-P 

NASA Technical Memorandum 88989 

ANALYSIS AND TEST OF SUPERPLASTICALLY FORMED TITANIUM HATSTIFFENED PANELS UNDER COMPRESSION 

HANDALL C. DAVIS, DICK M. ROYSTER, AND 

THOMAS T. BALES 

(IiASA-TM-8898s) ANALYSXS ALE !LEES? CP N87- 18 113 

SUPEKPLASlICAlLk EGIiI'lEI; TIIALICE 

hAl-S'IIFEENED E&hELS UhDEb CCEFKESSLON 

{NASA) 11 p CSCL 20R Unclas 

G3/39 43404 

JANUARY 1987 

National Aeronautics and 

Space Administration 

Langley Research Center 

Hampton. Virginia 23665 
ANALYSIS AND TEST OF SUPERPLASTICALLY FORMED TITANIUM HAT-STIFFENED PANELS UNDER COMPRESSION 

Randall C. Davis 

Dick M. Royster 

Thomas T. Bales 

NASA Langley Research Center 

Hampton, Virginia 

Abstract 

Four hat stiffened titanium panels with two 

different stiffener configurations were fabricated by superplastic forming/wel d brazing and 

tested under a moderately heavy compressive 

load. The panels had the same overall dimensions but differed in the shape of the hat 

stiffener webs; three panels had stiffeners 

with flat webs and the other panel had stiffeners with beaded webs. Analysis indicated that 

the local buckling strain of the flat stiffener 

web was considerably lower than the general 

panel buckl ing strain or cap buckl ing strain. 

The analysis also showed that beading the webs 

of the hat stiffeners removed then as the critical element for local buckl i ng and improved 

the buckling strain of the panels. The analytical extensional stiffness and failure loads 

compared very well with experimental results. 

I n troduc ti on 

Recent advances in the state of the art of 

superplastic foning (SPF) of certain 

metals'-7 have made it possible to fabricate 

design shapes that depart from past experience. With the SPF process the designer has 

considerably more freedom to design mass efficient structures. Parts with intersecting compound contour surfaces can be made that would 

be impossible to fabricate with more conventional methods. 

The hat stiffened panel, shown in figure 

1, is a typical construction for aircraft 

structure. The superplastic forming/weld brazing (SPF/WB) fabrication process has been 

ex lored for this type of geometry in titanium'~~. The primary function of the material 

in the webs of the hat stiffener is to support 

the load carrying caps. For this purpose the 

webs should be made as thin as possible yet 

have enough bending stiffness to provide adequate support for the caps5. For a panel with 

stiffener caps and webs fabricated from a single sheet of material * the requirement for thin 

webs conflicts with the requirement for a cap 

with high local buckling strain. 

Beads in the stiffener webs (see figure 2) 

increase their transverse bending stiffness to 

provide more cap support and produce a web with 

a high local buckling strain. Thus if the 

panel configurations were optimized the beaded 

webs would allow for a deeper stiffener that 

would increase the general panel buckl ing 

strain. However, because of the fabrication 

constraints of this study (discussed bel ow) 

such an optimization was not carried out. The 

geometry shown offered a simple test specimen 

that could be formed by SPF/WB to, check out 

fabrication process details. Titanium sheet 

material 0.050 inch thick was used because it 

was readily available. 

For this study, two hat-stiffened panel 

geometries were fabricated using the SPF/WB 

process. The only fabrication variable available for change was the geometry of the stiffener. Three panels were fabricated with flat 

webs and one panel was fabricated with beaded 

webs. The two geometries used the same basic 

mold and started with the same nominal gage 

sheet material. The panel length was sufficiently low to preclude general panel buckling. The purpose of this report is to present 

the results of the analytical and experimental 

investigation of the effect of beading the webs 

on the buckling strain for hat-stiffened panel s. 

