February 2006 

NASA/TM-2006-214274 

Damage Tolerance Analysis of a Pressurized 

Liquid Oxygen Tank 

Scott C. Forth, Stephen F. Harvin, Peyton B. Gregory, Brian H. Mason, Jon E. Thompson, and 

Eric K. Hoffman 

Langley Research Center, Hampton, Virginia 
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February 2006 

NASA/TM-2006-214274 

Damage Tolerance Analysis of a Pressurized 

Liquid Oxygen Tank 

Scott C. Forth, Stephen F. Harvin, Peyton B. Gregory, Brian H. Mason, Jon E. Thompson, and 

Eric K. Hoffman 

Langley Research Center, Hampton, Virginia 
Available from: 

NASA Center for AeroSpace Information (CASI) National Technical Information Service (NTIS) 

7121 Standard Drive 5285 Port Royal Road 

Hanover, MD 21076-1320 Springfield, VA 22161-2171 

(301) 621-0390 (703) 605-6000 

The use of trademarks or names of manufacturers in the report is for accurate reporting and does not 

constitute an official endorsement, either expressed or implied, of such products or manufacturers by the 

National Aeronautics and Space Administration. 
Abstract 

A damage tolerance assessment was conducted of an 8,000 gallon 

pressurized Liquid Oxygen (LOX) tank. The LOX tank is constructed of 

a stainless steel pressure vessel enclosed by a thermal-insulating vacuum 

jacket. The vessel is pressurized to 2,250 psi with gaseous nitrogen 

resulting in both thermal and pressure stresses on the tank wall. Finite 

element analyses were performed on the tank to characterize the stresses 

from operation. Engineering material data was found from both the 

construction of the tank and the technical literature. An initial damage 

state was assumed based on records of a nondestructive inspection 

performed on the tank. The damage tolerance analyses were conducted 

using the NASGRO computer code. This paper contains the 

assumptions, and justifications, made for the input parameters to the 

damage tolerance analyses and the results of the damage tolerance 

analyses with a discussion on the operational safety of the LOX tank. 

Introduction 

A damage tolerance assessment was conducted of a 

pressurized LOX Tank. The LOX tank consists of a 

stainless steel pressure vessel enclosed by a thermalinsulating vacuum jacket. The subject LOX tank was 

utilized for rocket motor testing at Edwards Rocket 

Test Facility in California beginning in 1960 and 

concluding sometime prior to 1982, when the tank was 

first examined by NASA Langley personnel as surplus 

property. NASA Langley obtained the LOX tank 

during a project to upgrade the Eight-foot High 

Temperature Tunnel (8’HTT) in the mid 1980’s. 

During operation, 85 pounds per second of cryogenic 

oxygen at 2250 psi is required to flow from the LOX 

tank to the facility. To achieve this flow rate and 

pressure, the LOX tank is pressurized to 2,250 psi with 

gaseous nitrogen for each operational cycle. A typical 

operational cycle has a two minute duration. The 

pressurization and subsequent depressurization of the 

tank produce both thermal and pressure stresses on the 

vessel wall. Finite element analyses were performed to 

characterize the stresses from operation. Engineering 

material data for the stainless steel was found from 

both the construction of the tank and the technical 

literature. An initial damage state needed for the 

damage tolerance analysis was assumed based on 

records of a nondestructive inspection performed on 

the tank. The damage tolerance analyses were then 

conducted using the NASGRO computer code [1]. 

This paper contains the assumptions, and justifications, 

made for the input parameters to the damage tolerance 

analyses, and a discussion of the results of the damage 

tolerance analyses. 

LOX Tank Construction 

The Liquid Oxygen Tank (LOX) was manufactured in 

1959 by Yuba Consolidated Industries Inc., Southwest 

Welding and Manufacturing Division under contract to 

Rocketdyne, a Division of North American Aviation. 

The fabrication specifications are contained in the 

1959 Rocketdyne Specification 218-7 with two 

addendums [2]. The tank and its vacuum jacket were 

required to meet the 1959 requirements of the ASME 

Boiler and Pressure Vessel Code, Sections VIII [3] and 

IX [4]. The fabrication of the pressure vessel was 

supervised by the Division of Industrial Safety for the 

State of California. 

The tank is constructed of an inner stainless steel 

pressure vessel surrounded by a thermal insulating 

vacuum jacket. The stainless steel vessel was 

constructed using austenitic 347 stainless steel 4.73 

inch thick plates welded together radially and 

circumferentially as shown in Figure 1. The top of the 

vessel is a 347 stainless steel 11 inch thick forged 

manway that was welded to the spherical vessel walls. 

347 Stainless Steel 

The AISI 300 series austenitic stainless steels were 

developed as corrosion-resistant alloys. These alloys 

posses excellent oxidation resistance and good hightemperature strength and creep resistance and are used 

extensively at cryogenic temperatures. They alloys 
2 

can only be hardened through cold work (non heat 

treatable). 

The two main alloying elements in the austenitic 

stainless steels are chromium and nickel. Chromium 

(Cr) imparts corrosion and oxidation resistance while 

nickel gives the alloys its austenitic structure with its 

associated toughness and ductility. The base 

composition of these alloys is 18% chromium and 8% 

nickel. Varying the base composition and alloying 

with other elements creates grades with special 

characteristics. Type 347, which was used in the 

construction of the LOX run tank, has additions of 

columbium and tantalum added to stabilize the alloy 

for service in the 800°-1600°F range and to minimize 

carbide precipitation when welding for resistance to 

intergranular corrosion [5, 6]. The chemical 

composition and mechanical test requirements for the 

SA-240 type 347 are presented in ASME BPVC 

Section II [7]. 

