”, 1.0”, 2.0”, 

3.0”, 4.0”, 5.0”, 

6.0”) 

12.7mm (0.5”) thick Cryolite 

buffer, 25 layers, Aluminum 

penetration 

2 Buffer Thickness 6.35mm, 12.7mm 

25.4mm (0.25”, 

0.5”, 1.0”) 

Cryolite Buffer 

For 76.2mm (3”) and 152.4mm 

(6”) penetrations 

3 MLI Layers 25, 50, 75, 100 

layers 

For 76.2mm (3”) penetration 

with 12.7mm (0.5”) buffer 

For 152.4mm (6”) penetration 

with 25.4mm (1.0”) buffer 

4 Strut Materials Aluminum, 

Carbon Fiber 

Repeat Case 1 with Composite 

Repeat Case 2 with 76.2mm (3”) 

composite penetration 

Repeat Case 3 with 76.2mm (3”) 

composite penetration 

5 Cold Boundary 

Temperature 

LN2, LH2 Repeat Cases 1 thru 4 for LH2 

temperature at cold mass 

4.1.3 Cryostat Detailed Model Validation 

The models were validated by comparing the test chamber heat leak, measured by LN2 boiloff, 

to the model heat leak when the model temperatures matched test conditions. Parameters 

representing setup conditions such as the amount of contact between the MLI and the cold 

mass were varied in order to match temperatures between the model and test results. An 

acceptable match was considered to be an average of 5% as calculated by the following 

ൌ ݃ݒܽݎݎܧ

ሻ ݈݀݁݋ܶ݉ െ ݐݏ݁ݐܶሺݏܾܽ ∑

ݐݏ݁ݐܶ

݊

1

݊
30

The following measurements were used for validation 

Strut Temperatures T4, T5, T6, T7 

MLI Layer Temperatures MLI Layer 3 T8, T9, T10, T14 

MLI Layer 8 T15, T16, T17, 

T18 

MLI Layer 16 T19, T20, T21, 

T25 

MLI Layer 25 T26, T27, T11 

Table 8 below shows how well the model predicted test temperatures along the penetration and 

in the MLI as well as the predicted change in heat leak due to the penetration compared to 

actual measurements. 

Table 8: Penetration Heat Load Predicted From Model Compared With Test Result Calculations

Test Average Percent 

Error In Temperature 

Match Between Test 

Results And Model 

Change In Heat Load 

Due To Penetration 

Strut MLI Delta Q – 

Test (W) 

Delta Q – 

Model (W) 

12.7mm AL Strut, 

25.4mm buffer 

P106 Run 

1 

1.03% 1.87% 0.262 0.0199 

6.35mm AL Strut, 

25.4mm buffer 

P115 Run 

1 

1.03% 0.71% 0.288 0.0175 

25.4mm AL Strut, 

12.7mm buffer 

P116 Run 

1 

0.25% 2.96% 0.656 0.0806 

25.4mm AL Strut, 

25.4mm buffer 

P117 Run 

1 

0.23% 2.61% 1.135 0.0785 

25.4mm Composite 

Strut, 25.4mm 

buffer 

P121 Run 

1 

2.13% 1.12% 0.252 0.002 

For the model the change in heat load due to the penetration was determined according to the 

following: 
31

1. Setting the temperature at the warm end of the strut to match the test results by varying 

the passive heater to strut contact resistance. Normally T7 was used since T13 

consistently read high. 

2. Changing the contact resistance with the cold mass to match the temperature profile 

along the strut. 

3. Changing the amount of MLI contact with the cold mass to match the temperatures of 

each of the layers of MLI. 

4. The following model results were recorded: 

QTotal – sum of heat transferred to the cold mass 

Qstrut – heat transferred from the strut to the cold mass 

5. The model was run a second time without the penetration to determine the baseline heat 

leak: 

QMLI_0 – heat from the unpenetrated MLI 

QMLI_Pen – heat leak through a strut sized section of MLI 

This last term is the amount the heat leak that would have happened through MLI had 

the strut not been there. 

6. The change in heat load was calculated from: 

∆QMLI = QTotal - Qstrut + QMLI_Pen – QMLI_0 

The model consistently underestimated the change in heat leak due to the penetration. There 

are several reasons for this. 

 

1. Penetration Heat Load Calculation 

The model gave higher estimates for the penetration contribution to the total heat load 

compared to the method used to calculate the test data. This is shown in Table 9. 