Analysis 

The analytical local buckling strain of 

any flat element depends on a boundary condition factor times the square of the ratio of 

the element thickness to width. If all of the 

elements in the cross-section have a similar 

boundary condition factor, then their local 

buckling strains can be ordered by comparing 

their ratio of thickness to width. Using this 

criteria, the local buckling strain of the flat 

stiffener webs (figure 1) is lower than the 

other elements in the cross-section. The skin 

under the caps has the next lowest local buckling strain. The caps have a local buckling 

strain less than the double layer skin between 

stiffeners which has the highest local buckl ing 

strain of all of the elements. 

Beading provides a web with low extensional stiffness and with extra transverse 

bending stiffness so that, based on the ratio 

of thickness to width, the skin under the 

stiffeners now has the lowest local buckling 

strain of the remaining elements. Thus beading 

the stiffener webs will raise the strain level 

needed to cause local buckling in the panel 

cross-section. However, the webs are no longer 

load carrying elements of the panel and the net 

effect on the total load the panel could carry 

must be considered. 

A PASCOO analysis of a model of the conventional hat stiffened panel cross-section 

verified the order of the local buckling 

strains among the elements given by the thickness to width analysis and showed that they 

were all below the critical strain for a general panel buckling mode. The PASCO code uses 

colnpatibil i ty between elements of the crosssection and does not have to assume a boundary 

condition factor for the elements, thereby giving an accurate computed element buckling 

strain. However, the program cannot account 
for loss of panel stiffness when an element 

buckles, therefore the code cannot be used to 

find the higher buckling loads or the maximum 

load for the panel. 

The panels are made from two sheets of 

0.05 inch material. Regardless of reshaping by 

the SPF process, the cross-sectional area, mass 

and extensional stiffness (EA) of the panels 

have values equal to those of two sheets 0.05 

inch thick. Using a nominal modulus of 

17,000,000 psi, the computed extensional stiffness (EA) is 54,000,000 lbs. Using this value 

of cxteiisi cnal stiff ness and the coi.!,CteC 1 ocdl 

buckling strains for the flat webs, the stiffener webs are calculated to buckle at a panel 

load of 200,000 lbs. To compute a maximum load 

for the panel, the stiffener webs and the skin 

under the stiffeners are assumed to be buckled 

and their effective widths are reduced to one 

half. Using this criteria for the reduced load 

carrying area and a yield stress of 133,800 

p:i, the maximum load computed for the conventional hat stiffened panel is 340,000 lbs. 

A PASCO analysis of the beaded hat stiffened panel shown in figure 2 is not so straight 

forward as the conventional hat-stiffened panel. The PASCO code is a linked strip analysis 

and requires that the geometry of the strips be 

uniform along the panel length. Therefore, 

stiffness properties of a uniform strip equivalent to the beaded webs had to be generated. 

To this end, the beaded web geometry was analy- sed as a separate inodel on PASCO to determine 

the web overall bending and extensional stiffnesses. Then using the computed bending and 

extensional stiffnesses the geometry and modulii of an equivalent element were determined5. 

Reading the sides of the stiffeners reduces the computed EA by 11% to 48,300,0001bs. 

Based on this analytical EA and the computed 

local buckling strains, the skin under the 

stiffeners is calculated to buckle at about 

212,500 lbs. This implies a 6% improvement in 

the lowest local buckle load over the conventional hat stiffened panel. Assuming the skin 

buckles under the stiffener leaving an effective width of one half, the beaded hat stiffened panel is calculated to carry a maximum 

load at yield of 316,600 lbs. This implies a 

7% reduction in maximum load capacity for the 

beaded hat stiffened panel compared with the 

conventional hat stiffened panel. 

Panel Fabrication 

The panels were SPF Ti-6A1-4V alloy multistiffener sheets joined to a Ti-6A1-4V alloy 

skin. The nulti-stiffeners were SPF from one 

0.05 inch thick sheet in a single forming operation. Prior to SPF, the sheet was chemically 

cleaned, sprayed with a die release compound 

and placed between the cover plate and mold as 

indicated in figure 3. The mold assembly was 

positioned between resistance heated ceramic 

platens which were mounted in a hydraulic 

press. The mold assembly was heated to the 

forming temperature of 1700°F by means of the 

resistance heaters in the ceramic platens. 