At cryogenic service temperatures, all of the austenitic 

stainless steels have the good strength, ductility, and 

toughness. The typical mechanical properties of type 

347 stainless steel at room temperature and -320°F are 

shown in Table 1 [3, 6, 8, 9]. 

Weldments 

The shell plate segments and the forged manway and 

nozzles were welded to each other in a double-V 

groove weld (butt joint) configuration with full 

penetration back-up weld seams as shown in Figure 1. 

Weld process was shielded metal arc with a designated 

weld rod of 19-9Cb with 4% max C and 4-7% ferrite 

[10]. This is not a standard designation for welding 

rods but appears to be an AWS classification E347 

weld rod. Based upon AWS A5.4 Specification for 

Stainless Steel Electrodes for Shielded Metal Arc 

Welding [11] (identical to ASME BPVC Section II, 

Part C SFA-5.4 [7]), the chemical composition and 

mechanical property requirements for the weld metal 

are defined. 

Arc Welding 

The austenitic stainless steels can be readily welded by 

almost all of the usual arc welding processes provided 

adequate steps are taken to prevent oxidation and 

carburization of the weldment. In arc welding, an 

electrical discharge arc consisting of a thermally 

ionized plasma gas is established between the welding 

electrode and the workpiece. The electrode may be a 

consumable and provide the filler metal for the weld 

deposit such as in shielded metal arc or it may be nonconsumable and use a secondary filler metal such as in 

gas tungsten arc. In either case, the weld metal is 

protected from atmospheric oxidation by a fluxing slag 

or a chemically inert shielding gas. Commonly used 

arc welding processes include shielded metal arc, gas 

tungsten arc, gas metal arc, plasma arc, flux cored arc, 

and submerged arc welding [12, 13]. 

Filler Metal 

Weld filler metal selection is based upon the service 

conditions and the properties that the weld metal has to 

display. In general, the deposited weld metal 

composition should nearly match the base metal 

composition when welding austenitic stainless steels to 

themselves [5, 13]. 

Carbide Precipitation 

Precipitation of carbides at the grain boundaries can be 

a major problem in welding austenitic grade steels. 

Carbide precipitation is accelerated by exposure to 

temperatures within the sensitizing range (800 – 

1600°F) and by increased time at temperature. When 

intergranular Cr carbides are precipitated, resistance to 

corrosion and stress corrosion decrease markedly. 

Decrease in corrosion resistance is due to the presence 

of Cr-rich carbides at the grain boundaries which 

results in the depletion of Cr in the adjacent matrix 

material. Sensitization may result from slow cooling 

from annealing temperatures, stress relieving or 

welding. Because annealing or stress-relieving 

requires the metal to be held at temperatures for a 

relatively long period of time, it is possible that the 

entire workpiece will be sensitized [5, 6, 12, 13]. 

Carbide precipitation can be prevented or corrected by 

any one of three means. 

• Solution heat treat to put the carbides back 

into solution thus restoring the normal 

corrosion resistance. This is done by heating 

the weldment to a high temperature range 

(usually above 1850°F) and then cooling 

rapidly. 

• Use extra low C grades (0.03% max C). 

These grades have sufficient immunity to 

carbide precipitation in the 800-1600°F 

temperature range to undergo normal welding 

or stress-relieving operations without 

impairment of the corrosion resistance. 

• Use stabilized grades such as 321 or 347 

which contain Ti, Cb, or Ta. These 

stabilizing elements have a greater affinity for 

C than does Cr and carbides of these 

stabilizing elements have minimum 
3 

susceptibility to precipitation at the grain 

boundaries.

Crack Susceptibility of Weldments 

Welded joints in austenitic stainless steels are 

susceptible to hot cracking during solidification of the 

weld metal unless suitable precautions are taken. The 

degree and frequency of these cracks depend on the 

composition of the weld metal, the amount of stress 

developed in the weld as it cools, the thickness of the 

joint, ductility of the weld metal at high temperatures, 

and the presence of notches [5, 8, 13]. 

Weld metal with a fully austenitic microstructure is 

considerably more susceptible to cracking than weld 

metal with a duplex structure of ferrite in austenite. 

Generally, ferrite is beneficial when the welds are 

restrained, the joints are large, and when cracks can 

adversely affect service performance. Too much 

ferrite, however, may have detrimental affects on 

corrosion resistance in some environments and it also 

is detrimental to toughness in cryogenic service, and in 

high temperature service when it can transform to 

brittle sigma phase. Two-phase structures of austenite 

and ferrite exhibit the least cracking susceptibility 

when the ferrite is from 4-8%, with the remainder 

austenite. This is sufficient ferrite to break up the 

coarse grain structure of as-cast austenite and is 

effective way to prevent hot cracking. The 

compositions of most filler metals are adjusted by the 

manufacturers to ensure weld deposits that have the 

desired ferrite content. Type 347 weld metal is 

designed to resolidify with a small amount of ferrite to 

minimize cracking susceptibility. 

Ferrite content can be measured on a relative scale by 

various magnetic instruments or can be calculated from 

the chemical composition of the weld deposits based 

upon constitutive diagrams such as Espy, DeLong, and 

WRC-1988 diagrams [5, 8, 11, 13]. The WRC-1988 

diagram is considered the most accurate, and is the 

preferred diagram, in the welding industry for 

predicting the ferrite content for 300 series stainless 

steel weld metals. 

Residual Stresses and Post-Weld Heat 

Treatment 

During welding, as the molten weld metal pool begins 

to solidify and shrink, it begins to exert stresses on the 

on the surrounding weld metal and heat affected zones. 

