SPACE SHUTTLE FILAMENT WOUND CASESOLID ROCKET MOTOR STATIC TEST RESULTS (DM-6)

C. A. SADERHOLM

MORTON THIOKOL, INC./WASATCH DIVISION

BRIGHAM CITY, UTAH

ABSTRACT

The filament wound case solid rocket motor (FWC-SRM) is presently under development toincrease the payload capability of Shuttle launches from the Vandenberg Air Force Base (VAFB) by4 ,600 pounds . The fi rs t o f th ree s ta t i c tes ts (DM-6) to qua l i f y th is des ign fo r fl igh t wassuccessfully conducted in October 1984.

The steady state and dynamic thrust performance of DM-6 was similar to that of the currentfl ight s tee l case SRM. In addi t ion, there was l i t t le d i f ference between the thrust t race for DM-6and that for the steel case design during ignit ion and tai loff. This meant no change would benecessary in the SRM thrust loads used for Shuttle ascent loads analyses.

Overal l structural performance of the FWC was excel lent. The axial and radial growth duringoperation met or exceeded design requirements. Bending stiffness was 20 percent less thanpredicted requiring some adjustment in the FWC dynamic model used for Shuttle dynamic analyses.

Performance of the tape wrapped nozzle inlet ablative rings with the new ply angles and lowsodium carbon c loth mater ia l was super ior to the r ings in the current fl ight nozzles. Nozzleinlet parts for the two remaining FWC-SRM static tests wil l be fabricated with the new ply anglepr ior to making any decision to incorporate the change in the flight nozzle parts.

INTRODUCTION

In May 1982 work was initiated to develop a filament wound case (FWC) for use on the SpaceShutt le sol id rocket motor (SRM) as a replacement for the l ightweight steel case for specialmissions requiring increased payload capability. The FWC-SRM was designed, developed, and is nowbeing manufactured by Morton Thiokol, Inc. at its Wasatch Division in Utah under a contract managedby the Marshall Space Flight Center of NASA. The FWC design was developed and manufactured by theHercules-Votaw Joint Venture under subcontract to Morton Thiokol , Inc. At the incept ion of theFWC development program, the maximum inert weight for the FWC segments was 65,000 lbm. Duringdevelopment, s ignificant design problems resul ted in inert weight increases total ing about 7,000lbm. As a consequence, about a year after initiating the development program the maximum inertweight was increased to 72,000 lbm.

In June 1985, three years after development was initiated, loaded FWC-SRM segments for thefirst fl ight wi l l be avai lab le for del ivery to Vandenberg Ai r Force Base (VAFB) for launch inearly 1986. The inert weight of these flight motors wil l be less than the maximum requirementthus providing a weight reduction of about 25,000 lbm compared to the lightweight steel case.This is expected to increase the Shutt le l i f t capabi l i ty by 4,600 lbm.

DISCUSSION

When the filament wound case SRM was selected as a performance improvement option for theSpace Shutt le i t was necessary to establ ish bal l is t ic performance and structural design requirements that would al low it to be used interchangeably with the steel case SRM without significantlya f fec t ing the fl igh t per fo rmance o r s t ruc tu ra l behav io r o f the Shut t le veh ic le . To p rov ideequivalent bal l ist ic performance i t was establ ished that the internal dimensions of the FWC wouldbe the same as those of the steel case. To provide acceptable structural propert ies i t wasnecessary to establ ish requirements that contro l led axia l and radia l growth and axia l s t i f fness ofthe FWC-SRM during motor operation. The static test of DM-6 offered the first opportunity todetermine if the full-scale FWC-SRM would meet these design requirements and provide the confidenceto proceed with the manufacture of the FWC segments for the first flight motors.

Work completed under contract NAS 8-30490 with Marshall Space Flight Center. Release to governmentagencies and their contractors; a l l other inquir ies referred to MSFC.

