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NATIONAL ^ITOTF. OF STANDARDS

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United States Department of Commerce

National Institute of Standards and Technology

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REFERENCE PUBLICATIONS

NISTIR 3966

A CRYOGENIC 4.4 MN MECHANICAL TEST

SYSTEM FOR LARGE SCALE TESTS OF

COMPOSITE SUPPORT STRUTS

R.P. Walsh

R.P. Reed

J.D. McColskey

W. Fehringer

J.R. Berger

-QC—

100

.U56

3966

1992

C.2

N/STIR 3966

A CRYOGENIC 4.4 MN MECHANICAL TEST

SYSTEM FOR LARGE SCALE TESTS OF

COMPOSITE SUPPORT STRUTS

R.P. Walsh -

R.P. Reed

J.D. McColskey

W. Fehringer

J.R. Berger

Materials Reliability Division

Materials Science and Engineering Laboratory

National Institute of Standards and Technology

Boulder, Colorado 80303

March 1992

U.S. DEPARTMENT OF COMMERCE, Barbara H. Franklin, Secretary

TECHNOLOGY ADMINISTRATION, Robert M. White, Under Secretary for Technology

NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY, John W. Lyons, Director

TABLE OF CONTENTS

1.0 EXECUTIVE SUMMARY 1

2.0 INTRODUCTION 2

3.0 MECHANICAL TESTING: FACILITIES AND INSTRUMENTATION 3

3.1 Facility Development 3

3.2 Instrumentation 6

4.0 EXPERIMENTAL RESULTS AND DISCUSSION 7

4.1 Ultimate Compressive Strength: Short Specimens 7

4.2 Ultimate Compressive Strength: Long Specimens 8

4.3 Fatigue Tests 11

5.0 SUMMARY AND CONCLUSIONS 15

iii

1.0 EXECUTIVE SUMMARY

The mechanical tests reported here were conducted by the Materials

Reliability Division of the National Institute of Standards and Technology.

These tests were intended to validate the full scale structural behavior of fiber

reinforced polymer (FRP) support struts in a thermal and mechanical environment

similar to the actual service conditions.

Two types of tests were performed: ultimate compressive strength and

fatigue. Thermal conditions for the tests ranged from isothermal room

temperature tests to an imposed thermal gradient of cryogenic- to -room

temperatures. The development of the unique facility necessary for the large

scale cryogenic tests is detailed in this report.

A summary of all tests is presented in Table 1. Many tests demonstrated

that a problem exists with the aluminum alloy (7075-T6) end fittings. Failure

seemed to propagate from the failed end fittings into the tube , thereby

preventing the tube from reaching its theoretical ultimate strength.

1
2.0 INTRODUCTION

The superconducting magnetic energy storage (SMES) coil magnet requires an

array of large radial compression members to resist dimensional change of the

coil. The design for the compression members uses FRP composite tubes. These

FRP tubes must withstand cyclic compressive loads, in a vacuum environment, at

temperatures ranging from cryogenic (1.8 K) to ambient (295 K) . The Materials

Reliability Division of the National Institute of Standards and Technology (NIST)

was contracted to assess the viability of the proposed tube design (Figure 1)

.

Towards such an assessment, NIST performed full scale thermal and mechanical

qualification tests of prototype composite support struts.

As a precursor to the full scale tests, a 4.4 MN capacity servo-hydraulic

test machine was modified for testing at cryogenic temperatures. Because of the

uniqueness of this facility, the details of the cryostat design are presented in

the first part of this report. Additional details pertaining to instrumentation

are also presented in the same section of the report.

The test program divides into three distinct areas:

(1) Ultimate compressive strength tests of short specimens.

(2) Ultimate compressive strength tests of full scale specimens.

(3) Compressive fatigue tests of full scale specimens.

The experimental results for each of the above tests are discussed in the final

section of the report. Additional details are presented on the failure analysis

for each type of test, especially for those tests where failure originated with

the aluminum end fittings of the tube.

2
3.0 MECHANICAL TESTING: FACILITIES AND INSTRUMENTATION

3 . 1 Facility Development

Full scale mechanical qualification tests of the coil support struts were

needed for design optimization, fitness for service, and safety. Two types of

full scale qualification tests were desired: ultimate compressive strength tests

and compressive cyclic fatigue tests. The tests required mechanical loading (to

loads in excess of 3000 kN) of the composite tube and an imposed thermal gradient

(4 to 295 K) over the length of the tube. To meet these specifications we

developed a mechanical test cryostat for a 4.4 MN capacity servo -hydraulic test

machine (Figure 2)

.

The cryostat design was driven by thermal specifications, load capacity,

and the size constraints of the test machine. The specimen would be tested

between two compression platens at extremely different temperatures. One platen

was maintained at room temperature while the other was cooled to extremely low

temperatures using liquid helium. The cold platen represents the superconducting

coil operating at 1.8 K and the warm platen represents the ambient temperature

support trench that the struts are anchored to. In service, each coil support

strut thermally loads the 1.8 K coil. To minimize this thermal load, the struts

are equipped with liquid nitrogen thermal intercepts at a strategic position

along the length of the strut. To best simulate the service environment, the

test specimen and test fixture were also equipped with a liquid nitrogen

intercept at the appropriate location. The specimen would cool in vacuum by

thermal contact conductance with the platens and thermal intercept before

testing

.