When the weld metal first solidifies, it is hot and 

relatively weak and exerts little stresses. As it cools 

further, however, the stresses in the weld areas 

increase and eventually reach the yield point of the 

base metal and heat affected zone. In the case of 

multi-pass welds, the weld bead metal laid down first 

resists the shrinkage of subsequent built-up weld bead 

layers. As a result, the weld metal laid down first is 

strained in tension down the length (longitudinal 

direction) of the weld bead. For thick-plate butt joints, 

little shrinkage in the transverse direction is possible 

due to the constraint provided by the plate and the 

stiffening effect of the underlying weld bead. As a 

result, compressive stresses develop in the transverse 

direction to the weld bead. Figure 3 shows a schematic 

representation of the residual stress distribution in the 

welded butt joint [25]. 

When weldments are made in thick plate (> 1 in.), 

residual stresses can vary significantly through the 

plate thickness. Figure 4 shows the distribution of 

residual stresses in the three directions in the weld 

metal of a butt joint. Both the longitudinal and 

transverse stresses are tensile in areas near the surfaces 

of the weld while the residual stresses were primarily 

compressive below the surface. 

Residual stresses can adversely affect the performance 

of welded structure. In tension, residual stresses may 

lead to high local stresses in weld regions of low 

toughness. This high local stress can initiate cracks 

that propagate by the applied working stresses that 

would not ordinarily propagate. Weldments with 

heavy sections may be susceptible to underbead 

cracking if the stress levels due to restraint are high. 

Weldments of the Cb stabilized grades such as type 

347 are most susceptible to this type of cracking 

because of strain induced precipitation of Cb carbides 

[8, 14]. 

To relieve residual stresses and reduce the tendency for 

weld cracking, a post-weld stress relief is usually 

conducted. However, this may not be feasible 

depending upon the size of the welded structure. 

Stress relieving can be performed over a wide range of 

temperatures. The normal annealing temperature range 

for 347 steel is 1800° - 2000°F. While the primary 

purpose of the anneal is to obtain softness and high 

ductility, these steels may also be stress relieve 

annealed within the carbide precipitation range of 800° 

- 1500°F without increasing the susceptibility to 

intergranular corrosion. Recommended time at 

temperature ranges from about 1 hr/inch of section 

thickness at temperatures above 1200°F to 4 hr/ inch of 

section thickness at temperatures below 1200°F. For 

maximum ductility, the higher annealing temperature 

range of 1800° - 2000°F is recommended [5, 8, 14]. 

Based upon these annealing treatments for welds in 

austenitic stainless steels, the estimated percentage of 

residual stresses relieved at various temperatures are as 

follows: 

 1550° – 1650°F 85% stress relief 
4 

 1000° – 1200°F 35% stress relief 

To relieve the residual stresses in the weldments and 

reduce the tendency for weld cracking, the entire LOX 

run tank was post-weld stress relieved after fabrication. 

Indicated stress relieving at 1150°F for 44 hrs was 

required. A special note was added revising this heat 

treatment due to the 11 in.-thickness of the manway 

forging. The revised post-weld heat treatment for the 

LOX run tank was as follows: 

Heat to 800°F and hold for 22 hours. Heat as rapidly 

as possible from 800°F to 1800°F and hold for 11 

hours. Air quench. Based upon this stress relief and 

full annealing treatment, the estimated percentage of 

residual stresses relieved in the welds was 85 to 100 

percent [8, 14]. 

Material and Weld Process Qualification 

Weld qualification was conducted prior to fabrication 

to ensure that the actual weldments used in the 

construction of the LOX tank would have the required 

properties for the intended service conditions. 

Material and weld process qualification was conducted 

using weld procedure specification (WPS) and ASME 

BPVC, Section II, Parts A and C, Section VIII, 

Division 1 and Section IX [3, 4, 7, 11, 15]. In 

addition, a special material specification listed on 

Drawing No. D-8979 was used to qualify the materials 

and weld process used in the construction of the LOX 

tank for cryogenic service. This specification is as 

follows: 

Special Certified Low temperature property tests, in 

addition to those required by the various codes, shall 

be required on each plate, forging, pipe, cap screws, 

and weld metal in the “as welded state” that is to be 

subjected to the working pressure stress of the inner 

vessel as follows: 

Minimum requirements of metal in the annealed 

state at -300°F 

Ultimate Tensile Strength: 150 ksi 

Yield Strength: 35 ksi 

Charpy Impact Strength: 30 ft-lbs 

Elongation (In 2 Inches): 25 % 

Reduction in Area: 50 % 

Weld metal Charpy Impact Strength: 15 ft-lbs 

Performance qualifying records for the 347 plate, 

forgings, and weld metal are shown in Tables 2-4. 

Tables 2 and 3 show the qualifying mechanical 

properties of the weld metal, shell segment plate, and 

forged flanges at -320°F while Table 4 shows the 

chemical analysis of the shell plate. No chemistries 

were listed for either the forged flanges or the weld 

metal. As per ASME BPVC Sec. VIII, Division 1 [3], 

no weld metal chemistries are required for qualifying 

welds for code approved base materials. 

Fully austenitic stainless steel weld metals possess 

excellent toughness at cryogenic temperatures. To 

ensure freedom from brittle failure, ASME BPVC, 

Section VIII, Division 1 [3] requires weldments 

intended for cryogenic service to be qualified by 

Charpy V-notch testing. The criterion for acceptability 

is the attainment of a lateral expansion opposite of the 

notch of greater than 0.015 inches at -320°F. At the 

time of manufacture, the qualifying tests consisted of 

tensile and Charpy V-notch fracture toughness 

minimum values only. Based upon the above 

referenced test results, base material and weld metal 

mechanical properties are within code and special 

material specification and demonstrate sufficient 

strength, ductility, and fracture toughness for service at 

-320°F. 

LOX Tank Repairs 

The LOX tank was repaired on three separate 

occasions during it’s service life. Additionally, NASA 

Langley modified the tank for use at the 8’HTT by 

welding plugs into 6 of the small nozzles (P1 thru P6). 