During testing the penetration heat load is determined using thermocouple data along 

the strut in one of two ways. 

Test Method 1 

The heat load is calculated at a specific point on the strut where the crosssection changes from tubular to solid. Temperature measurements on either 

side of this transition are used to determine the heat load at the transition by 

knowing the material properties, strut cross-section, and distance between the 

measurements. The load at the transition is used as the penetration heat load. 

Using this method any heat that is transferred between the strut and the buffer is 

not included in the penetration heat calculation. 

Test Method 2 

Temperature readings all along the strut are used to determine heat flowing 

between each set of thermocouples. The average heat load is used as the 

penetration heat load to the cold mass. 

Both methods assume there is no heat transfer to/from the strut other than the 

passive heater. Both use Fourier’s law of conduction 

Qstrut = kA∆T/x 
32

Model Method 

The model determines the struts contribution to the Test Chamber heat leak by 

matching the temperatures along the strut just as the second method does but it 

also accounts for heat transferred between the strut and its surroundings 

including the buffer and the heat leak through the strut MLI. 

Table 9: Penetration Heat Load - Test vs Model 

Test Penetration Heat Load 

Test (W) Model (W) 

12.7mm (1⁄2”) AL Strut, 

25.4mm (1”) buffer 

P106 Run 1 2.906 0.844 

6.35mm (1⁄4”) AL Strut, 

25.4mm (1”) buffer 

P115 Run 1 0.645 0.916 

25.4mm (1”) AL Strut, 

12.7mm (0.5”) buffer 

P116 Run 1 0.56 0.872 

25.4mm (1”) AL Strut, 

25.4mm (1”) buffer 

P117 Run 1 1.476 1.78 

25.4mm (1”) Composite 

Strut, 25.4mm (1”) 

buffer 

P121 Run 1 0.153 0.397 

The model shows an appreciable heat exchange between the strut and surroundings. 

Table 10 shows an example of test data with an aluminum strut compared to model 

output and the closeness of the model fit between known temperatures. With the 

exception of T13, the model predicts the strut temperatures fairly accurately. 

Table 10: Test P115 Strut Temperature Data

Thermocouple 

Reference 

Designation 

Test 

Temperature 

(K) 

Model 

Temperature 

(K) 

Distance 

from 

cold 

mass 

(mm) 

Estimated Q 

between 

temperatures 

Estimated 

Q (W) 

(From 

Test 

Calcs) 

T13 263.3 224.6 200 T13 – T7 1.875 

T7 189.85 192.8 118 T7 – T6 0.765 

T6 167.1 168.6 61 T6 – T5 0.72 

T5 153.0 152.2 25 T5 – T4 0.57 

T4 146.2 144.5 4 - 

Heat to cold mass from strut – Test 0.645 

Heat to cold mass from strut - Model 0.916 

Another model was created to eliminate uncertainty in the actual penetration heat leak. 

This model found the change in MLI heat leak by summing the MLI and buffer 

contributions only. The strut was thermally disconnected from the cold mass and was 

given a fixed temperature gradient representative of the test results. This model also 

showed the penetration has a very small effect on the MLI heat leak for buffers 6.35mm 

(1/4”) and thicker. 

2. Convection 

The standard convection Sinda-Fluent equations used assume the convective surface is 

contiguous and uniform in temperature or heat flux, however, the tests setup creates a 
33

high heat flux in the center of the cold mass that causes a temperature gradient across 

the cold mass. That could create changes in the convective currents that would alter the 

heat transfer to the cold mass. The convective model also treats the guard chamber as 

a contiguous surface where it should also be modeled as a ring. 

Figure 22 shows the predicted temperature across the cold mass along with some actual 

test data. 

Figure 22: Predicted Temperature Accross Cold Mass

A simplification of the problem is possible where the model includes only the buffer and 

MLI. The temperature profile of the third layer of MLI is used as the boundary 

temperature and the temperature profile of the strut is used as another. This avoids 

uncertainty in convection and conduction along the strut and includes only the portion of 

the system that is well instrumented but it would only capture the penetrations effect on 

MLI layers 4 through 25. The effect of the buffer and first three layers of MLI would not 

be captured. 

3. Test – Guard Chamber Heat Exchange 

The model shows there is some heat exchanged between the test and guard chambers. 