After heating, a load was applied to the platens to establish a gas-tight seal between the 

,titanium sheet material, the upper cover plate, 

and the mold. Argon gas was then injected 

between the upper surface of the titanium sheet 

and the cover plate. The argon gas pressure 

was then increased to approximately 100 psi to 

superplastically form the sheet. The applied 

load of the hydraulic press was increased to 

react the forming pressure. At 100 psi gas 

pressure, the forming time required to SPF a 

mu1 tiple stiffener sheet was approximately 60 

minutes. 

The configuration of the two male die 

inserts for the panels is shown in figure 4. 

The conventional hat stiffener dies were 

machined from 17-4 stainless steel bar stock. 

The beaded hat stiffened die inserts were cast 

to shape using 17-4 stainless steel. The conventional hat and beaded hat dies were interchangeable and fastened by screws to the 

retainer plate. 

Following removal from the mold and chemical cleaning, the SPF multiple stiffener sheets 

were joined to the Ti-6A1-4V panel skin sheets 

by two rows of eight spot welds between each 

stiffener. The spot welds were sufficient to 

maintain alignment and no additional tooling 

was required. Strips of 0.016 inch thick 3003 

aluminum braze alloy were placed adjacent to 

the joints. The assemblage was first placed in 

a vacuum brazing furnace, heated to 1250°F and 

held for 5 minutes to complete the brazing, and 

finally furnace cooled. At the brazing temperature, the molten braze alloy was drawn into 

the faying surfaces of the joints by capillary 

action. 

Test Procedure 

The panel ends were potted and machined 

flat and parallel to each other. A machined 

steel bar was placed between each end of the 

panel and faces of the test machine platens. 

Feeler gage measurements were used to fit shims 

between the steel bars and the platens to 

ensure bending-free contact with the panels. 

2esting against the steel bars, the flat panel 

ends offered considerable restraint against 

panel end rotation. The panel edges were mounted in knife edge supports to simulate simply 

supported edges. 

The specimens were loaded to failure in 

compression between the displacement control led 

platens of a one million pound test machine at 

the rate of 60,000 pounds per minute. The panels were instrumented with electrical resistance strain gages. Platen displacements were 

monitored with 1 i near vol tage displacement 

transducers (LVDT). The load, strain gage 

responses and di spl acements were inoni tored by a 

data acquisition system that recorded the data 

on a digital computer tape. 

Three of the failed panels (panels designated 2, 3 and 4 in Table 1) were cut up to 

make material coupon test specimens. The COUpon specimens were marked with uniform lines 

across the test section and instrumented with 

electric resistance strain gages. The coupons 

were pulled to fracture in a tensile test 

machine. Uniform elongations were measured from 

marked gage lines over a 1 inch section of the 

specimen away from the fracture zone. Elongations across the fracture were measured from 

L 

2 
marked gage lines over a 2 inch section across 

the fracture zone after the specimen had been 

put back together. 

Results and Discussions 

Tensile property data from the coupon 

tests are given in tables 2, 3 and 4. The 

average modulus was 17.000 ksi and the average 

material yield stress was 133,800 psi. Comparing the average coupon data obtained from the 

cap and skin areas, indicate that the fabrication process had little or no effect on the 

basic material properties. 

Conventional Hat Stiffened Panels 

Local thinning of the cross section by 

the SPF process nust be taken into account to 

properly interpret the experimental results. 

The titanium sheet prior to forming rests on 

top of the male die insert in the mold cavity. 