These repairs and modifications are documented below 

in order to provide a complete source of information 

regarding the repair history of the tank for future 

inspection and analysis of the LOX tank. The first 

weld repair of the LOX run tank occurred in 1965. 

The second and thrid repairs of the tank occurred in 

1986 and 1991 respectively. 

First Repair 

There are records of a repair required on the tank in 

1965. In an initial letter dated 12 January 1965 

(attachment 1), it was stated that an inspection in the 

tank conducted on 5 January revealed cracks in the 

manway boss and nozzle P4. In a follow-up letter 

dated 14 January 1965 (attachment 2), a repair to 

Nozzle P4 was conducted, there is no information 

relative to why the repair was required. There is also 

no further mention of the manway boss crack. There is 

a hand written note on the 12 January note stating that 

the manway boss must have just been a surface scratch 

since it was never mentioned again. However, 

following the weld repair on P4, the tank was 

pneumatically pressure tested to 125% of operating 

pressure, then cleaned with acid, flushed, dried and reinspected. The dye penetrant re-inspection revealed 

that nozzles P1 and P2 had cracks in the welds, and 

nozzle P3 with a surface scratch. Some additional 

analysis was conducted by Rocketdyne and on January 

20, 1965 the decision was made to make the welds on 

all seven nozzles P1 thru P7 full penetration welds. 
5 

The weld repairs were completed and documented 

following proper procedures for an ASME code 

stamped vessel. The repair is documented in 

attachment 3, the U1A Report for the vessel after 

completing the repair. Following these repairs, no 

other indications were noted in this 1965 inspection. 

The qualifying weld procedure report for the weld 

repair is listed in Table 5 [16]. 

The Military Specification MIL E-22200 2A referred 

to in Table 5 is entitled Electrodes, Welding, Covered 

(Austenitic Chromium-Nickel Steel) [17] and applies 

to the shielded metal arc welding of austenitic stainless 

steels. For this repair weld, type E-308 ELC-16 

electrodes were used. These electrodes are intended 

for general welding of 18Cr, 8Ni extra low carbon 

corrosion-resistant steels and are used where 

intergranular corrosion should be limited but where 

corrosive conditions are not the most severe. Although 

this welding electrode is not of the same composition 

as that of the 347 plate, it is permissible to use this 

electrode for 347 stainless steel. The recommended 

ferrite number shall be from 4 to 10 as determined by 

AWS A5.4 [11]. 

The reported values of 145 and 26 for the 1/8” and 165 

and 26 for the 5/32” appear to the ultimate tensile 

strength (ksi) and Charpy V-notch (ft-lbs) for the weld 

metal at -320°F. The 1/8” and 5/32” indicate the weld 

rod diameter used. Due to the poor quality of the 

report, the description of these numbers cannot be 

determined. If we assume the same Special Certified 

low temperature property tests required at the time of 

manufacture, then the weld metal meets code for 

strength and fracture toughness at -320°F. 

Based upon the FN determined from AWS A5.9 [11], 

the undiluted 308L weld metal had a FN of 9 based 

upon the nominal mid-range chemistry of the E308L 

weld rod. Weld metal dilution with the plate material 

will usually result in somewhat lower ferrite content 

than the undiluted weld metal depending upon the 

amount of dilution and the composition of the base 

metal. These values are within the recommended 4-

9% ferrite range for reducing crack susceptibility in 

347 weld metal [13]. 

No post-weld stress relief was used following the weld 

repair. Therefore, the residual stresses in the vicinity 

of the weld repair can be high and approach the yield 

point of the surrounding base metal and heat affected 

zone [14, 25]. 

Second Repair 

A dye penetrant inspection was conducted after 

moving the tank from California to NASA Langley 

dated 28 July, 1986 (see attachment 4). The dye 

penetrant inspection report document states that the 8 

circumferential welds, the 6 insert welds (with the 

seventh being inaccessible), the penetration at the entry 

hole, and all other penetrations of the shell were 

inspected. The liquid outlet at the bottom of the tank 

was listed as being inaccessible for inspection. The 

results from this inspection revealed two incomplete 

fusion indications in the top of the vessel, one 

incomplete fusion in the bottom of the vessel, and one 

crack-like indication at the bottom of the vessel. In a 

note, the inspector stated that the diffusers on the 

manifold were loose and that the interconnecting rods 

were worn. These items were addressed in subsequent 

repair work – however they did not require any work 

on the pressure shell itself. 

On 5 December of 1986 a second dye penetrant 

inspection inside the tank was completed following 

light grinding of the four indications revealed in the 

earlier inspection. The inspection notes that the light 

grinding (less than 1/32”) had removed the four 

indictations (see attachment 5). Following this 

inspection, a memo dated 16 January 1987 from the 

Fracture Mechanics Engineering Section stated that the 

LOX run tank had been judged safe for continued 

operation. Following this memo, the LOX tank 

installation into the 8’HTT began. The LOX run tank 

was not used for facility operation, including shake 

down testing, until July of 1992. 

NASA records show that on 31 August, 1989 Chicago 

Bridge and Iron (CB&I) the Prime contractor for the 

8’HTT modification project submitted the completed 

R-1 forms for welded modifications made inside the 

LOX run tank. The documentation for the weld 

modifications is covered in Section C of this report. 

The modifications consisted of welding plugs into 

nozzles P-1 through P-6. The plugs were used to seal 

off each of the shell penetrations which were not 

required for operation at the 8’HTT. The lower nozzle 

P7 is used for a differential pressure measurement 

which measures the height of the liquid in the run tank. 