The method of calculating the change in heat leak due to the penetration from test 

results relies on the test and guard chambers being at thermal equilibrium so any heat 

exchanged results in a bias. For the model the heat exchanged is driven by very slight 

differences in the temperature along the test section of the cold mass and is on the 

same order of magnitude as the unpenetrated MLI heat load. 

4.1.4 Detailed Model Validation Discussion 

MLI Performance 

The model and test program may underestimate the effects of the penetration. The tests 

were performed with a relatively poor performing MLI which did not seem to be in 

contact with the cold mass for most of the tests. This is evidenced by the relatively low 

difference in temperatures between layers and the high temperature on layer 3. 
34

Because the MLI stack was at relatively high temperature there was a fairly small 

gradient between the strut and MLI. In a flight application the penetration may be 

warmer or colder than the surrounding MLI and so would have a different effect on a 

better performing MLI. 

Test Strut Temp at 

Base – T4 (K)

MLI Temp 

at Layer 3 

(K) 

MLI Temp 

at Layer 16 

(K) 

MLI Heat 

Flux 

(W/m^2) 

P112 NA 217.9 239.0 8.55 

P117 107.3 181.2 261.1 NA 

4.1.5 Scaling Model Results 

The model was scaled to find the change in the MLI heat leak for conditions that were not 

tested. The strut was thermally disconnected from the cold mass and the end temperatures 

were fixed to match a typical temperature profile from the testing, the base of the strut at 150K. 

The model was run with and without the penetration and the change in heat load was calculated 

from: 

∆QMLI = Qbuffer + (QMLI – QMLI_0) + QMLI_Pen

The advantage of this method is that it is independent of penetration parameters such as 

thickness. The penetration is still thermally connected to the buffer insulation and the remainder 

of the thermal network is unaffected. 

Case 1: Variation of Penetration Diameter 
35

Figure 23 shows the effects of varying the penetration diameter on the MLI heat leak using a 

6.4mm (0.25”) Cryolite buffer. The model was run for a cold boundary temperature of 20K and 

77K and with composite and aluminum struts. The variation in predicted heat leak between 

these cases was about the magnitude of the model noise so the figure shows the average of all 

model runs. 

Figure 23: Change in Heat Leak with Penetration Diameter
36

Case 2: Variation of Buffer Thickness 

Figure 24 below shows the effects of varying the buffer thickness around the penetration. Again 

the base of the strut was held at a fixed temperature of 150K with the strut thermally 

disconnected from the cold mass. The model was run for a cold boundary temperature of 20K 

and 77K and with composite and aluminum struts. As with the penetration diameter the 

variation in predicted heat leak between these cases was about the magnitude of the model 

noise so the figure shows the average of all model runs. 
37

 

Figure 24: Change in Heat Leak with Buffer Thickness

Case 3: Variation of MLI Layers 

Figure 24: Change in Heat Leak with Buffer Thickness

 below shows the effects of varying the number of MLI layers. The model was run for a cold boundary 

temperature of 20K and 77K and with for both 76.2mm (3 inches) and 152.6 mm (6 inches) aluminum struts. 
38

However for the 76.2 mm strut the variation in predicted heat leak between these cases was about the 

magnitude of the model noise so 

Figure 24: Change in Heat Leak with Buffer Thickness

 shows the average of those model runs. 

Figure 25: Change in Heat Leak with number of MLI layers
39

5 Test Results and Analysis 

A total of 23 tests were run in order to complete the test matrix in section 3.5, some of these involved subtests 

where parameters such as the warm boundary temperature was changed or the test was repeated. These 

tests and the results are summarized in 

Figure 26. Six of these tests were no penetration tests, three involved no integration (also 

known as hot poker), six were involved in the material trade study, two involved temperature 

matching, six were involved in the strut size analysis, and three tests were run using the 

composite strut. An additional test was run where the MLI was disturbed to create a non-ideal 

MLI blanket. 