The SPF forming pressure stretches the sheet 

over the male die insert. The sheet material 

lying on top of the die insert undergoes very 

1 ittle stretching and becomes the stiffener 

cap. As the sheet superplastically forms into 

the mold, it stretches and makes contact with 

the sides of the die insert and forms the webs 

of the stiffeners. As the sheet progressively 

forms into the mold, the stretching causes the 

stiffener webs to be progressively thinner 

toward the bottom of the mold cavity. The 

resulting stiffener cross-section has a cap 

with a thickness nearly equal to the original 

sheet thickness and webs with a tapering thickness (see figure 5). 

As the titanium sheet forms into the mold 

cavity, at the ends of the die insert, it not 

only stretches transverse to the stiffener as 

it contacts the sides of the die insert, but it 

also stretches along the stiffener as it contacts the side of the mold between the dies. 

This bi-directional stretching causes greater 

thinning to occur in the sheet near the bottom 

of the mold. As a result there is more thinning of the stiffener webs and of the skin section between stiffeners near the panel ends 

than in the middle of the panel. 

This additional thinning causes the 

position of the neutral surface of the panel 

to vary along the panel length. Its position 

is closer to the skin near the ends than in the 

middle of the panel. This variation in position of the neutral surface produces an offset 

load path for the axial load that causes the 

panel to bend. This could be prevented by cutting off more of the ends of the panels. The 

offset load path explains why the strain gage 

response data for the conventional stiffened 

panels showed that the panel s were bending from 

the onset of loqding, with the skin side bowing 

inward (see figure 6). 

The experimental extensional stiffness 

values (from load displacement plots such as in 

figure 7) for the conventional hat stiffened 

panels averagedout to 51,400,000 lbs(see table 

1). This value is 5% lower than the computed 

extensional stiffness value. Since coupon tests 

determined that the material modulus was close 

to the nominal value used for the computations, 

the slightly lower than expected value of EA is 

attributed to the greater thinning near the 

panel ends. 

The stiffener caps in all three conventional hat stiffened panels buckled near the 

potted ends. Because the potted ends of the 

panel prevent rotation as the panel begins to 

bend, the stiffener caps near the ends were 

subjected to an additional compressive component due to the end moment required to prevent 

rotation. Thus these cap buckles were caused 

by increased compression strains combined with 

thinner webs near the ends. Conventional hat 

stiffened panels 1 and 2 had no strain gages 

located near the panel ends. However, several 

gages were placed on the stiffener caps 1.5 

inch above the potting material on panel 3. 

The response curves from these strain gages, 

for two different stiffeners, are shown in figure 8. One gage was nearly centered on a local 

cap buckle and the other was just outside the 

local cap buckle. These strain gages showed 

that strain reversal occurred near the panel 

ends before any reversal occurred near the middle of the panel (compare figures 8 and 9). 

The lowest strain reversal in panel 3 occurred 

in the caps at the panel ends at a load of 

193,000 lbs. This is slightly lower than the 

200,000 lb. computed panel load based on local 

buckling strain in the stiffener webs. The 

buckling analysis, however, did not include any 

of the bending effects that appeared in the 

experimental buckling results. It can only be 

be assuined that panels 1 and 2 also had reversals at lower loads near the panel ends as found 

in panel 3. Since panel 3 was the only conventional hat stiffened panel with gages at the 

panel ends, for comparative purposes, the initial strain reversal data shown in Table 1 are 

from gages located near the panel midsections. 

The average of these loads is 19% higher than 

the computed local buckling load based on the 

analytical buckling strain in the stiffener 

webs. 

Local buckles in the caps near the ends of 

the panels increase the bending moment at the 

panel midsection because of the additional 

shift in the location of the neutral surface. 

At the midsection the bending strains were 

reversed from those at the panel ends. Strain 

gage response data showed that the compressive 

bending strains combined with the axial compressive strain to cause a higher strain in the 

panel skin than in the caps. As can be seen by 

the strain gage response curves shown in figure 

10, the skin in the panel midsection buckles 

before the caps. 