On 14 June 1990 thickness measurements were taken 

in the areas in which the four indications found in 1986 

had been ground out. The measurements showed that 

wall thicknesses in the repair area were greater than 

4.71 inches. 

Following the NASA inpections and weld 

modifications completed by CB&I, a hydrostatic tank 

was conducted on the LOX tank. On 30 June 1990 the 

hydrostatic test report documents that the LOX run 

tank was hydrostatically tested to 1.5 times maximum 

operating pressure (1.5*2290 = 3435 psig). The 

hydrotest report document is shown in attachment 6. 

Following the successful hydrostatic test, the LOX run 

tank cleaning process (required for liquid oxygen 

service) was initiated. 
6 

Third Repair 

The documentation then shows that as a check of the 

LOX cleaning, a portion of the nitrogen distribution 

manifold internal to the run tank was inspected and 

poor quality welds were found in the distribution 

manifold. In October of 1990, the weld review board 

met, and determined that the large diameter tee just 

inside the run tank should be cut, the distribution 

manifold lowered to the bottom of the tank, and weld 

repairs on this piping be completed. The weld review 

board meeting minutes are provided in attachement 11. 

The distribution manifold was repaired to meet ASME 

code requirements. 

Records show (attachment 7) that also beginning in 

October of 1990, dye penetrant inspections inside the 

tank of all nozzles were conducted and documented in 

examination reports. These dye penetrant inspection 

reports are provided in attachment 8. In report number 

109 dated 22 October 1990, six cracks were found 

around the perimeter of nozzle “V”. It is not clear why 

these cracks were not found in the 1986 inspections. 

There were no operation cycles of the LOX tank over 

this period. The only pressurization occurred during 

the hydrostatic test. The process to grind out these 

cracks was initiated using dye penetrant inspections to 

insure that the entire crack had been ground out prior 

to beginning repair. 

Dye penetrant inspection reports 110, 111, 112, 113, 

and 114 show that all other nozzles into the run tank 

were also inspected in December of 1990. Please 

recall that nozzles P1 through P7 had plugs welded 

into them by CB&I and had previously been inspected, 

therefore they were not re-inspected. Defects were 

found in nozzles M, A. No defects were found in 

nozzles N, D, or P. The defects found in nozzles M and 

A were removed via grinding and no weld repair was 

required. The records show that after grinding, each 

nozzle was re-inspected and no indications were found. 

For the weld repair around nozzle V, all repair work 

was conducted and documented properly and inspected 

by the National Board. The welding procedure is 

documented on 28 November 1990 and the completed 

weld repair is and documented in a Report of Welded 

Repair submitted on 1 May 1991. The dye penetrant 

inspection record number 109 shows that following the 

completed repair, nozzle V was dye penetrant 

inspected on 1 February 1991 and no defects were 

found. The lox run tank oxygen cleaning resumed and 

the tank was cleaned for oxygen service. The LOX run 

tank was certified clean on 29 February 1992. 

The welding procedure for the repair of the cracks 

found in the weld surrounding nozzle “V” and for 

welding the plugs into nozzles P-1 through P-6 was gas 

tungsten arc welding (GTAW) using bare ER316L 

stainless steel welding electrodes. Although this weld 

process differs from that used in the manufacture 

(shielded metal arc welding), it is an acceptable 

alternative. 

Filler metal ER316L has a nominal composition (wt. 

%) of 19Cr, 12.5NI and 2.5Mo with low carbon (0.03 

% max). The low carbon in this filler metal reduces the 

possibility of intergranular chromium carbide 

precipitation and thereby increases the resistance to 

intergranular corrosion without the use of stabilizers 

such as columbium or titanium. The molybdenum 

provides enhanced creep resistance at elevated 

temperatures. The filler metal is used primarily for 

welding low-carbon molybdenum-bearing austenitic 

alloys. The ER316L filler metal is often used for 

service at cryogenic temperatures because of its high 

toughness due to its high Ni content and it does not 

crack at very low ferrite levels [13, 18]. 

The weld metal tensile strength at ambient temperature 

averaged 83 ksi while the Charpy-V notch fracture 

toughness at -320°F averaged 31 ft-lbs with a lateral 

expansion opposite the notch opening exceeding the 

0.015 inch minimum allowable threshold. The 

reported 10% shear fracture area is a qualitative 

measurement that estimates the relative amount of 

shear fracture versus that of cleavage or brittle fracture. 

Lateral expansion is also used to indicate the relative 

amount of ductile tearing during fracture. Lateral 

expansion is reported as the increase in width of the 

broken specimen over the unbroken specimen width. 

A brittle specimen will not have significant tearing at 

the sides and the width of the specimen will be very 

close to that of the unbroken specimen. A tough 

specimen will have extensive shear lips that deform the 

specimen into the width direction. 

Based upon the FN determined from AWS A5.9 [18], 

the undiluted 316L weld metal has a FN of 8 based 

upon the nominal mid-range chemistry of the ER316L 

weld rod. Further dilution between the 347 plate metal 

and the 316L weld metal would somewhat lower the 

ferrite content. This value is within the recommended 

4-8 percent ferrite range for reducing crack 

susceptibility in the weld metal [13]. The weld repair 

qualification records for the tensile and Charpy Vnotch of the 316L weld metal meet the current 

requirements for strength, ductility, and fracture 

toughness for -320°F service. 

No post-weld stress relief was used following the weld 

repair surrounding nozzle V (Figure 5). Therefore, the 

residual stresses in the vicinity of the weld repair can 

be high and approach the yield point of the 

surrounding base metal and heat affected zone [14, 

25]. 
7 

LOX Tank Stresses 

The LOX tank is an 8,000 gallon capacity, 2,290 

maximum operating pressure, vacuum jacketed, vessel. 