Figure 26: Test Result Summary for Penetration Calorimetry
40

5.1 No Penetration – Null testing 

Prior to punching a hole in any MLI blanket, a null test was run to determine the no penetration 

heat load. Temperature sensors were installed in the blanket per Figure 7 to get baseline data 

for temperature gradients in these particular undisturbed MLI blankets. The results for the no 

penetration testing are shown in Table 11. It is interesting to note the apparent dependence of 

the heat load based on the diameter of the blanket. However, the temperature profile of P100 

(see 

Figure 27) shows that the edges of the layers were shorting to the versify edge guard. Figure 

28 (temperature profile of P118) shows that by removing the outer 5 mm or so, that shorting 

was generally avoided, though there are still some gradients and edge effects. When the 

diameter of the sample was kept the same, the heat loads were repeatable within 6% for nearly 

identical blankets. This was an important result as it removes any biasing of the test data based 

on null blanket variations. 

Table 11: No penetration MLI test results

Series 

# 

Blanket Serial 

Number 

Outer 

Diameter 

(mm) 

Heat load 

(W) 

Blanket 

Usage 

P100 SN2 305 0.130 Buffer 

Materials 

P107 SN3 305 0.122 None 

P109 SN5 300 0.203 Temperature 

Matching 

P112 SN6 285 0.294 Penetration 

Sizes 

P118 SN3 295 0.194 Composite 

Strut 

P122 SN7 300 0.191 Retesting 
41

Figure 27: Temperature Profile for P100 

Figure 28: Temperature Profile for P118 
42

5.2 No Integration 

No integration testing was run with both aluminum and composite struts. For the aluminum 

strut, a 0.0159 m (0.625 inch) hole was punched through the MLI blanket which the 0.0127 m 

diameter strut was placed. The 0.0259 m (1.020 inch) diameter composite strut was placed 

through a 0.0270 m (1.063 inch) hole in the MLI blanket. The delta heat load was 0.50 W from 

the aluminum strut and 0.31 W for the composite strut. The degradation radius was over 100 

mm for both tests. The temperature profiles for the aluminum strut are shown in Figure 29. The 

No Integration testing served as a general reference point as what was thought to be the worst 

integration method to prevent degradation of the blanket. However, based on the test data, it is 

clearly better than too large of a buffer or a buffer that is not highly pliable, but is not as good as 

a properly sized buffer. 

Figure 29: Temperature profiles for No Integration testing on Aluminum strut

5.3 Buffer Materials 

Buffer material testing was performed using the 0.5” aluminum penetration. The penetration 

was wrapped with four layers of MLI when not within the MLI/buffer area. Three different 

materials and various thicknesses were tested at two different sizes (0.5” diameter, and 1.0” 

diameter). Additionally the 0.5” diameter (0.25” radius) buffer was tested with no fill material, 

essentially a vacuum buffer. Based on the performance of P106, the Cryo-Lite system was 

determined to be the best performing. However, all insulation systems outperformed the 

vacuum buffer case. Figure 30, Figure 31, and Figure 32 show the different materials installed 

as a buffer. 

During the buffer testing, it was noticed that the degradation radii was not necessarily related to 

the integration thermal losses. This is because the buffer allows heat load through it while is 

isolates the MLI from the strut. Thus the buffer operation is a function of how well it thermally 

isolates the MLI from the penetration in addition to the thermal properties of the buffer. So a 

buffer material that has better surface contact properties (i.e. less thermal contact), yet a higher 

thermal conductivity, could provide a better overall result as that material could be installed in a 

thinner application. 
43

In addition to outperforming the other materials thermally, Cryo-Lite was also the easiest 

material to handle. It can be easily squeezed down to fit within the cut out hole, allowing it then 

to puff back out and completely fill the hole. The aerogel blankets and bead pack could not 

completely refill the hole, and thus were susceptible to cracks forming between the buffer 

material and the MLI blanket. This pliability in combination with less material to thermally touch 

with the MLI surfaces made the Cryo-Lite the best performing material. 

Table 12: Results from buffer material testing at different buffer thicknesses

Material ∆Q at 0.5” 

(W) 

∆Q at 

1.0” (W) 

Aerogel Blanket 0.764 0.942 

Aerogel Beads N/A 0.759 

Cryo-Lite 0.75 0.262 

Vacuum 0.979 N/A 

Figure 30: Aerogel blanket buffer - 0.5" (0.25" radius) 
44

Figure 31: Aerogel beads packed in nylon and installed on calorimeter 

Figure 32: 1" (0.5" radius) Cryo-Lite buffer

Following the materials trade study, testing was done on multiple sized penetrations. The 

penetrations are described in section 3.3. Aluminum struts of 0.25” and 1.0” were added to the 