At failure, a catastrophic collapse 

occurred that produced permanent buckles in the 

caps and webs near the panel ends and in the 

skin at the panel midsections (see figures 11 

and 12). None of the weld braze joints in the 

panels failed during the test. The corresponding maximum or collapse loads recorded from the 

panel tests are shown in Table 1. The 296,870 

lb average for these loads agrees within 131, of 

the computed 340,000 lb maximum load. 

3 
- 

Beaded Hat Stiffened Panel Conclusions 

The neutral axis location for the beaded 

hat stiffened panel does not depend on the 

beaded stiffener webs because the webs are not 

load carrying. Thus greater thinning of the 

beaded webs near the panel ends has little 

effect on the location of the neutral surface 

panels. Hence the beaded hat stiffened panel 

has less bending than the conventional hat 

stiffened panel. This explains strain gage 

response data from the beaded hat stiffened 

panels showing nuch less bending in the panel 

with loading compared to the conventional hat 

stiffened panels (see figure 13). 

The experimental results from LVDT data 

show that the beaded hat stiffened panel EA 

stiffness (Table 1) was 11% lower than the EA 

for the conventional hat stiffened panel. This 

compares very well with the 11% reduction 

expected based on the computed EA values for 

the two types of panels. 

The experimental EA value for the beaded 

hat stiffened panel was 5% lower than the computed value. Since the nominal modulus value 

used in the computations was close to the experimental value obtained from coupon tests, the 

slightly lower panel EA stiffness value froin 

test is attributed to iwre thinning of the 

sheet material between stiffeners near the ends 

due to bi-directional stretching. 

Since the bending occurring near the ends 

of the beaded hat stiffened panel was slight 

based on strain gage response data (see figure 

141, the caps near the ends did not buckle 

early as was observed in the conventional hat 

stiffened panels. As the panel strain level 

approached material yield, the skin under the 

stiffeners at the panel midsection began to 

buckle (see figure 13) causing the panel skin 

side to bow inward. This induced local buckles 

in the skin which began to occur at a load of 

268,500 lbs. The experimental local buckling 

load for the beaded hat stiffened panel was 13% 

higher than the corresponding load for the conventional hat stiffened panel. This is more of 

an increase in buckling strain between the two 

types of panels than was expected, based on 

computations for beading the sides. However, 

most of this additional improvement is 

attributed to less bending in the beaded hat 

stiffened panel compared to the conventional 

hat stiffened panel. 

The local buckles in the panel skin 

resulted in a pronounced inward bowing of the 

panel skin and a subsequent catastrophic collapse of the beaded hat stiffened panel across 

the midsection. The resulting collapse at 

280,000 lbs (see table 1) caused fractures to 

occur at the midsection and produced deep buckles with fractures near the panel ends (see 

figures 15 and 16). There were no separations 

in the weld brazed joints observed in the 

test. The failure load was 11% below the computed maximum load. The maximum load for the 

beaded hat stiffened panel was 6% below the 

average failure load for the conventional hat 

stiffened panel which agrees well with the 7% 

computed reduction. 

Four titanium multi-stiffener panels were 

successfully fabricated using the SPF/WB 

technique. Three panels had conventional hat 

stiffeners (flat webs) and one panel had hat 

stiffeners with beaded webs. The panels were 

tested to failure in axial compression. None 

of the weld braze joints in the panels failed 

or separated during the tests. 

Analyses were made for both types of 

panels. According to analysis, the local 

buckling strain of the flat webs is lower than 

the buckling strain of the other elements in 

the cross section. Analysis showed that 

beading the webs of the hat stiffeners raised 

the buckling strain of the panel by 6%. However 

it also removed 11% of the load carrying 

material from the cross section. This 

reduction in extensional stiffness reduced the 

computed maximum load by 7%. 

Test results for the conventional hat 

stiffened panels showed bending occurred in the 

panels under an axial load. The bending is 

attributed to variation in the position of the 

non-uniform neutral surface, caused by 

differential thinning of material along the 

panel length. This could be prevented by 

cutting off more of the ends of the panels. At 

the panel ends this bending combined with the 

axial load causing a higher strain in the 

stiffener caps than in the skin. Thus the caps 

of the conventional hat stiffeners buckle 

locally near the ends of the panels. The 

location of the neutral surface along the 

length of the beaded hat stiffened panel is not 

affected as much by SPF thinning as the 

conventional hat stiffened panel. Hence the 

experimental results for the beaded hat 

stiffened panel showed very little bending. 