The FWC-SRM design combines three case components from the segmented steel case (forward andaft domes and External Tank attach) with four graphite filament wound cylindrical segments to formfour casting segments as shown in Fig. 1. Steel adapter rings are attached to the ends of thefilament wound cyl inders by a double row of pins that provide either a tang or clevis feature thatpermits the filament wound cylinders to be attached to each other or to the steel segments duringmotor assembly. The baseline FWC-SRM design configuration is shown in Fig. 2. An i l lustration ofthe typical composite to composite joint construction is shown in Fig. 3 and the construction of atypical composite segment is shown in Fig. 4.

The first FWC segment for DM-6 was delivered to the Morton Thiokol Wasatch Divison in March1984. Prior to receipt of the FWC segments extreme care was taken to clean and abrade the insidesurfaces of the segments to assure a sound and reliable bond would be made between the caseinterna l insu la t ion and the FWC in ter ior dur ing the vu lcanizat ion process. Th is is ext remelyimportant because, unlike a metallic case, a filament wound composite material is porous and mustbe sealed on the inside to prevent any gas leak during motor operation. At the Wasatch Divisionthe FWC components were assembled into casting segments and processed just like the steel casecasting segments.

By early September 1984, DM-6 was assembled in the test stand with the support chocks removedand a l l jo in t leak tes ts complete . Af ter ins ta l l ing a record number o f ins t ruments prov id ing atotal of 401 digital and 81 FM channels of data and completing all pretest checkouts, DM-6 wassuccessfully tested on 25 October 1984.

D6AC STEELFORWARDDOME

GRAPHITE/EPOXY COMPOSITEFORWARD SEGMENT CASE CYLINDER

IGNITERD6AC STEEL TANG AND CLEVISADAPTER MP35N PINS (JOINT)INCONEL 718 PINS(ATTACHMENT)

GRAPHITE/EPOXYCOMPOSITE CENTERCASE CYLINDERSEGMENT

GRAPHITE/EPOXY COMPOSITECENTER CASE CYLINDER SEGMENT

D6AC STEEL ETATTACH SEGMENT

GRAPHITE/EPOXYCOMPOSITE AFTSEGMENT CASECYLINDER WITHINTEGRAL STIFFENERRING

HPM NOZZLE/EXIT CONE

D6AC STEEL AFT DOME

Figure 1. FWC-SRM Configuration

Figure 2. FWC-SRM Design Configuration

GRAPHITE COMPOSITESHELL -\

JOINT PIN (MP 35N) ANDRETAINER BAND -,

CLEVIS -,R I N G /

COMPOSITEPIN/SEALGROUP(132 PINS/ROW)

COMPOSITESEAL

ASSEMBLYSEAL

CAPTUREFEATURE

Figure 3. Typical Composite-to-Composite Joint Construction

MEMBRANE

-OVERWRAP (2 PLACES)HOOP WOUND GLASS

JOINT—Iregion I

(2 PLACES)

SUBSTRATE (2 PLACES) 1.346 in. FWD (1.376 in. OVER 202 in. OF FWD SEGMENT)GRAPHITE CLOTH

JOINTCONSTRUCTION

1.322 in. CENTERS1.309 in. AFT

SEGMENT

12 PLIES 90 deg HOOP

22 HELICAL LAYERS 33.5 deg

48 LAYERS UNIDIRECTIONALBROADGOODS

GLASS 90 deg HOOPOVERWRAP AND GRAPHITECLOTH SUBSTRATE

MEMBRANECONSTRUCTION

19-22 90 deg HOOP PLIES

22 HELICAL LAYERS 33.5 deg3 CUT DOUBLE HELICALLAYERS

GRAPHITE CLOTH SUBSTRATE

-TAG END ——I(TEST SPECIMENS)

MATERIALAS-4 GRAPHITEAND 55A EPOXYRESIN

AS-4 GRAPHITE3501 EPOXY RESINGLASS ROVING55A EPOXY RESIN

Figure 4. Typical Composite Segment Construct

This d iscussion of the DM-6 test wi l l cover the fo l lowing four subjects re lat ive to the testob jec t i ves cons idered to be o f par t i cu la r in te res t : ba l l i s t i c per fo rmance, ign i te r per fo rmance,case performance, and nozzle performance.