0.5” tests already performed. Cryo-Lite total thicknesses of roughly 0.5” and 1.0” were tested on 

each strut. Figure 33 shows a typical two dimensional temperature profile for the buffer testing, 

where generally the impacted area was very small. 
45

Figure 33: P115, 1 inch buffer, 0.25" strut 2-D thermal projection 

Table 13 shows the test results for the various sized penetration cases. It is noted that there 

was much more uncertainty with the 1” aluminum penetrations because the temperatures of the 

struts were colder as the warm boundary temperature was not able to be maintained on the 

strut, but was maintained on the MLI. The top section shows the test numbers, the middle 

section the parasitic heat loads, and the bottom section shows the degradation radius. The 

results indicate that the Cryo-Lite material is best suited for use as a thin material to isolate the 

penetration from the MLI. For the 0.25” penetration, the testing shows that a quarter inch thick 

radius buffer is sufficient and that larger thicknesses hurt performance. However, for the larger 

penetrations, the half inch radius buffer appears to be closer to the optimal thickness. The 

degradation radius appears to be within 2 penetration diameters on the preferred thicknesses. 
46

Table 13: Test Results for Cryo-Lite Buffers on various penetration sizes

To assess the performance of the Cryo-Lite buffer across variable environments, the P116 

configuration was run at a warm boundary temperature of 330 K. To account for the null MLI 

heat load difference, the subtracted out load was multiplied by the ratio of warm boundary 

temperatures raised to the fourth power. The resulting parasitic heat load increased from 0.68 

W in the 297 K test to 0.78 in the 325 K test, and increase of 15%. The degradation radius still 

was greater than 100 mm, so no change was seen. If these two tests are used to create a 

power factor for Warm Boundary Temperature, then the exponent solves to 1.56 yielding the 

equation: 

Additional testing with buffers was done on the composite strut. With the 0.5” buffer, a parasitic 

heat load of 0.256 W was measured with a degradation radius of 51 mm. With the 1.0” buffer, a 

parasitic heat load of 0.252 W was measured with a degradation radius of 35 mm. There was 

little thermal benefit from increasing from the 0.5” to the 1.0” around the composite strut. 

However, the composite strut showed much lower thermal penalties than the aluminum strut. 

This is due mostly to the lower thermal conductivity of the strut itself. 
47

Figure 34: Composite strut installed on calorimeter with a Cryo-Lite buffer 

5.4 Temperature Matching 

Temperature matching is the analytically optimal method for integrating a penetration into an 

MLI blanket. It involves determining the temperature gradients through the installed MLI blanket 

as well as along the penetration and then cutting an adapter to attach the MLI to the penetration 

where the temperatures are the same. Table 14 shows a summary of the temperature matching 

locations. 

Temperature gradient predictions were run prior to any testing (see Figure 35) using the double 

aluminized mylar, double silk netting Lockheed equation for the MLI temperature gradients and 

NIST data of Aluminum 6061-T6 and G10 (composite) for the penetration temperature gradients 

(8) (9). The MLI was done on a layer by layer basis (25 layers), while the penetrations were 

each done in with 21 nodes and 20 even conductors, assuming a constant cross section the 

length of the penetration. The warm boundary was assumed to be 293 K, the cold boundary 77 

K, and the interstial pressure 5x10-6 torr. The layer density of the MLI was assumed to be 20 

layer/cm. 

Following the initial no penetration testing of both SN2 (P100) and SN6 (P112), the actual 

temperature gradients through the MLI, as well as along the penetration (from P110), were 

compiled and used to determine where the adaptions between the MLI and penetration should 

be. Figure 36 shows the temperature gradients through the MLI for the Lockheed predicted 

case in addition to P100 and P112. In addition to testing the “As Measured” temperature 

matching at a warm boundary temperature of 293 K (the design temperature), the warm 

boundary temperature was ramped up to 325 K to test off-nominal conditions of the temperature 

matching and assess the flexibility of temperature matching. 
48

Table 14: Temperature matching locations along the strut (total length .203 m)

Layer # Lockheed 

Strut Location 

(m) 

As Measured 

Strut Location 

(m) 

1 0.020 N/A 

3 0.046 0.076 

8 0.102 0.140 

16 0.157 0.178 

25 0.203 0.203 

Note: As measured test did not include layer 1 

Figure 35: Predicted MLI and penetration temperature gradients 

The initial testing of the Lockheed and NIST temperature matched sample showed a heat 

penalty (Q) of 0.75 W. Testing of the “As Measured” temperature matched sample shows a 

much lower heat penalty of 9.5 mW. This suggests that equations used did not adequately 

predict the temperature gradients of both the MLI and the penetration. Contact resistance was 

neglected between the penetration and the cold mass. This may have played a role in the 
49

errors on the penetration side, but as shown in Figure 36, the MLI temperatures were off as 

well. 