Consequently the beaded hat stiffened panel did 

not buckle near the ends as did the 

conventional hat stiffened panels. 

The experimental extensional stiffness (EA 

or ratio of load to strain) values were only 5% 

lovJer than calculated presumably due to more 

thinning in the stiffener sheets near the ends 

of the stiffeners. Both the computed and the 

experimental results for the beaded hat stiffened panel showed a 11% reduction in the EA 

stiffness value compared to the conventional 

hat stiffened panel. 

The buckling loads from the tests showed the 

beaded hat stiffened panel had an increase of 

13% in local buckling load over the conventional hat stiffened panel. This is higher than 

the 6% calculated increase for the beaded hat 

stiffened panel buckling load and is attributed 

to a combination of decreased bending as well 

as to beading of the stiffener sides. The 

bending observed in the experimental results 

was not included in the analysis. Maximum 

loads for the panels were governed by yield and 

differed only slightly between the beaded and 

conventional hat stiffened panels. Experimental maximum load for the beaded panel was 6% 

lower than the average of the conventional 

panel maximum loads compared to a computed 7% 

reduction. 

4 
References 

1. Bales, Thomas T., ed.: SPF/DB Titanium Tech2. Bales, Thomas T.; Royster, Dick M.; and 

Arnold, Winfred E., Jr.: Development of the 

Weld-Braze Joining Process. NASA TN D-7281, 

1973. 

3. Royster, Dick M.; Bales, Thomas T.; and 

Wiant. H. Ross: Superplastic Forming/Weld-Brazing of Titanium Skin-Stiffened Compression 

Panels. Materials Overview for 1982, Vol. 22, 

National SAMPE Symposium and Exhibition, 1982 , 

4. Royster, Dick M.; Davis, Randall C.; Shinn, 

Joseph M.; Bales, Thomas T.; and Wiant, H. 

Ross: Fabrication and Eva1 uation of Superpl astical ly Formed/Wel d Brazed Corrugated Compres- sion Panels with Beaded Webs. NASA Tp 2512, 

1985. 

5. Davis, Randall C.; Mills, Charles T.; 

Prabhakaran,R.; and Jackson, L. Robert: Structural Efficiency Studies of Corrugated Compression Panels with Curved Caps and Beaded Webs. 

NASA TP 2272, Feb. 1984. 

6. Royster, Dick M.; and Bales, Thomas T.: Elevated Temperature Behavior of Superplastical ly 

Formed/Weld Brazed Titanium Compression Panels 

having Advanced Shaped Stiffeners. NASA TP 

2123, 1983. 

7. Royster, Dick M.; Bales, Thomas T.; and 

Davis, Randall C.: Superplastically Forming and 

Weld Brazing for Fabrication of Titanium Compression Panels. Presented at: Symposiun on 

nology. NASA CP-2160. OCt. 1980. 

pp. 569-582. 

Welding, Bonding and Fastening, NASA. Langley, 

OCt. 23-25, 1985. 

8. Anderson, Melvin S.; and Stroud, W. Jefferson: A General Panel Sizing Computer Code and 

Its Application to Composite Structural Panels. AIAA J., Vol. 17,.NO. 8, Aug. 1979, pp. 

892-897. 