Since the specimen is long (2m) and one end needed to be at liquid helium

temperature (4 K) , the cryostat was designed to have the cold end isolated deep

in the bottom of a closed bottom super- insulated Dewar. This dictated that a

vertical thermal gradient be set up in the test machine with the warm end at the

top, thermally anchored to the test machine's heavy framework. This allowed the

warm end temperature to be controlled naturally as in the actual application.

The cooling of the lower platen had to be considered carefully from the

standpoint of feasibility and cost. The system (cooling and structural) had to

be designed efficiently due to the cost of the expendable liquid helium coolant.

As in many other design problems, there were material and design selection tradeoffs. The conflict was that a 4.4 MN capacity load frame needed to be sized for

minimal conductive heat loss and mass. The cooling system for the lower platen

(a liquid helium flow heat exchanger) needed to cool this large mass and

equilibrate the constant conductive heat losses.

The structure was designed with the lower platen suspended in the Dewar by

a four-post cage. The lower platen was efficiently sized to react the bending

stress at the 4.4 MN maximum load. The four-post cage also had to react to this

load in tension while supplying a direct conduction path to the platen. The

lower (cold) platen was a 100 mm thick by 533 mm diameter maraging steel disc

with four 60 mm diameter holes for mounting. Four 60 mm diameter titanium rods

were used to mount the platen and to transfer the load from the platen to an

3
upper reaction plate. Titanium was used because of its high strength- to -weight

ratio and low thermal conductivity. Maraging steel was used for the platen

because its high strength allowed for minimal sizing of the part that needed to

be cooled with liquid helium. Another reason for selecting maraging steel was

the need to solder the copper tube heat exchanger directly to it.

With the optimum physical size of the load frame defined, the cooling

system could now be designed. Since the actual support struts in a vacuum are

cooled mainly by conduction, the system was designed to cool the specimen by

conduction. The cryostat was designed with the specimen and load frame in a

vacuum chamber. Liquid helium (4 K) and liquid nitrogen (76 K) flowed through

tubing that kept them isolated from the test chamber vacuum space. The liquid

cryogens flowed through double -walled, vacuum- insulated plumbing lines until they

reached the temperature station that they were supposed to cool (liquid helium

at the lower platen and liquid nitrogen at the thermal intercept) . Then singlewalled copper tubing was used as a heat exchanger. For the lower platen this

single -walled copper tubing was soldered directly to the bottom and the

circumference of the maraging steel disc. This allowed cooling of the platen by

conduction from the convectively cooled copper tubing. The liquid nitrogen

thermal intercept was a large copper ring with the single-walled copper tubing

soldered to it. This ring was thermally anchored to both the specimen and

fixture with copper straps and copper braided cable respectively. Thus, the

copper ring, the specimen, and the fixture were cooled by conduction from the

convectively cooled copper tubing. The exhaust plumbing for both heat exchangers

brought the boiled gases to the outside atmosphere.

The liquid nitrogen thermal intercept was simple compared to the liquid

helium cooling system. This was due to a number of factors: the ease with which

liquid nitrogen can be handled, the lower thermal load it was required to handle,

and the relatively low cost compared to that of liquid helium. For this reason,

the liquid helium cooling system is described in more detail.

The liquid helium was responsible for cooling the mass of the lower platen

and for equilibrating the heat input from the four titanium support rods. The

liquid helium was supplied to the copper tube heat exchanger in 5 mm diameter

vacuum- j acke ted tubing. A 1.0 1/min liquid helium flow rate was assumed; then,

fluid mechanics and heat transfer equations were used to design the heat

exchanger. The objective was to have as much length of tubing as possible

soldered directly to the platen. The tube diameter was increased ~3 mm, for

every 6 m of length, for flow efficiency. The length of copper tubing used to

make the liquid helium heat exchanger was -24 m, starting with an -5 mm diameter

and exhausting through a -16 mm diameter tube.

The discussion of the facility development has focused on the cryostat

internal construction. Two other important areas of the design are the external

load frame construction and the accommodations for vacuum sealing.

The configuration of the servo -hydraulic test machine differs from the test

machines we normally use for cryogenic mechanical testing. The hydraulic

actuator of the 4.4 MN capacity machine is located at the base of the machine

while the normal servo -hydraulic machine used for cryogenic tests has a

4
crosshead-mounted actuator. Normally, cryogenic tests are done in a closedbottom Dewar and the load frame reacts the load against the upper crosshead.

Using a closed-bottom Dewar necessitated building an external load frame.

The frame consisted of two large reaction plates held apart by four long (3m)

columns on the corners. The lower reaction plate (normally the lower platen of

the test machine) was mounted to the actuator. The four long columns were

mounted vertically on the outside corners on the top of the lower reaction plate.