BALLISTIC PERFORMANCE

The DM-6 test motor contained 1,105,486 lbm of propellant and weighed a total of 1,222,560lbm. The propellant core design and inhibitor patterns were the same as those used on the steelcase High Performance Motor (HPM) design. Minor differences did exist in the burn out pattern dueto a sl ightly smaller inside case diameter and an insulation change at the segment joints to covera capture feature used to restrain the inner leg of the clevis when pressurized. The last steelcase HPM (QM-4) was tested in March 1983. Performance summaries for DM-6 and QM-4 are presented onTable I. The specific impulse for DM-6 was 0.3 percent higher than that for QM-4 but very closeto that del ivered by the first stat ic test HPM (DM-5).

The propellant formulation and raw material vendors for DM-6 and QM-4 were the same excep_for the burning rate catalyst, iron oxide, which was increased to 0.366 percent to achieve thehigher burn rate required to compensate for the lower ambient temperatures at VAFB. The delivered

Table I. DM-6 and QM-4 Performance Summary (Vacuum, 60°F)

DM-6 QM-4ACTION TIME IMPULSE (Ibf-sec)

ACTION TIME lsp (Ibf-sec/lbm)

ACTION TIME (sec)

BURN TIME (sec)

BURN TIME, AVERAGE VACUUM THRUST (Ibf)

ACTION TIME, AVERAGE VACUUM THRUST (Ibf)

BURN TIME, AVERAGE STAGNATION PRESSURE (psia)

ACTION TIME, AVERAGE STAGNATION PRESSURE (psia)

MOP (psia)

MOF (Ibf)

BURN RATE (ips)C)

AVERAGE NOZZLE EXPANSION RATIO

296 .62X106 297.53 X 106

268.32 267.53

117.90 121.28

107.62 110.93

2,703,000 2,617,000

2,516,000 2,453,000

649.0 645.4

604.5 604.7

937.1 927.8

3,481,000 3,329,000

0.378 0.369

7.44 7.48(1)625 psia AND 60°F

burn rate for DM-6 was calculated to be 0.378 in./sec at a reference pressure of 625 psia and 60°For about one percent above the target of 0.374 in/sec. This difference was within the experienceband from previous stat ic and flight tests as shown in Fig. 5. The ful l-scale HPM burn ratepredict ions are made by mul t ip ly ing the propel lant burn rate measured in 5 in. d ia bal l is t ic testmotors by a scale factor determined from prior ful l-scale HPM tests. Table II presents a comparison of burn rate scale factors determined for the 17 HPM static and flight tests conducted upto and including DM-6 using both 5 in. test motor and acoustic strand burn rate measurements. Notethat the mean values for the two burn rate measuring techniques are different, but both havecoeffic ien ts o f var ia t ion o f about 0 .5 percen t .

DM-5 QM-4 STS-8n 1 rSTS-9 STS-11 STS-13

t i rSTS-14 STS-17 DM-6

i rSTS-19 STS-20

PLOTTED IN CHRONOLOGICAL ORDER

Figure 5. Difference Between Actual and Target Burn Rates for 17 Staticand Flight HPM Tests

Ta b l e I I . Burn Rate Scale Factor Comparison Ratio of Full-Scale to Subscale andStrand Burn Rates

SUBSCALE*1) ACSBR<2>

NO. OF FULL-SCALE TESTS 17 17

RANGE 1.0095-1.0242 0.9943-1 .0130

MEAN 1.0138 1.0033

COEFFICIENT VARIATION (%) 0.496 0.543<1)5 in. DIA BALLISTIC TEST MOTOR (-7.7 lb)(^ACOUSTIC STRAND BURN RATE (CUT CURED STRAND 1/4 in. x 1/4 in.)