5 
Table 4. Material prDperty coupon test results for panel #4. 

panel 

Table 1. Experlmntal and analytical results 

Extensional Cornpuled Reversal Maximum Computed 

stiffness, EA e?ensioMl Weight id load mx 

Ibs Ibs stiffness, EA Ibs Ibs 

Ibs X lo6 Ibs 106 

- 

.cationspecimen 

1-1 

1-2 

1-3 

Avg 1 

2-1 

2 -2 

2 -3 

Avg 2 

3-1 

3 -2 

3-3 

Avg 3 

- 

- 

1 

2 

3 

- 

Area 

in2 

53.3 54.1 10.96 219.764 u)4,130 340,ooO 

54.6 54.1 NA* 273,660 319,134 340,ooO 

46.2 54.1 10.71 217,000 267,360 340,ooO 

- 

Max 

load 

Ibs 

6Mo 

6620 

65M 

647 3 

3590 

36M) 

3560 

3583 

3950 

4055 

4090 

4032 

- 

- 

Yield Ultimate 

stress stress 

Kpsi Kpsi 

131.9 139.6 

133.0 141.1 

133.6 141.3 

132.8 140.7 

133.0 141.9 

132.4 142.3 

134.8 142.4 

133.4 142.2 

134.8 142.6 

134.5 142.8 t 133.5 142.0 

134.3 142.5 

modulus elongation across fraduri 

(1 Inch) I2 Inch) 

17.3 11 11.5 

17.3 6.5 

17.3 7.7 10 

19.0 10 

18.8 0 

. M51 

,0468 

.o060 

.Mho 

.m3 

. m3 

. mo 

. m2 

. m7 

. m84 

. mas 

- 

,0283 

16.7 8 9.5 

:::; I 78.3 I i.5 Nd available 

Location of specimens a cut from panel 

13 

1.12 

--I b3.53 -4 

10 

ty coupon 

Yield 

stress 

- 

KpSi 

Table 2. Material ora I test results for panel #2, conventional hat. 

135.3 

138.8 

137.0 

127.9 

1265 

127.2 

143.0 

145.6 

144.3 

2 

- 

z -Location d specimens 11. cut from panel 

13 

T 

wide I 

9 stiffeners 20.98 

k31.804 

tt 

Fig. 1 Geometric details of conventional hat 

stiffened panels (dimensions are in inches). 

Table 3. Material property coupon test results for panel #3. conventional hat. 

- JI r 180° 

Area 

in2 

. MM 

.M32 

. 0636 

,0633 

.W46 

. Q2M 

.m24 

. Q234 

. Q246 

. U241 

. w4 

. Q239 

- 

- 

I E 

modulus 

psi x 106 

16.5 

16.7 

16.8 

16.7 

17.4 

17.9 

18.0 

17.7 

16.6 

16 6 

16.8 

16. 7 

Locationspeclmen I 

b.54 

f 

k- 31.80 4 Section AA 

6090 13'2.0 141.6 

130.7 140.0 

6100 130.7 139.9 

131.1 140.5 

1268 133.9 

32260 130.4 140.5 

b 

A4 1 

2 -2 

2-3 

Avg 2 

3-2 

3-3 

Avg 3 

7.3 10.3 

3 3 

5 I 1:5 I 

3150 130.8 140.6 8 

E I ::::: I E:: I b 

5.3 7.5 

6 

:.7 I thl J 

Location d speclmens a cut from panel 

13 Fig. 2 Geometric details of the beaded hat 

stiffened panel (dimensions are in inches). 

ORIGINfiL pi?. '2 f7 

OF Pcm oa*q: "fv 6 
A- Argon gas inlet 

Fig. 3 Schematic of tooling for superplastic 

forming of the hat stiffeners. 

1 Stiffener die Stiffener die 

Retainer plate Retainer plate 1 

(a) conventional hat 

stiffened panel stiffened panel 

configuration configuration 

(b) beaded hat 

Fig. 4 Configurations of die inserts. 

Fig. 5 Photograph of a section through the 

conventional hat stiffened panel showing 

effect of SPF thinning on the thickness. 

3oo r Stiffener 

Load, 

Klbs 

200 

100 

0 .002 .004 .006 .m 

Strain 

Fig. 6 Midsection back-to-back strain gage 

resDonse curves for the conventional 

300 

200 

Load, 

Klbs 

100 

0 

Fig. 7 

hat' stiffened Panel 3. 

load, 

Klbs 

.E .10 .15 

Deflection, in. 