The upper reaction plate was mounted on top of the four columns. The upper plate

was machined with a clearance hole (170 mm diameter) in the center to allow a

150 mm diameter push rod to pass through. This rod (chromium AISI 304) is a

compressive push rod that rigidly mounts the upper (warm) platen, located inside

the cryostat, to the test machine's crosshead, outside the cryostat. The upper

reaction plate also has a four hole pattern for mounting the four- titanium- post

cage containing the lower platen. When the hydraulic actuator is moved, the

motion of the lower reaction plate is transmitted by the four columns, to move

the upper reaction plate. Since the four titanium posts are also connected to

the upper reaction plate, they move the lower platen in unison with the actuator.

This large double -cage apparatus moves with the actuator while the compressive

push rod and upper platen remain stationary. To apply a load the test specimen

must be in place on top of the lower platen. The specimen is compressed when the

actuator is moved up and the specimen contacts the stationary upper platen.

The next area to be discussed is some of the design details of the vacuum

chamber. Since the cryostat was designed to fit in a closed bottom, open mouth

Dewar all of the vacuum seals were at the top of the Dewar. The Dewar was

constructed with an 0-ring flange top with a mounting bolt pattern. The upper

reaction plate of the test fixture acted as the top of the Dewar that needed to

be sealed. With the cryostat assembled, the Dewar fit inside of the four-column

external load frame

To accommodate the necessary access ports for the internal vacuum chamber,

an adapter plate was machined to fit between the Dewar flange and the upper

reaction plate. Vacuum feed-through ports were needed for the cryogenic plumbing

lines and instrumentation lead wires. The adapter plate was machined from a

canvas -phenolic composite material and was 0-ring sealed when mounted to the

upper reaction plate. The necessary plumbing inlet and exhaust ports were epoxypotted into machined grooves in the composite plate. Pipe- threaded access ports

were machined to accommodate commercial instrumentation vacuum feed-throughs.

A bolt stud pattern was used to seal the Dewar flange to the composite adapter

plate

The compressive push rod (150 mm diameter chromium-plated AISI 304) was

sealed where it entered the vacuum chamber, at the top of the upper reaction

plate, using an 0-ring flange specifically designed for the dynamic motion of the

push rod.

The Dewar was equipped with a vacuum pump port into the inner chamber. A

large (760 1/min) mechanical vacuum pump was used to evacuate the chamber. Lower

vacuum pressure than the pump could attain was assumed possible due to cryopumping from the liquid helium cooled components inside of the test chamber.

5
3 . 2 Instrumentation

This section details the various types of instrumentation used for all

tests. The physical data to be recorded were load, displacement, temperature,

and vacuum pressure. The data were recorded with a computer data acquisition

system and an analog x-y chart recorder. Since all tests were basically

compression tests of a tube, one requirement common to all tests was a record of

load versus displacement. The load was measured using the test machine's load

cell. Other data measurements varied and are described for individual tests.

Three alternative methods of displacement measurement were used: localized

specimen strain measurement using strain gages bonded directly on the specimen,

overall specimen displacement measurement using LVDT's (linear variable

differential transformer) from specimen end-to-end, and as a backup method, the

test machine's stroke displacement was also measured with an LVDT. Strain gages

were bonded directly to specimens 005, 005-1, 006-1, and 008. Each specimen had

four strain gages longitudinally mounted and radially spaced 90° apart. For

short specimens, the gages were located half-way along the length. For long

specimens, the gages were located ~500 mm from the top (warm) end fitting. The

strain gage signals were conditioned with a commercial strain indicator and a

three leadwire, quarter bridge configuration was used. The leadwires were

soldered to the leadwire vacuum feed-through port for the long specimen test.

Two LVDT's were used to measure displacement on specimens 005, 005-1,

006-1, and 009. Overall specimen displacement was measured using two LVDT's,

radially spaced 180° apart, on opposite sides of the specimen. The core of the

LVDT was attached to the upper platen to ensure that it remained warm during

testing. An extension rod (6 mm diameter Invar) was attached to the center rod

of the LVDT. It extended down the length of the specimen and contacted the top

of the lower platen. After the test on specimen 009, one of the LVDT's was

damaged, so for the next test on specimen 007, only one LVDT was used. For the

two short specimens 005-2 and 006-2, without end fittings, displacement data were

obtained from the test machine stroke LVDT.

Temperature data was obtained for all the thermal gradient tests using

type-E (chromel/constantan) thermocouples. The specimen and the test fixture

were instrumented with thermocouples to monitor the thermal gradient. Each

thermocouple bead was securely attached to the specimen or fixture at strategic

locations along the length of the apparatus . The thermocouple leadwires were

soldered to a thermocouple reference junction/vacuum feed-through port. Shielded

copper wire pairs connected each thermocouple reference junction to a 20 channel

digital voltmeter. The thermocouples were calibrated in this configuration. The

thermocouple data were taken directly from the digital voltmeter into the

computer data acquisition system.

The vacuum pressure in the test chamber was monitored continuously with a

thermocouple gage and analog readout unit. The system was capable of reading

pressures ranging from atmospheric down to 0.133 Pa (1 mTorr)

6
4.0 EXPERIMENTAL RESULTS AND DISCUSSION

In this section we present the details of all tests performed on the

composite support struts. Load- versus- displacement records and temperature

gradient data are provided for each test. Finally, a short discussion and

conclusions are given for each group of tests. Some remarks are also given for

the suspected mode of failure for each test.