Plots of the measured headend pressure data for DM-6 and QM-4 during the ignition transientare compared on Fig. 6. Both pressure traces exhibi ted the same character ist ics dur ing igni t ionwith the DM-6 trace lagging the QM-4 trace by about 5 milliseconds. The maximum pressure riserate in DM-6 was lower (86.6 psi/10 msec) than for QM-4 (90.8 psi/10 msec) but within the experienceband from previous tests.

800

* m + / /n / /g. 600 / /—• ' /UJ / /ccD , y</> <yS

/ /cc 400 I fQ. I f11////

200 t lt l

DM-6

0QM-4

0 . 0 0 0 . 0 5 0 . 1 0 0 . 1 5 0 . 2 0 0 . 2 5 0 . 3 0 0 . 3 5 0 . 4 0 0 . 4 5 0 . 5 0

TIME (sec)

Figure 6. Comparison of DM-6 and QM-4 Headend Pressure During Ignition

A comparison of the DM-6 and QM-4 pressure and thrust traces shows only minor differenceswhen corrected to the same reference burn rate and temperature, as shown on Figs. 7 and 8. Themost noticeable differences occurred during the first 20 sec of operation and several secondsbefore web burnout . The d i f fe rence dur ing the firs t 20 sec is a t t r ibu ted to greater rad ia lexpansion of the FWC. The different trace shape at burnout is probably due to the case insulationchanges made in the DM-6 FWC. As a further measure of the similarity of the performance deliveredby the two motors a comparison of vacuum thrust to headend pressure ratio for both DM-6 and QM-4 isshown on Fig. 9. The maximum difference between the two plots is less than one percent which iswell within the normal variation expected due to nozzle throat erosion differences and measurementaccuracy.

An area of special interest to those in the Shuttle vehicle loads community was the pressureand thrust oscil lations generated during the DM-6 test compared with those developed during thesteel case HPM tests (DM-5 and QM-4). The frequency range of primary interest was 5 to 20 Hzwhich spans the first longitudinal acoustic mode (1-L) in the motor cavity varying about 14 to 15

0 1 0 2 0 3 0 4 0 5 0 6 0 7 0 8 0 3 0 1 0 0 1 1 0 1 2 0 1 3 0TIME (sec)

♦CORRECTED TO 0.368 ips BURN RATE AND 60°P

T T -4 0 5 0 6 0 7 0 8 0 3 0 1 0 0 1 1 0 1 2 0 1 3 0

TIME (sec)CORRECTED TO 0.368 ips BURN RATE AND 60°F

Figure 7. QM-4 and DM-6 Pressure Data Comparison* Figure 8. QM-4 and DM-6 Thrust Data Comparison*

4,250 -,

t -UJCC 4,000 -

C/> 33 COCC COX UJ 3,750 -2DD

QZUJ 3,500 -

O Q2 <IIIX 3,250 -

3,000 t i i 1 1 1 r

0 1 0 2 0 3 0 4 0 5 0 6 0 7 0 8 0 9 0 1 0 0 1 1 0TIME (sec)

Figure 9. Comparison of Thrust to Pressure Ratio Data Fro QM-4 and DM-6

Hz. Activity at the second longitudinal acoustic mode (2-L), which ranges around 29 to 30 Hz,was only of minor interest. To provide a qual i tat ive comparison of the pressure and thrustoscillations measured during the FWC and steel case static tests, data from the dynamic pressuregage (P006) and the F001 axial thrust measurement were used to generate the waterfall plots shownon Fig. 10. Table III presents the maximum pressure and thrust values in the 1-L and 2-L frequencyranges based on the waterfall plots for DM-6 and QM-4. The method used to generate the plotsaverages the measurements over 4 sec intervals so the magnitude of transient oscil lations isa t tenuated. Compar ison o f the pressure osc i l la t ion data ind icates that the overa l l acoust iccharacterist ics of the two motors are very similar to the osci l lat ions in DM-6, being somewhatlarger in magnitude but less organized than those in QM-4. The frequency response of the thrustoscil lations at the 1-L and 2-L modes is also similar in both motors with the magnitude of theoscillations in QM-4 being more pronounced around the 1-L mode and in DM-6 more pronounced aroundthe 2-L mode. There was concern before testing DM-6 that the lighter mass FWC would enhancecoupling between the 1-L acoustic mode and a test stand dynamic frequency near the end of motoroperation. This coupling did not occur to any greater degree in DM-6 than in any previous test.