The LOX tank is maintained in a cryogenic condition 

at all times, full or at least partially full of liquid 

oxygen. The run tank sits at atmospheric pressure at 

all times except during a facility run. The LOX tank is 

pressurized during operation using gaseous nitrogen. 

The gaseous nitrogen enters at the top of the tank at 

520 degrees Rankine (60 degrees Fahrenheit). During 

operation, a pressure control valve is used to ramp the 

run tank pressure up to 2,250 psig, then hold this 

pressure level for the duration of the run, typically one 

minute. Immediately after each run, the LOX tank 

nitrogen gas is vented and the tank returns to 

atmospheric pressure. A plot showing the 

pressurization cycle for a typical run is shown in the 

attached Figure 6. 

Pressure Stresses 

Two finite element models were constructed to model 

the pressure stresses in the LOX tank during operation. 

The first model presented was constructed using Pro 

Engineer [19] and analyzed using Mechanica [20]. 

The second model was constructed using Patran [21] 

and analyzed using Abaqus [22]. The details of each 

analysis are presented separately followed by a 

discussion of the modeling results. 

Mechanica Model 

Three structural Finite Element Models were 

developed using Pro Engineer. For the first two 

models, top and bottom, a spherical coordinate system 

was used (R, θ, φ). The constraints along the shell 

boundary are θ and φ and the models have constant 

pressures on the inside surfaces of the spherical shell 

and the penetrations. The third model, which is a slice 

of the shell, uses a cylindrical coordinate system (R, θ, 

Z) with constraints of θ along the shell boundary and a 

vertical constraint Z on the lower surface of the 

support bracket. The third model uses a variable 

pressure that is a function of the height to obtain a 

vertical reaction due to the weight of the LOX and 

shell. A negative pressure is applied to the surface 

where a pipe would connect to a penetration on the 

outside to account for the open area of the penetration. 

Figure 7 is a solid model representation of the ProEngineer model used to develop the Mechanica finite 

element models with the nozzles labeled. 

The top, side and bottom of the tank were analyzed 

with refined models to evaluate the stress 

concentrations from the nozzles and the support 

brackets on the side of the tank during pressurization. 

The maximum stress of 57.7 ksi was found to be at the 

interior of Nozzle N during the pressurization of the 

tank, as shown in Figure 8, and is the only one of the 

three models shown herein. 

Abaqus Model 

Three finite element models of varying discretization 

were created to analyze the LOX tank. Models with a 

coarse mesh having 11,062 8-node hexahedron 

elements, a medium mesh having 11,062 20-node 

hexahedron elements and a fine mesh having 37,388 

20-node hexahedron elements are shown in Figure 9ac, respectively. In Figure 9 the nozzles are blue, the 

welds between the nozzles and the tank are red, the 

section of the tank from the original model is green, 

and the rest of the tank hemisphere is yellow. 

The von Mises equivalent stress distribution near 

Nozzles V, D and N is shown in Figure 10. The 

maximum von Mises stress was located at the same 

location (near Nozzle N and shown in the figures) in 

all cases and had a value of 40.86 ksi, 45.24 ksi and 

45.66 ksi for the Coarse, Medium and Fine mesh 

models, respectively. The difference between the 

values obtained for the Medium and Fine mesh models 

is less than 1%, indicating that the discretization is 

converged. 

Discussion of Pressure Stress Models 

The maximum value of von Mises stress (45.66 ksi) 

predicted by the Abaqus model is lower than the 

maximum value (57.7 ksi) predicted by the Mechanica 

model. This difference is due to several factors 

including geometric differences in the two models, as 

shown in Figure 11. In Figure 11, a cross-section of 

the tank near the manhole cover is shown along with 

the design drawing from which it was created. The 

solid and dashed lines represent the local configuration 

of the Mechanica and Abaqus models, respectively. 

This geometry difference causes a shift in the location 

of the corner at which the curved tank wall and the flat 

surface of the manhole cap intersect. In the Mechanica 

model, this corner intersects the edge of the hole of 

Nozzle N, as shown in Figure 12. In the Abaqus 

model as in the design drawings, this corner does not 

intersect the edge of the hole of Nozzle N, as shown in 

Figure 12. In physical structures, stress concentrations 

occur at the edges of holes and at the corners. Because 

the hole and the corner overlap in the Mechanica 

model, both stress concentrations occur at the same 

location resulting in a magnified maximum stress. In 
8 

the Abaqus model, the stress concentrations do not 

occur at the same place and the maximum stress is not 

magnified. 

Thermal Stresses 

The thermal stresses are calculated by computing a 

conservative temperature profile. An analysis was 

conducted in which the nitrogen mass flow rate into 

the LOX tank during an operational cycle was 

estimated from the depressurization rate of the 

dedicated nitrogen field providing the pressurization 

gas. The mass flow rate of nitrogen was used to obtain 

the pipe flow Reynolds number, which was in turn 

used to estimate the convective heat transfer 

coefficient on the inside surface of nozzle “V” (the 

nitrogen inflow nozzle). For the transient thermal 

analysis it was assumed that the entire structure was 

initially at the temperature of the liquid oxygen (-300 

°F = 160 °R), while the surface of the nitrogen inlet 

nozzle (nozzle V) was convectively heated by the 

gaseous nitrogen flowing through nozzle “V”. A 

summary of the calculation method used for the 

transient heat transfer analysis is on final in the 

document archives at the 8’HTT. From this 

information,transient temperature profiles were 

calculated. The thermal stresses were calculated using 

finite element models developed in Patran [21]. 

Nastran [23] was used to compute the resulting thermal 

stresses. 