4 . 1 Ultimate Compressive Strength: Short Specimens

Four short specimens (two, 0.3 m long and two, 0.6 m long) of the full

scale diameter tube were available for testing. They were tested at ambient

temperature to obtain some base properties of the prototype tubes. The two

shorter tubes (specimens 005-2 and 006-2) were tested without any special end

fittings. The ends of the composite tubes were set directly on the compression

platens. The 0.6 m long tubes (specimens 005-1 and 006-1) were equipped with

aluminum (7075 -T6) end fittings. The end fittings, shown in Figure 3, slide onto

the ends of the tube and are designed to prevent brooming of the tube ends. If

end fittings are not used, the tube fails by brooming and the materials maximum

strength is not actually measured. The load- versus- displacement curves for all

four tests are shown in Figure 4.

For the short-specimen tests, the cryostat test fixture was not used. The

specimens were compressed to failure between platens in the 4.4 MN capacity test

machine. A compression swivel platen was attached to one of the test machine's

rigid compression platens to ensure axial loading of the test specimen. The load

was applied using a ramp function in a displacement control mode. Overall

specimen displacement was measured using LVDT's from platen to platen.

Additional strain measurements used strain gages bonded to the specimen. Load

versus displacement curves (LVDT) were plotted on an x-y recorder.

Specimen 005-1

The undamaged portion of specimen 005 from the full scale tests was cut off

at 0.66 m (26 in). The end fitting from the warm end of the tube in the full

scale test was intact and not removed. A new end fitting was machined for use

on the other end of the short specimen. This end fitting was pressed on the

short tube and potted in epoxy under a 445 kN (100 kips) compressive load.

The specimen was tested at 295 K in a 4.4 MN (1000 kips) servohydraulic

test machine. The specimen failed at a maximum compressive load of 1.6 MN (369

kips). The specimen failed in primarily a shell-buckling mode, but a longitudinal crack running the length of the specimen was found in post-mortem

analysis. One end fitting had an exterior surface crack, radially oriented,

about 150 mm (6 in) long.

Specimen 006-1

Full scale specimen 006 failed during the installation of the end fittings

as described later in this report. The damaged ends of specimen 006 were cut off

7
and provided a 0.58 m (23 in) specimen without end fittings. New end fittings

were then installed before testing.

The specimen failed at a maximum compressive load of 1.7 MN (384 kips) in

the 4.4 MN (1000 kips) test machine. The primary failure mode was shell buckling

which was also the failure mode of specimen 005-1. The shell buckling was

preceded by a large longitudinal crack which initiated at the rigid aluminum end

fitting and propagated the length of the specimen.

The aluminum end fitting exhibited premature failure at approximately half

the maximum load. The end fitting failure was characterized by radial cracking

on the exterior surface . The location of the cracking showed that the cracks

initiated at the bottom of the internal groove. The cracks were 75 to 100 mm (3

to 4 in) long in about four places.

Specimen 005-2

An additional 0.31 m (12 in) specimen was available from the undamaged

portion of specimen 005. This short tube was tested with no end fittings to gain

additional data on the material's ultimate compressive strength. The test was

performed under isothermal conditions at room temperature (295 K) . The specimen

failed at a maximum compressive load of 2.4 MN (536 kips). The specimen failed

catastrophically by brooming on one end.

Specimen 006-2

This 0.36 m (14 in) specimen was removed from the undamaged portion of

full-scale specimen 006. Similar to specimen 005-2, this specimen was tested

under isothermal conditions at room temperature (295 K) . The specimen failed at

a maximum compressive load of 2.4 MN (538 kips), nearly identical to the failure

load of specimen 005-2. The mode of failure was again brooming from one end of

the specimen.

Summary of Results

The results of the two tests without end fittings (specimens 005-2 and

006-2) are almost identical and are shown in Table 1. Both specimens failed at

approximately 2.4 MN (540 kips) compressive load and both failed due to brooming

on one end. The average cross-sectional area of the specimens was 64 cm2 (10

in2

) , and the average ultimate strength was 375 MPa (54 ksi)

.

The other two tests (specimens 005-1 and 006-1) failed at loads of 1.6 MN

(369 kips) and 1.7 MN (384 kips) giving an average ultimate strength of 262 MPa

(38 ksi) . The failure mode for these two tests was shell buckling, but both

tubes had longitudinal cracks emanating from one end (see Figure 5) . The end

fittings also exhibited initial signs of failure on both tests. Circumferential

cracks on the exterior surface of the end fitting (relative to the base of the

internal machined groove) were observed (see Figure 6)

.

8
Conclusions

The short specimens without end fittings failed at higher stresses than the

longer specimens with end fittings. It cannot be concluded that end fittings

were detrimental to the tube strength since there was also the variable of

specimen length. The decrease in tube strength is due to the change in mode of

failure from end brooming to shell buckling. The initial signs of failure of the

end fittings indicate a problem area where large stresses are too close to a

geometric stress concentration (the root radius at the base of the machined

groove)

.