DYNAMIC PRESSURE (P006)

1 5 2 0 2 5FREQUENCY (Hz)

1 0 1 5 2 0 2 5 3 0 3 5 4 0FREQUENCY (Hz)

D M - 6 Q M - 4DYNAMIC THRUST (CALCULATED FROM F001 DATA)

1 0 1 5 2 0 2 5 3 0 3 5 4 0FREQUENCY (Hz)

1 0 1 5 2 0 2 5 3 0 3 5 4 0FREQUENCY (Hz)

DM-6 QM-4Figure 10. Waterfal l Plots Comparing Pressure and Thrust Osci l lat ions

in DM-6 and QM-4

Table III. Maximum Pressure and Thrust Oscillations for DM-6 and QM-4from Water fa l l P lots

PRESSUREOSCILLATIONS

THRUSTOSCILLATIONS

FREQUENCY TIME MAX VALUEMOTOR MODE (Hz) (sec) (0-TO-PEAK)

DM-6 1-L 15.5 76 0.51 psi2-L 29 86 0.78 psi

QM-4 1-L 14 93 0.21 psi2-L 29.5 80 0.41 psi

DM-6 1-L 15 100 20,900 Ibf2-L 29.5 76 40,530 Ibf

QM-4 1-L 15 91 25,900 Ibf2-L 29.5 79 12,000 Ibf

In addition there was concern about feedback between the thrust and pressure oscil lations nearmotor burnout resu l t ing in an ampl ifica t ion o f the pressure osc i l la t ions . Th is concern a lso d idnot mater ia l i ze resu l t ing in the overa l l conc lus ion tha t the acous t ic and s t ruc tu ra l in te rac t ion o fthe FWC during the DM-6 stat ic test was not significantly di fferent from that already experienceddur ing s tee l case tests .

IGNITER PERFORMANCE

The igniter for the FWC-SRM is the same as that used for the steel case SRM. The main ignitergrain is a 30 point design containing about 137 lbm of TP-H1178 propellant. The maximum mass flowrate of the DM-6 igniter was 358 lbm/sec and the maximum pressure at the test temperature of 76°Fwas 1,989 psia. The igniter had been cast 64 months and stored as part of an ageing program beforei t was fired in DM-6. I t represented the oldest igni ter tested to date in the SRM program. Thepressure traces for the DM-6 igniter and four other igniters cast from the same propellant mix areplotted on Fig. 11. The DM-6 performance was found to be within the variation determined fromother igniters tested that had been cast from the same mix. Three of the pressure traces plottedon Fig. 11 are from open air tests which allowed the igniter chamber to return to ambient pressureafter burnout. I t was concluded from the DM-6 test that there was no significant change in thebal l is t ic performance of the igni ter af ter more than 5 years of storage.

0 . 4 0 . 5TIME (sec)

"CORRECTED TO 80°F

Figure 11. SRM Igniter Pressure Traces-Age Life Evaluation Tests

CASE PERFORMANCE

The overall performance of the DM-6 FWC was excellent. There was no apparent growth of anydelaminations in the composite cyl inders known to be present before the static test. The measuredand predicted values for axial and radial growth as well as bending stiffness are compared withtheir requirements in separate tables as referenced below:

Axial Growth - Well within the requirement as shown on Table IV.

Radial Growth - A requirement existed only for the forward membrane which was well within therequirement as shown on Table V.

Bending Stiffness - 20 percent lower than predicted due to joint interaction as shownin Table VI .