When the environment of the “As Measured” sample was changed to a warm boundary 

temperature of 325 K, the heat penalty increased from 9 mW to 0.57 W. This shows that 

changing the environmental temperature, even only 30 K, can drastically affect the performance 

of the temperature matched penetration. 

In theory temperature matching is ideal, however, since conduction and radiation heat transfer 

vary to different powers of temperature, in general a temperature matched solution can only be 

applied at a single environmental temperature. For locations where the temperature might 

greatly change (such as in low earth orbit), this appears to have unintended consequences in 

off-nominal conditions. 

Figure 36: Temperature gradients through MLI showing predicted, P100, and P112. Also shown for reference are 

the strut temperatures.
50

6 Conclusion 

Over twenty-three tests were run to help characterize the thermal performance impacts of 

penetrating MLI. Testing included the development and fabrication of a new calorimeter and 

test method for two dimensional thermal performance testing. The testing included null testing 

of every blanket, no integration testing, buffer testing, and temperature matching testing with 

different size and material penetrations. The preferred method of isolating penetrations was 

shown to be the buffer method, and Cryo-Lite the best material to use as a buffer. The thermal 

degradation or parasitic heat load was shown to be a function of strut diameter, buffer thickness, 

buffer material, warm boundary temperature, and penetration material. The buffer method was 

shown to be easier to develop, more robust, and less variable over multiple conditions and 

environments. 

Figure 37: Comparison between different integration methods for a half inch aluminum penetration

Similarly, a thermal model was developed and validated against the test results. The model was then used to 

extrapolate predictions to systems that are larger than could be tested in the current set up. While an overall 

mathematical model could not be developed, a set of representations were developed for the main driving 

variables. These equations should be used to determine specific cases by using the ratio of a single chosen 

case and the actual case in each of the plots compared to the reference data point (using a 6.4mm thick 

Cryolite buffer and either a 152.4 or 76.2 mm penetration). Since the same buffer thickness is not enveloped 

between the penetration diameter ( 
51

Figure 23) and number of layers ( 

Figure 24: Change in Heat Leak with Buffer Thickness

), the second reference point is required. The second reference point can be used on either of 

the two graphs, but it is shown on the number of layers calculation. Th should be in the units of 

Kelvin. 

Thus an overall approach to the integration of penetrations into MLI blankets has been 

developed. Thermal modeling was anchored to testing on a modified flat plate calorimeter. 

This is readily usable for designing and preliminarily specifying designs for cryogenic systems. 
52

References 

1. Coston, R. M., Murray, D. O. and Lambert, L. A Study On High‐Performance Insulation Thermal Design

Criteria. Sunnyvale, CA : Lockheed Missiles & Space Company, 1967. pp. 42‐104. LMSC‐A847882, Vol. I.

2. Sumner, I. E. Degradation of a Multilayer Insulation Due to a Seam and a Penetration. Cleveland, OH :

Lewis Research Center, 1976. TN D‐8229.

3. Johnson, W. R. and Sprague, E. L. Analytical Investigation of Thermal Degradation of High

Performance Multilayer Insulation in the Vicinity of a Penetration. Cleveland, OH : National Aeronautics

and Space Administration, Lewis Research Center, 1968. NASA TN D‐4778.

4. Lord Corporation. Aeroglaze Z306 Polyurethane Coating. [Online] Lord Corporation, 2012. [Cited:

March 23, 2012.] http://www.lord.com/products‐and‐solutions/coatings/product.xml/611.

5. Coffman, B. E., et al., et al. Aerogel Blanket Insulation Material for Cryogenic Applications. [book

auth.] J.G. Weisend II. Advances in Cryogenic Engineering Vol. 55B. NY : American Institute of Physics,

2010, pp. 913‐920.