4. 2 Ultimate Compressive Strength: Long Specimens

One full scale cryogenic test (specimen 009) is discussed in detail and

another unsuccessful full scale test (specimen 005) is briefly discussed. Both

specimens were full length (~2 m) and were tested in the 4.4 MN cryogenic test

machine. The specimens were equipped with the aluminum end fittings described

previously. The cryogenic test of specimen 009 was a controlled thermal gradient

compression test, using liquid nitrogen (76 K) for cooling. The test of specimen

005 preceded this test and was intended to be an ambient temperature, full scale

compression test as well as a proof test of the cryostat load fixture; however,

this specimen was inadvertently heated during testing, which caused premature

failure

.

Specimen 005

Specimen 005 was shortened to 1.61 m (63.5 in) from its original 1.98 m

(78.0 in) length to properly scale the buckling conditions from cryogenic to room

temperature. Also, a small portion of the tapered end of the tube was removed

because of some visible delaminations. The end fittings used for this ultimate

compression strength test were fabricated from 7075 aluminum and did not use the

lead pad. The end fittings were press-fitted to the specimen, potted in epoxy,

and subjected to a 4.4 kN (100 kips) compressive load during curing of the epoxy.

The test on specimen 005 was a test of the material as well as a learning

exercise on the assembly and operation of the 4.4 MN cryostat. The test set-up

was identical to a full scale cryogenic test without the super- insulated Dewar.

This made it possible to view the specimen during testing. In hopes of visually

observing specimen damage upon loading, two 100 W light fixtures were installed

inside the tube. The lights were turned on when the specimen was initially

loaded, then after a ~30 min instrumentation and fixture proofing procedure, the

test was started. The specimen was ramp loaded and failed prematurely at 445 kN

(100 kips) by column buckling near the midpoint (approximate location of the

lights). Immediately after failure, the buckled specimen was examined and found

to be warm. The external surface of the specimen was approximately ~25°C above

room temperature. Specimen creep was observed as the load was applied, but the

temperature rise was not observed until after the tube failed.

Specimen 006

This was originally to be a full scale room- temperature test with lead-pad

end fittings. The full length specimen was cut to 1.61 m (63.5 in), removing

9
material from the cold end (straight section) to reduce the critical column

buckling stress that is the dominating failure criteria for warm tests.

This test was never run because a longitudinal crack initiated when the end

fittings were pressed on the tube. This failure is probably due to the

interference fit between the inside diameter of the tube and the inside diameter

of the end fitting groove. The larger diameter of the groove in the aluminum end

fitting caused a large diametral strain and hoop stress on the tube that caused

initiation of the longitudinal crack.

Specimen 009

Specimen 009 was a full scale test conducted in a thermal gradient. The

specimen was equipped with rigid end fittings which necessitated the use of a

swivel platen to ensure axial loading of the specimen. Liquid nitrogen (76 K)

was used to cool the lower two-thirds of the test fixture and specimen. The goal

was to test a full length (2 m) tube with an imposed cryogenic thermal gradient

and to verify the cryostat thermal efficiency before liquid helium (4 K) tests.

Some problems, such as cryostat vacuum leaks and specimen thermal intercept

inefficiency, were noted for future repair.

With liquid nitrogen, the lowest temperature attainable for the thermal

gradient is 76 K. The ideal conditions would be isothermal for the lower twothirds of the specimen at 76 K, then a gradient from 76 K to ambient (295 K) over

the upper one -third of specimen length. This ideal gradient was not obtained due

to time constraints and the low thermal conductivity of the specimen. The actual

gradient attained on the specimen is shown schematically in Figure 7 and ranged

from 115 to 278 K.

The specimen was loaded in compression using manual displacement control.

The load was applied in 2 kN (450 lb) increments. At each load step, the test

was interrupted for data acquisition and all transducers were scanned. The

compressive load versus displacement curve is shown in Figure 8. The specimen

failed with a loud acoustic emission at 3.34 MN (752 kips).

Summary of Results

Specimen 009 failed in the mid- region, where the liquid nitrogen thermal

intercept was attached (Figure 9). At this location, several factors could

influence the stress field in the material. The mechanical attachment of the

thermal intercept (Figure 10) could increase the localized stress on the specimen

since it confines diametral strain when the tube is compressed. This is also the

location along the length of the specimen where the specimen changes from a

constant cross-sectional area (see Figure 1) to an increasing outside diameter

and increasing cross-sectional area.

The ultimate compressive strength of specimen 009 was 518 MPa (75 ksi)

.

The temperature in the region of fracture was 140 K. The elastic modulus

calculated from the load-displacement curve was 45 GPa (6.5 Msi)

.

Total strain

to failure was 0.0135.

10
The specimen failed catastrophically by a combination of failure modes:

shell buckling, column buckling, and longitudinal cracking. Column buckling is

indicated by the midlength fracture and the angle at which the specimen came to

rest after the failure event. Shell buckling is inferred from the appearance of

the fracture area: both halves of the specimen are broomed, splayed, and pushed

together (Figure 9). Both end fittings exhibited the initial signs of failure

described for the previous short specimen tests: circumferential cracking on the

exterior surface relative to the base of the internal machined groove (Figure 6) .