Table IV. DM-6 Case Axial Growth

• REQUIREMENT: 0.6-in. NOMINAL AT 1,004 psi (0.56 in. AT 935 psi)

TOTAL AXIAL GROWTH (A)

TOTAL AXIAL GROWTH(B + SC)

TOTAL AXIAL GROWTH (SD)

FWD/FWD-CTR JOINT

FWD-CTR/AFT-CTR JOINT

FWD-CTR MEMBRANE

AFT-CTR MEMBRANE

MEASURED AT 935 psi PREDICTED AT 935 psi(in.) (in.)

0.34 0.41

0.36 0.41

0.30 0.41

0.10 0.10

0.09 0.09

-0 .085 - 0 . 0 5 3

- 0 . 0 7 8 - 0 . 0 5 3

STA

FWD

(D) (D)

I601

I771

FWD-CTR

I931

AFT-CTR

(C)

AFT

( C ) ( C ) ( C )

( D ) ( D )+ • —

1.251

I I I I1.530 1,560 1.647 1.687

Table V. DM-6 Case Radial Growth

• REQUIREMENT: 0.66 in. MAXIMUM AT 1,004 psi (0.61 in. AT 935 psi)

MEASUREDHOOP STRAIN

(jLcin./in)

RADIAL CONSTRAINT REGIONSFORWARD MEMBRANE

PREDICTEDHOOP STRAIN

(/Jn./in)

MEASUREDRADIAL GROWTH

(in.)

STA 630 AT 7 sec 7,700 7,595 0.57

STA 710 AT 7 sec 7,550 7,445 0.56

STA 630 AT 60 sec 5,470 5,260 -

STA 710 AT 60 sec 5,500 5,260 -

NOTE: ERROR IN PREDICTION DUE TO NONLINEAR HOOP STRAIN RESPONSE.MATERIAL PROPERTIES CALIBRATED TO PREDICT RESPONSE AT PROOFPRESSURE

Table VI. DM-6 Case Bending Stiffness

• REQUIREMENT: NOMINAL Ezt = 10.6 x 1Q6 lb/in.* (66% OF STEEL CASE STIFFNESS)

TOTAL BEAM

TOTAL SAG

SAG VS TIME SLOPE

BEAM BENDING FREQUENCY• FWC-SRM AXIAL STIFFNESS IS 20% LESS THAN PREDICTED

• FWC JOINT BEARING COMPLIANCE IS CAUSE

Et COMPARED COMPARED TOTO PREDICTION STEEL CASE

(%) (%)

- 2 0 54

- 2 0 54

- 2 0 62

♦INTEGRATED Ezt FOR FWC IS 10.4 X 10© lb/in., BUT IS NOT DIRECTMEASUREMENT OF STIFFNESS-IT NEGLECTS JOINT BEARING COMPLIANCE

Al l jo int seals funct ioned normal ly. Plot ted on Fig. 12 is the seal pressure between theprimary and secondary O-rings measured at the FWC center field joint compared with a similarmeasurement made at the center field joint of the QM-4 steel case. Also plotted is the chamberpressure to i l lustrate the response of the joint seal pressure to the changes in chamber pressure.The difference between the FWC and steel case seal pressures is due to the volume changes in thecav i ty resu l t ing f rom d i f fe ren t re la t i ve mot ions o f the tang and c lev is por t ions o f the jo in tduring motor operat ion.

Tabulated on Table VII is another comparison between the performance of the FWC (DM-6) and asteel case i l lustrating the greater deflection or sag of the FWC compared to the steel case.

In summary there was some discrepancy between predicted and measured membrane strains. Axialstrains were up to 60 percent less than predicted due to the difference in small value measurements, and hoop strains were up to 7 percent greater than predicted apparently due to the nonlinear

FWC Steel Case18.0

16.0 FWC (DM-6)—-\Seal Pressure \ /^

14.0\ i12.0 r

11 0 . 0 1 1Q. - - — I2 8 . 0 V /3 ^ " ^ ^ /2 6 . 00. Chamber-'

* - - - - - ^ ^/ /

Pressure > ^4.0 (PX 1,000)s \

2.0S *

s\ \

0.0^ ^

^^Steel Case (QM-4)

- 2 . 0 Seal Pressure

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140Time (sec)

Figure 12. Field Joint Seal Pressure Comparison FWC vs Steel Case

Table VII. DM-6 Case Performance Comparison to Steel Case (QM-4)

SAG AFTER CHOCK REMOVAL (in.)