Three thermal stress finite element models were 

developed. The structural models were not used 

because many more elements are needed through the 

thickness of the tank and nozzle walls to capture the 

thermal stresses. The model was developed in a 

cylindrical coordinate system (R, θ, Z) and is wedge 

shaped with symmetry in the θ direction. The heat 

transfer coeffient versus run time is shown on Figure 

13. The heat transfer coefficient was calculated using 

an equation from Kays and Crawford for a circular 

tube with fully developed velocity and temperature 

profiles, constant heat rate and Reynolds’s number < 

105

 [24]. The physics of thermal stress analysis is 

complicated because the thermal profile is not fully 

developed, the Reynolds’s number is one to two orders 

of magnitude larger, the geometry has been simplified 

and backside natural convection and radiation have 

been ignored. In order to insure that the analysis was 

still conservative, a factor of ten was applied to the 

heat transfer coefficient. The modified values of the 

heat transfer coefficient used in the analysis are shown 

on Figure 14. 

For the inside surface of the nozzle V, the highest 

thermal stresses occured at a time of 62 seconds. The 

resulting temperature profile is shown in Figure 15. 

The thermal stresses are shown in Figure 16. The 

maximum thermal von Mises stress is 107 ksi. For the 

backside of the nozzle on the inside of the shell, the 

highest thermal stresses occur at a time of 245 seconds. 

The resulting temperature profile is shown in Figure 

17. The thermal stresses are shown on Figure 18. The 

maximum thermal von Mises stress in this region is 33 

ksi. 

Damage Tolerance Model 

The area of interest for a damage tolerance analysis is 

the top of the spherical LOX tank where a forged cover 

and three nozzles are welded to the tank. When 

pressurized for operation, the top of the tank witnesses 

the highest pressure and thermal stresses. The pressure 

stresses are high in this region because of the nozzle 

geometric details. The thermal stresses are highest 

because of the introduction of gaseous nitrogen to 

pressurize the tank through Nozzle V. 

The input to the damage tolerance model is broken 

down into separate sections on geometry, residual 

stresses, applied stresses, initial crack size, and 

material properties. 

Geometric Input 

The forged top of the tank is nearly flat, so the 

contributions of out-of-plane cracking will be 

minimized, such that the simplification of the crack 

growth model to a think-plate solution is reasonable 

(and 19). The geometric model was constructed to 

simulate cracking between the two Nozzles V and N 

(Figures 19 and 20). The geometric dimensions 

(inches) used are: t = 4.73, W = 7.0, ai = 0.5, and a/c = 

1.0. 

Stress Distribution 

Three distinct stress distributions affect the damage 

tolerance analyses at the top portion of the LOX tank. 

The LOX tank is pressurized with room-temperature 

gaseous nitrogen during operation leading to pressure 

and thermal stresses. The construction and subsequent 

repair of the LOX tank left residual stress fields around 

the welments. In the damage tolerance analysis, the 

pressure (Figures 8, 10 and 21) and thermal stresses 

(Figures 16, 18 and 22) are superimposed with the 

residual stresses from construction (Figure 2) and 

repair welding (Figure 3) where the x and y 

coordinates are defined in Figure 23. 
9 

Pressure stress 

The pressure stresses were computed using Mechanica 

and Abaqus based on the pressure cycle shown in 

Figure 6. The highest pressure stresses on the tank 

were computed using the Mechanica model and were 

located in the vicinity of Nozzles N and V as shown in 

Figure 8. Therefore, the pressure stresses used in the 

damage tolerance analyses are: 

Principal pressure stresses are 16 - 22 ksi between 

the Nozzles (Figure 21) 

Assume tank stresses are approximately zero 

between runs, i.e. load ratio, R = 0 (Figure 6) 

Pressure at run = 2250 psi 

Pressure between runs = 5 psi 

Thermal stress 

Nozzle V is the nitrogen fill nozzle for pressurizing the 

tank and therefore experiences the highest thermal 

stresses. A thermal stress analysis was performed 

using convective heat transfer into the nozzle based on 

the thermal cycle the tank undergoes, as shown in 

Figure 14. The thermal stresses used in the damage 

tolerance analyses are: 

Local thermal stresses are a constant 3.74 ksi 

between Nozzles V and N (Figure 22) 

Residual stress 

The construction of the LOX tank involved welding 

several sections together to form the sphere and then 

welding a forged cover and several nozzles to the top 

of the tank (Figure 1). The welds are all butt-welds as 

shown in Figure 1. It is assumed that the residual 

stresses from the initial construction of the tank have 

the profile of tension at the tank surfaces and 

compression in the interior (shown in Figures 2 and 3) 

based on information presented in the Welding 

Handbook [12-14] and ASTM STP 776 [25]. 

The tank was repaired in 1990 as reported in contract 

18367C [26]. The inspection report details the weld 

repairs around Nozzle V as shown in Figure 5. The 

repair welds were not heat treated, and based on 

ASTM STP 776 [25] the residual stresses are tensile 

along the weld-line and their variation along the length 

(x-direction) is presented in the bottom part of Figure 

23. The crack growth analyses are performed by 

adding the residual stresses as steady stresses 

(increasing the stress ratio only), since this is an 

accurate representation of how the residual stresses 

would impact crack growth. 

Assume residual stresses from manufacture are 15% 

of room temperature yield 

σres = 5.5 ksi 

Profile is tension at the surfaces and compression at 

the depth (Figure 23) 

Assume residual stresses from the weld repair to be 

equal to room temperature yield 

σres = 35 ksi 

Profile is tension at the weld repair and transitions 

linearly to zero stress at a dimension of 3⁄4 of the 

thickness (3.55 inches) in both the thickness and 

width directions. 