Conclusions

The full scale, ultimate compressive strength test on specimen 009 was

successful in characterizing the mechanical performance of the FRP tube subjected

to a thermal gradient. The failure stress (518 MPa) was slightly lower than

expected (620 MPa) . The location of the failure is a potential problem that

needs further design evaluation. It is also important that the specimen failed

in a cold (140 K) region. Further down the specimen (Figure 7), the temperature

was higher (190 K) and the stress was equivalent (518 MPa) . The end fittings

again exhibited cracking which possibly induced premature failure of the FRP

tubes

.

4 . 3 Fatigue Tests

Two extreme thermal gradient tests, using liquid helium cooling, are

discussed in this section. The first (specimen 007) was a cyclic fatigue test

that had multiple failure events before total catastrophic failure at 385 cycles.

The second (specimen 008) was a single load cycle test performed with

approximately the same environmental and mechanical conditions to help analyze

the results of the first test. The two fatigue specimens were fitted with selfaligning end fittings. The difference in design between the two types of end

fittings is shown in Figure 11.

Specimen 007

Since the full scale test of specimen 009 was successful, we were able to

return to the original test plan for the remaining two specimens. The original

test plan called for fatigue tests to be performed with the 4 to 295 K thermal

gradient along the length of the test specimen. The test plan specified that the

tube be cyclically loaded in compression to approximately 310 MPa (45 ksi) at the

warm end for 22 000 cycles. The 310 MPa (45 ksi) stress is based on 1.5 times

the operating design stress of 207 MPa (30 ksi) . The actual fatigue conditions

were cyclic loading to 276 MPa (40 ksi) using a sawtooth waveform with a stress

ratio R ratio of 0.1 and a frequency of 0.25 Hz (Figure 12). The slightly lower

maximum stress was used since the ultimate strength of the material was 20

percent lower than expected.

The end fittings were epoxied into the specimen under a 445 kN (100 kips)

compressive load. The specimen was wrapped with 50 mm (2 in) wide aluminized

tape on the outside diameter to provide a layer of thermal radiation insulation.

The inside of the tube was filled with crumpled, aluminized Mylar film, also for

thermal radiation insulation. With the tube in the fixture, the whole fixture

11
was wrapped with aluminized Mylar sheet, another thermal radiation barrier. The

thermal gradient measured just before testing is shown in Figure 13.

Due to the complexity of this full scale test, the first load cycle was

controlled manually. The load was applied with the machine in displacement

control in approximately 4.4 kN (1000 lb) increments. The load-versus

displacement curve, shown in Figure 14, was linear to 1.2 MN (282 kips), when a

loud noise occurred. The specimen load dropped to 1.0 MN (230 kips), and the

load- versus -displacement curve shifted an amount that indicated 3.5 mm (0.14 in)

of platen- to -platen closing. The specimen was then manually unloaded to 0.76 MN

(170 kips) and had the same linear elastic slope as before the noise. The

specimen was then reloaded to 1.0 MN (230 kips) where the manual incremental step

loading was restarted. The desired maximum fatigue load of 1.8 MN (400 kips) was

reached without incident. The specimen was unloaded to the minimum fatigue load

of 0.18 MN (40 kips) and the test machine was switched to load control in

preparation for the computer controlled fatigue waveform.

The specimen was loaded with a triangular waveform, in load control, at a

frequency of 0.25 Hz as shown in Figure 12. The fatigue cycle continued for 385

cycles until another loud noise occurred. The fatigue cycle was stopped and

again the specimen was still able to sustain a load. The specimen was reloaded

manually from zero load and the load-versus- displacement curve was monitored.

The curve again had the same linear elastic slope but had shifted 5.3 mm (0.21

in) from the initial loading curve (cycle number 1) . This manual reloading cycle

was curiously normal until another loud noise at 1.7 MN (380 kips) signaled total

catastrophic failure of the specimen: no load could be sustained with further

displacement of the actuator.

The specimen failed at the bottom (or the cold end) as shown in Figure 15.

The failure mode appeared to be shell buckling. The failed area covered about

300 mm (1 ft) of length starting 25 mm (1 in) from the bottom end of the

specimen. The specimen also contained numerous longitudinal cracks emanating

from the bottom end of the specimen. Examination of the longitudinal cracks

revealed approximately ten, 300 to 900 mm (2 to 3 ft) long cracks and many short

50 to 150 mm (2 to 6 in) cracks spaced approximately 13 mm (0.5 in) apart.

The longitudinal cracking of the specimen on the cold end effectively

changed the specimen's geometry from a tubular cross-section into a series of

independent columns. The ultimate failure of the specimen was local buckling of

the slender columns.

Both end fittings on the specimen showed signs of failure as shown in

Figure 16. The lower (cold) end fitting completely fractured into multiple

pieces as shown in Figure 15. The band of material on the outside of the

machined groove in the lower end fitting fractured at the base of the groove and

was no longer in place to function as designed. This may be the reason for the

initiation of longitudinal cracks in the specimen. The upper (warm) end fitting

had radial cracks visible on the outside surface but had not completely fractured

as in the lower end fitting. Another problem for both end fittings was the area

machined for placement of the lead pad. Both end fittings fractured between a

fillet in this machined area and the free surface. The fracture path was

slightly different for each end fitting as shown in Figures 15 and 16.