SAG AFTER FIRING (in.)

SLOPE OF SAG VS TIME (in/sec)

BEAM BENDING FREQUENCY (Hz)

TOTAL AXIAL GROWTH (in.)

QM 4 DM 6

3.37 6 . 0

0.324 0 . 5

0.0286 0.0525

1.84 1.45

0.791 0.34

response of the composite material. This can be seen in Fig. 13 where the hoop strain measurementmade on a steel case motor is compared with the hoop strain measured on a DM-6 composite membraneat station 805. Both measurements are plotted as a function of the headend chamber pressure and ascan be seen, the variat ion in strain in the steel case is l inear after about 20 sec whi le thevariat ion in strain in the composite membrane is nonl inear. As a resul t , the strain measurementson the exter ior of the composi te mater ia l cannot be di rect ly re lated to the internal pressure.Additional analyses are being conducted to develop a method to correct the measured strains forthe pressure gradient through the propellant and composite material due to the radial growth ofthe composite membrane.

8,000

7,000

6.000

5,000

.£ 4,000Zk 3,000wzcc 2,0005

1,000

-1,000.100

STATION 805 - FORWARD CYLINDER MEMBRANE

DIVERGENCE FROM /

*

LINE)\R RESPONSE "f y S t = 20

sa /

DM-eFWC-SRM t 20 sec/ p r e

' B U ISSURE/LDUP

rHRUST

f t

^ S T EEEL CAS EHPM

100 2 0 0 3 0 0 4 0 0 5 0 0 6 0 0HEADEND PRESSURE (psig)

7 0 0 8 0 0 9 0 0 1 , 0 0 0

Figure 13. Comparison of Hoop Strain vs Headend Pressure for DM-6FWC and Steel Case HPM

NOZZLE PERFORMANCE

The DM-6 nozzle was fabricated to production nozzle design requirements with the followingmodificat ions to be eva lua ted fo r the fi rs t t ime dur ing a fu l l -sca le nozz le tes t . The two tapewrapped nozzle inlet ablat ive r ings, the forward nose and aft in let , incorporate new ply anglesand low sodium (150 ppm) carbon cloth material; a new carbon black filler was used in the forwardand aft carbon phenolic exit cone liners; and changes were made in the nozzle exit cone severancesystem. The DM-6 nozzle differences from the high performance motor (HPM) production nozzle arei l l u s t r a t e d i n F i g . 1 4 .

The ply angles of the forward nose and aft inlet rings were changed to eliminate pocketingerosion in the af t in let r ing and spal l ing between the forward nose and af t in let r ings. The lowsodium carbon cloth material was used to improve the erosion performance of the rings. The ply

NEW PLY ANGLE DESIGNLOW SODIUM CARBON CLOTH PHENOLIC

K U.S. POLYMERIC, 150 ppm MAX FABRIC\ / — N E W U . S . P O L Y M E R I C C A R B O N

I \ ^ - - — ~ ~ ~ ~ > ^ / B L A C K S U P P L I E R ( F W D E X I T C O N E )

■NEW FIBERITE CARBON BLACKSUPPLIER (AFT EXIT CONE)

MODIFIED NOZZLESEVERANCE SYSTEM

^ R J P E ^ -

Figure 14. DM-6 Nozzle Differences from HPM Production Design

CURRENT DESIGNTHROATINLET RING

THROAT53.86 DIA

NOSE CAP402 106.1 DIA

DM-6 DESIGNMODIFIED PLY 75 deg/ J^PA N G L E I N ^ ' j p &403 AND 404 /V J7R I N G S 1 0 5 , d e g /

Figure 15. Comparison of Current and DM-6 Nozzle Designs

angle redesign is compared to the current production design in Fig. 15. Increased bond gaps werealso introduced between the mating rings to reduce interface loads. Structural analyses concludethat high fiber strain is the primary cause of pocketing in the current production nozzle design.Studies were conducted to determine the lowest tensile fiber strain design that could be fabricatedwith the existing manufacturing tooling and equipment. As a result, ply angles of 105 and 75 degwere selected for the forward nose and aft inlet rings, respectively.