6. Koravos, J. J., et al., et al. Nanogel aerogel as a load bearing insulation material for cryogenic systems.

[book auth.] J. G. Weisend II. Advances in Cryogenic Engineering, Vol 55B. Melville, NY : American

Institute of Physics, 2010, pp. 921‐927.

7. Johns Manvill Corporation. [Cited: July 9, 2012.]

http://www.jm.com/engineered_products/hpi31_cryolite.pdf.

8. Keller, C. W., Cunnington, G. R. and Glassford, A. P. Thermal Performance of Multi‐layer Insulations,

Final Report. Sunnyvale, CA : Lockheed Missile and Space Company, 1974. Contract NAS3‐14377.

9. Cryogenic Technologies Group ‐ Material Properties. NIST Material Measurement Laboratory. [Online]

National Institute of Standards and Technology. [Cited: March 23, 2012.]

http://cryogenics.nist.gov/MPropsMAY/material%20properties.htm.
53

Appendix A – More data plots 
54
55
56
57
58

Appendix B – Sample Calculation 

Calculate the degradation due to a 104 mm (4 inch) pipe going through 60 layers of MLI using 

an 8 mm (~0.75 inch) Cryolite buffer. The system has a cold boundary temperature of 77 K and 

a warm boundary temperature of 297 K. 

For reference case one use 25 layers of MLI with a 76.2 mm penetration and a 6.4 mm Cryolite 

buffer. 

Qref equals 0.052 W from 

Figure 23. 

For reference case two use 25 layers of MLI with a 76.2 mm penetration and a 12.7 mm Cryolite buffer. 

Qref’ then equals 0.086 W from 

Figure 24: Change in Heat Leak with Buffer Thickness

. 

Q actual for the pipe diameter (using a 104 mm penetration with 25 layers of MLI & 6.4 mm Cryolite buffer) 

is 0.076 W from 
59

Figure 23. 

Q actual for the buffer thickness (using an 8 mm Cryolite buffer with 76.2 mm penetration and 

25 layers) is 0.088 W from Figure 24. 

Q actual for the number of layers (using 60 layers with a 12.7 mm buffer and a 76.2 mm penetration) is 0.038 

W from 

Figure 24: Change in Heat Leak with Buffer Thickness

. 

Since the WBT is 297 K, we can neglect that term as 1 
60

Appendix C – Uncertainty Analysis Derivations 

The effect of a penetration on the heat leak thru MLI was not measured directly; it is 

calculated based on the measurements of other parameters during two tests. In the first test 

the heat leak through MLI without any penetration is measured once steady state conditions 

are reached, then a penetration is added and the heat leak again measured. The effect of the 

penetration on the steady state heat leak of the MLI can be calculated: 

Where, 

∆QPen – Change in heat leak of MLI due to penetration 

Qmeas – Total measured heat leak of MLI and penetration during the second test 

QMLI – Heat leak thru MLI only measured in the first test 

QStrut – Heat leak due to the penetration 

The heat leak through the penetration, QStrut, is calculated from the temperature change over a 

section of the penetration outside of the MLI: 

Where, (2)

k – thermal conductivity of the strut material 

A – crossectional area of the strut in the area where ∆T is measured 

∆T –change in temperature along the length of the strut 

x – distance between temperature measurements 

The heat leak through the MLI, QMLI, is determined by measuring the rate of boiloff for liquid 

nitrogen: 

Where, 

V – Volumetric flow rate of nitrogen 

ρ – Density of gaseous nitrogen 

hfg – Heat of vaporization of nitrogen at atmospheric pressure 

Qmeas is determined the same way during the penetration test. So the change in heat leak is: 

∆Qpen = f(k, A, ∆T, x, V, ρ, hfg) 

Each parameter contains an amount of uncertainty which combines to cause error in the final 

calculated thermal conductivity value. Returning to the definition of uncertainty above, the 

total uncertainty is: 
61

From the definition of Qstrut,

The partial derivatives for these four variables are shown: 

Substituting and simplifying results in, 

From the definition of QMLI
62

The partial derivatives for these four variables are: 

Substituting and simplifying: 

And from the definition of Qmeas

The Qmeas and QMLI uncertainty terms are the same. Substituting the partial derivatives results 

in: 
63

 (8)

The error terms in brackets for QMLI and Qmeas will be the same in the previous equation. The 

flowrates, however, will be different for each test: 

And the percent error would be,