12
Specimen 008

Due to the uncertainty of the dominating failure mode on the previous test,

we performed an interrupted test on the final specimen. The plan was to

terminate the test after the first sign of failure (a loud noise or anomalous

load and

strain behavior) . The thermal conditions of the previous test were duplicated

as closely as possible.

The full length (2 m) specimen was fit with the self -aligning end fittings.

The cold end fitting fit the specimen well and was not epoxied on. The warm end

fitting fit well on the inside diameter of the tube, but had excessive clearance

between the outside diameter and the groove. There was an approximately 1.3 mm

(0.05 in) gap around the outside of the tube that we filled with fiber-glass

cloth and epoxy. It was necessary to fill this void to give the specimen side

support on the end. The epoxy cured with the specimen under a 44 kN (10 kips)

compressive load. This specimen was not wrapped with aluminized tape, but the

test fixture was skirted with aluminized Mylar sheet for thermal radiation

insulation. The inside of the tube was filled with crumpled aluminized Mylar

sheet also for thermal radiation insulation.

The specimen was placed in the test fixture at ambient temperature. The

fixture was sealed and the vacuum roughing pump and liquid nitrogen cooling were

started simultaneously. Liquid nitrogen was used to cool the three cooling coils

for approximately 75 h at an average of 19 1/h. The rough vacuum during this

interval was ~27 Pa (-200 mTorr)

.

As in previous tests the specimen temperature was higher than the fixture

temperature. After 75 h of phase I (liquid nitrogen only) cooling the lower

platen of the test fixture was 85 K, the lower end fitting was 91 K and 25 mm

(1 in) above the lower end fitting the temperature of the specimen was 137 K.

At this point, phase II of the cooling was started. Two 500-1 Dewars of liquid

helium were connected to the two 4 K cooling circuits. Approximately 10 to

15 min after the start of liquid helium flow, the test chamber vacuum pressure

dropped from 27 Pa (200 mTorr) to 1.3 Pa (10 mTorr). During phase II liquid

nitrogen flow was maintained in the 76 K thermal intercept.

The liquid helium/liquid nitrogen cooling of the test apparatus was

conducted for approximately 12 h before the start of the compressive test. The

respective flow rates for the liquid helium/liquid nitrogen cooling were 48 1/h

and 8 1/h.

After 12 h in phase II cooling, the lower platen and the lower end fitting

temperatures were both 35 K while 25 mm (1 in) above the end fitting, the

specimen temperature was 83 K. At the intermediate liquid nitrogen thermal

intercept the specimen temperature was 100 K. The thermal gradient measured just

before testing is shown schematically in Figure 13.

The specimen was loaded manually in displacement control. The loading was

suspended every 8 . 9 kN (2 kips) and all transducers were scanned. The resulting

load-versus- displacement curve, shown in Figure 14, was linear up to point A,

where a loud acoustic emission was heard. The load at this point was 1.16 MN

13
(260 kips) and it dropped instantly to 0.93 MN (210 kips). The load-versusdisplacement curve shifted to the right by an amount that indicated approximately

2 mm (0.08 in) of platen- to-platen closing. This load drop, curve shift, and

acoustic emission indicated some type of failure so we stopped the test to

inspect the specimen. The specimen unloading curve was linear and parallel to

the original loading curve

.

The test specimen remained intact but the lower (cold) end fitting was

completely fractured into multiple pieces. The failure of the lower end fitting

was essentially identical to the failure of the lower end fitting on specimen

007. The band of material on the outside of the machined groove completely

fractured at the base of the groove. There was also complete fracture of the end

fitting, in the area recessed for the lead pad, identical to previous failures

of the lower end fitting.

The specimen exhibited the start of numerous longitudinal cracks emanating

from the bottom of the specimen. The 50 to 100 mm (2 to 4 in) long cracks were

spaced approximately every 13 mm (0.5 in) about the circumference of the

specimen.

The load at which this end fitting failed was within 8 percent of the load

at the first failure event in the previous test. This coincidence of load and

the similarity of the two failed end fittings leads us to conclude that the first

failure event in the previous test also occurred in the lower (cold) end fitting.

Summary of Results

Examination of specimen 007 indicated that the specimen and both end

fittings failed. The specimen failed at the cold end as shown in Figures 15 and

16. The end fitting at the cold end had fractured into multiple pieces (Figure

16A) while the warm end fitting fractured into two pieces (Figure 16B)

.

The single cycle compressive test of specimen 008 caused the cold end

fitting to fail. The failure of the end fitting into multiple pieces was

essentially identical to the previous failure of the cold end fitting on specimen

007.

Conclusions

The most important conclusion from these tests is that the compressive

self -aligning end fittings must be redesigned to develop the full strength of the

composite tube. Previous tests performed using the rigid aluminum end fittings

alluded to a strength problem. The self -aligning end fittings contain more

geometric stress concentrations than the rigid end fitting. This combined with

the more brittle behavior of the 7075 -T6 aluminum alloy at cold temperatures was

detrimental to end fitting performance. Both tests failed on the first load

cycle at almost identical stresses (186 ±8.5 MPa). For specimen 007, it was not

possible to verify the cold end fitting failure as the first failure but that

conclusion was confirmed by the test of specimen 008.