Typical load-deflection response curve 

from LVDT placed between test machine 

platens (from Panel 3). 

0 .002 .OM ,006 .008 

Strain 

Fig. 8 Typical strain gage response curves from 

two stiffener caps near the end of the 

conventional hat stiffened Panel 3. 

7 
300 r 

OF 

Stiffener Stiffener 

Load, 

Klbs 

Strain 

Fig. 9 Strain gage response curves from a 

stiffener at the middle of a conventional 

hat stiffened panel. 

Load, 

Klbs 

300 

200 

100 

0 .001. OM. 003.004.005 

Strain 

Fig. 10 Midsection strain gage response curves 

from back-to-back gages that show skin 

strain reversal occurring before the 

cap strain reversal. 

Flg. 11 Conventional hat stiffened panel showing 

stiffeners after collapse. 

Fig. 12 Conventional hat stiffened panel showing 

buckles across skin side after collapse. 

300 r 

200 

Load, 

Klbs 

100 

0 .002 .004 .OD6 .DO8 

Strain 

Fig. 13 Typical back-to-back strain gage response 

curves from midsection of beaded hat 

stiffened panel. 

300 r 

Load, 

Klbs 

Strain 

Flg. 14 Typical strain gage response curves from 

caps of beaded hat stiffeners near the 

ends of the panel. 

8 
Fig. 15 Beaded hat stiffened panel showing 

damaged stiffeners after collapse. 

Fig. 16 Beaded hat stiffened panel showing 

skin side after collapse. 

9 
Standard Bibliographic Page 

HAT-STI FFENED PANELS UNDER COMPRESS I ON 

'. Author(s) 

Randall C. Davis, Dick M. Royster, Thomas T. Bales ~ 

b. Performing Organization Name and Address 

Report No. 

ASA TM- 88989 

6. RPrfnNning Organization Code 

505-63-81-01 

8. Performing Organization Report NO. 

10. Work Unit No. 

2. Government Accession No. 3. Recipient's Catalog No. 

17. Key Words (Suggested by Authors(s)) 

I I 

18 Distribution Statement 

NASA Langley Research Center 

Hampton, VA 23665 

12. Sponsoring Agency Name and Address 

11. Contract or Grant No. 

13. Type of Report and Period Covered 

Technical Memorandum 

National Aeronautics and Space Admini stration FSpoiisoring Agency Code 

Washi ngton, DC 20546 I- - 

15. Supplementary Notes 

Presented at the 26th AIAA/ASME/ASE/AHS Structures, Sturctural Dynamics and 

Materials Conference, San Antonio, TX, May 1986 

our hat-stiffened titanium panels with two different stiffener 16. Abstrap 

configurations were fabricated by superplastic forming/weld brazing and tested 

under a moderately heavy compressive load. 

dimensions but differed in the shape of the hat-stiffener webs; three panels 

had stiffeners with flat webs and the other panel had stiffeners with beaded 

webs. 

web was considerably lower than the general panel buckling strain or cap 

buckling strain. 

stiffeners removed them as the critical element for local buck1 ing and improved 

the buckling strain of the panels. The analytical extensional stiffness and 

failure loads compared very well with experimental results. 

The panels had the same overall 

Analysis indicated that the local buckling strain of the flat stiffener 

The analysis also showed that beading the webs of the hat 

Ti tani um 

Superplastically Formed 

Buck1 i ng 

Hat-Sti ffened Panels 

Unclassified - Unlimited 

I 

Subject Category 39 

20. Security Classif.(of this page) 21. No. of Pages 22. Price 

Unclassified ~____.__. Unci assi f i ___-_ ed --I107 A02 

~~ 

19. Security Clarsif.(of this report) 

--..

For sale by the National Technical Information Service, Springfield, Virginia 22161 

NASA Langley Form 63 (June 1985)