The nozzle aft exit cone severance system was modified to improve propagation of thedetonation wave across the linear shaped charge (LSC) segment interfaces and eliminate leakage atthe LSC segment end seal. This was accomplished by: (1) increasing the minimum LSC segment endseal gap from 0.0 to 0.04 in.; (2) cutting back the LSC retainer to allow measurement of segmentalignment; (3) reducing the maximum internal end seal-to-charge clearance from 0.012 in. to 0.008in.; and (4) replacing the end seal adhesive with Loctite RC680.

The overall performance of the nozzle ablative parts was excellent except for two locations orthe nose cap where char material was apparently wedged out at the nose cap-to-forward nose ringinterface. Based on this test it was concluded that the performance of the new ply angles used onthe forward nose and aft inlet rings was superior to the ply angles on current production rings.In addition, the performance of the exit cone liners with the new carbon black filler was as goodor better than with the old filler and the liners are considered to be qualified for use on theproduction flight nozzles.

Figure 16 presents erosion and char profiles of the nose inlet at the two locations (34 and354 deg) showing the wedged out condition of the nose cap material and the resulting effect it hadon the erosion of the adjacent forward nose ring. Close examination of the sectioned parts at 34deg indicated a 0.5 in. deep and 5 in. long piece of char was wedged out about 80 sec afterignition. The similarity of the erosion profile of the forward nose ring at the 354 deg locationindicates a wedge shaped piece of char may have been forced out here also but at an earlier timein the firing. Post-test thermal-structural analyses of the DM-6 nozzle indicate that the charredtape wrapped nose cap plies may have been wedged out as a result of higher cross-ply tensilestresses present due to the higher than design bond gap when combined with the relatively highinterlaminar shear stresses present during the entire period of motor operation. Increasing thebond gap between the two rings decreased the potential for char ply delamination in the forwardnose ring but increased the potential in the nose cap. Subsequent static test nozzles will befabricated with decreased bond gaps to reverse this trend.

EROSION

3 5 4 D E G ,A F T I N L E T v < ^R I N G ^ . / \ ~ 3 - I N .

34 DEG

CHAR DEPTH

FORWARDNOSE RING

EROSION

CHAR DEPTH

SIMILAR WASHOUTAREA IN FORWARDNOSE RING

WEDGE OUTIN NOSE CAP

WEDGE OUT OCCURREDEARLY IN TEST

WEDGE OUT OCCURREDABOUT 80 sec AFTER IGNITION

Figure 16. DM-6 Nozzle Inlet Erosion/Char Depth Profiles at WedgeOut Locations

CONCLUSION

The first FWC-SRM (DM-6) was successfully tested in October 1984. Analysis of the test dataverified that all test requirements were either met or exceeded resulting in approval to continueto process the FWC-SRM segments for use on the first flight from VAFB scheduled in early 1986.Two other static tests will be conducted in the FWC-SRM development program before the firstflight. Another development test (DM-7) is scheduled for May 1985 and a qualification test (QM-4)is planned for late 1985.

The tape wrapped nozzle inlet rings with the new ply angles tested on DM-6 did not experiencethe pocketing erosion which has been observed on several recovered flight nozzles. As a result,nozzle inlet rings with the new ply angles will be used on the two remaining FWC-SRM static tests.Any decision to change the current flight nozzles to incorporate the new ply angles will bedeferred until the data from the two additional static test nozzles confirm the DM-6 results.

The use of the FWC-SRM in the first Shuttle flight from VAFB will conclude the developmentprogram for the last funded SRM performance improvement option. However, due to the costeffectiveness of implementing performance improvements in the SRM, studies are underway to furthermodify the.SRM to provide additional payload gains for selected Shuttle flights several years inthe future.