Since the failure of the cold end fittings prohibited further determination

of the composite tube properties, the only other conclusion from the test on

14
specimen 007 is that the fatigue life of the tube for the documented thermal and

mechanical conditions is greater than 385 cycles for the loading conditions

described above.

5.0 SUMMARY AND CONCLUSIONS

The results of all tests of the FRP support tubes are summarized in

Table 1. For each test, the thermal conditions, ultimate load, and failure mode

are indicated. We conclude the following:

(1) The failure stress on the full scale, ultimate compressive strength

specimen (specimen 009) was nearly 20 percent lower than expected.

(2) The end fittings for the tubes present the greatest obstacle to

loading the tubes to their ultimate strength.

(3) The stress concentration of the end fittings is a problem that is

magnified with the present design of the self -aligning end fittings.

15
Table

1:

Test

result

summary

sheet. I

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<DO

exC

Bh

•oeW*

•oav

4—>

eo

•Oeoc T3e

<u

4—> co

6X3 C1oo5— XJ

<N

(3)

self-aligning

end

fittings

(3)

column

buckling
23.7 cm

26.7 cm

Start of taper

(increasing outside

diameter for upper

0.67 m)

.

Figure 1. Schematic of composite tube.

17
18

Figure

2.

4.45-MN

(1000

kips)

test

machine

and

cryostat.
19

Figure

3.

Rigid

end

fitting

(7075

T6

Al)
UM ‘avoi N0ISS3adlA|00

o o o o o

o o o o o

LO LO o LO

OvJ C\J T— T—

20

DISPLACEMENT,

mm
Figure 5. A failed short specimen with end fittings.

21
22

Figure

6.A

circumferential

crack

on

surface

of

end

fitting.
TEMPERATURE

(K)

23

Figure

7.

Schematic

for

thermal

gradient

of

specimen

009.
SPECIMEN

#009

O

CO

24

Figure

8.

Load-versus-displacement

curve

for

specimen

009.
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fo**HMMMiii iim ii M^

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Figure 9. Specimen 009 after ultimate strength test.

25
26

Figure

10.

Liquid

nitrogen

thermal

intercept

attachment.
Figure 11A. Rigid aluminum end fitting (7075 Al).

r 5 Aluminum

I

Figure 11B. Self-aligning end fitting.
300

28

Figure

12.

Schematic

of

fatigue

test

parameters

for

specimen

007.
TEMPERATURE

(K)

coo0

Z

IU

1

o

UJ

Q.

Cf)

in in CO

CD CM

CM CM Tm

CO

CM t- w co oo co co

29

Figure

13.

Schematic

for

thermal

gradient

of

specimens

007

and

008.
Figure 14.

v

Load-versus-displacement curves for specimens 007 and 008.

30
31

Figure

15.

Fatigue

specimen

007,

cold

end

failure.
Figure 16B. A failed self-aligning end fitting, warm end of specimen 007.

Figure 16A. A failed self-aligning end fitting, cold end of specimen 007.

32
BL-114A U.S. DEPARTMENT OF COMMERCE

(5-90) NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY

ADMAN 15.01

BIBLIOGRAPHIC DATA SHEET

1. PUBLICATION OR REPORT NUMBER

NISTIR 3966

2. ORGANIZATION REPORT NUMBER

3

- ]$Mai$y?ATE

4. TITLE AND SUBTITLE

A Cryogenic 4.4 MN Mechanical Test System for Large Scale Tests of Composite Support Stmts

R.P. Walsi, R.P. Reed, J.D. McColskey, W. Fehringer, J.R. Berger

6. PERFORMING ORGANIZATION (IF JOINT OR OTHER THAN NIST, SEE INSTRUCTIONS)

U.S. DEPARTMENT OF COMMERCE

NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY

BOULDER, COLORADO 80303-3328

7. CONTRACT/GRANT NUMBER

8. TYPE OF REPORT AND PERIOD COVERED

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (STREET, CITY, STATE, ZIP)

10.

SUPPLEMENTARY NOTES

1 1

ABSTRACT (A 200-WORD OR LESS FACTUAL SUMMARY OF MOST SIGNIFICANT INFORMATION. IF DOCUMENT INCLUDES A SIGNIFICANT BIBLIOGRAPHY OR

UTERATURE SURVEY, MENTION IT HERE.)

This report details a large mechanical test program undertaken by the NIST Materials Reliability Division. These tests were

intended to validate the full scale structural behavior of the composite support struts in a thermal and mechanical environment

identical to actual service conditions.

Two types of tests were performed: ultimate compressive strength and fatigue. Thermal conditions for the tests ranged from

isothermal, room temperature tests to an imposed thermal gradient of cryogenic-to-room temperatures. The development of

12.the unique facility necessary for the large scale cryogenic tests is detailed in this report.

KEY WORDS (6 TO 12 ENTRIES; ALPHABETICAL ORDER; CAPITALIZE ONLY PROPER NAMES; AND SEPARATE KEY WORDS BY SEMICOLONS)

composite; compressive platen; cryogenic; fiber-reinforced polymer; superconducting magnet; vacuum pump

13. AVAILABILITY 14. NUMBER OF PRINTED PAGES

X

UNLIMITED 36

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