Initial Crack Size 

The initial crack size for the damage tolerance analysis 

is based on the method of NDI used in the inspection 

report. For the repair procedure, dye penetrant was 

used and based on NASA STD-5009 [27] an initial 

corner crack of 0.1 inches in radius should be used for 

the damage tolerance analyses. Additionally, in this 

case, a dye penetrant inspection revealed a 1⁄2 inch 

crack in the weldment of Nozzle V (Figure 5). The 

crack was ground out and a weld repair was 

performed. A subsequent penetrant inspection did not 

reveal any indications at the weld repair. For this 

analysis, it was presumed that the repair did NOT 

remove the crack from the LOX tank. Therefore, the 

initial crack sizes chosen were 0.5 and 0.1 inch radius 

corner cracks initiating at the weld repair. 

Material Properties 

The LOX tank is primarily constructed of 4.73 inch 

thick 347 Stainless Steel plate welded together using 

347 SS wire. The top section, where the nozzles are 

located, is 10.25 inch thick 347 SS forging welded to 

the nozzles and sphere as shown in Figure 1. All crack 

growth considered in the damage tolerance analysis is 

contained within the weld connecting the forging to the 

sphere. Therefore, the material properties used in the 

analysis are representative of 347 Stainless Steel 

weldments at a temperature of -320 F. Table 2 

contains the tensile data was generated during the 

initial qualification of the tank [2], data for welds in 

this alloy was found in the literature [28], and data was 

generated for the repair welds (Table 3) . Fracture 

toughness was calculated using Rolfe, Novak and 

Barson [29] from Charpy data, such that 















 = 5 − 0.05

YS

VN

Ic YS

C K

σ

σ (1) 

where σYS is the yield stress at temperature for the 

material condition (parent or weld). The Stainless 

Steel 347 has a yield stress (parent material) of 36.5 ksi 
10 

at room temperature and 75.1 ksi at -325F (Tables 1, 2 

and 3). The properties used in the analysis are the 

average yield and ultimate stresses from the 

construction weldment and the average fracture 

toughness of the repair weld data. The fracture 

toughness values calculated from the Charpy data are 

similar to those found in the NASGRO database [1] 

and ASME code [3] for other austenitic stainless steels. 

Therefore, the material properties input into NASGRO 

for the damage tolerance analyses are: 

347 Stainless Steel Fracture Toughness and 

Yield/Ultimate data for weldments 

Yield = 48.9 ksi 

Ultimate = 172.2 ksi 

Fracture toughness, KIC = 83.0 ksi in1/2

Thickness dependent fracture toughness, KIe = 

116.0 ksi in1/2

304 Stainless Steel da/dN curve fit (Fits the high R 

cryo data and bounds weld data), units are US 

Costumary (Figure 24) 

NASGRO Coefficients: UTS = 172.2, Yield = 48.9, 

K1e = 116, K1c = 83, Ak = 0.75, Bk = 0.5, a0 = 

0.0015, Kth(s)/Kth(l) = 0.2, C = 6x10-10, n = 3.0, p 

= 0.25, q = 0.25, DK1 = 1.84, Cth = 0.63, Cth- = 

0.1, Alpha = 2.5, Smax/Flow = 0.3, Cth = 0 

throughout. 

Damage Tolerance Analyses 

The crack growth analyses were performed using 

NASGRO 4.22. The geometry chosen was CC09; the 

material properties were for 304 Stainless Steel at 

room temperature with modified values for yield 

stress, ultimate stress, KIC and KIe; a nonlinear stress 

profile in both the x (thickness) and y (length) 

directions were developed to include pressure, thermal 

and residual stresses; and a corner crack of 0.5 inches 

in radius that was detected and repaired during the 

inspection of the tank in 1990 [26]. The crack growth 

analysis predicts a critical crack depth of 1.17 inches, 

approximately 1⁄4 of the wall thickness, where the stress 

intensity factor in the depth direction exceeds KIe. KIe, 

as opposed to KIC, is the controlling factor in this case 

because the critical crack is a part-through crack [30]. 

The cycles to failure are 44,721, where one cycle 

equates to one pressurization cycle of the LOX tank. 

Figure 25 depicts the crack size relative to the 

geometry and Figure 26 shows the remaining cycles to 

failure for any given crack size from this analysis. 

Running the same analysis with an initial crack size of 

0.1 inches resulted in 131,923 cycles to failure. 

All subsequent sensitivity studies are performed using 

the initial crack size of 0.5 inches. The fracture 

toughness was varied from the baseline (83 ksi in1/2) to 

simulate a very low toughness weld (KIC = 40 ksi in1/2). 

The apparent fracture toughness was held to be 1.4 

times the plane strain fracture toughness based on the 

relationship between the two values for the base 

material, i.e. KIe = 1.4 KIC. This relationship between 

KIe and KIC is common for austenitic stainless steels 

based on data found in the NASGRO material database 

[1]. Variation of the fracture toughness and the 

resultant cycles to failure and critical crack size is 

plotted in Figure 27. The variation is similar to an Sshaped curve with a large linear portion. If the 

toughness drops to 60 ksi in1/2 then the predicted cycles 

to failure will be approximately 23,748 cycles, whereas 

a drop in toughness to 50 ksi in1/2 yields a failure time 

of 9,764 cycles compared to the baseline of 44,721 

cycles. 

Finally, the residual stresses due to the repair are 

modified based on observations from the nuclear 

industry to be the flow stress (62.5 ksi) locally and 

linearly decay to zero at 3⁄4 of the thickness in both the 

thickness and width directions. The result of changing 

the repair residual stress from 35 ksi to 62.5 ksi 

decreases the cycles to failure from 44,721 to 20,647. 

Summary 

The damage tolerance analysis of the LOX tank 

focused on crack growth between Nozzles V and N 

because of the following assumptions: 

1. Nozzle V experiences the greatest thermal 

stresses during tank pressurization. 

2. Nozzle N experiences the greatest pressure 

stresses during tank pressurization. 

3. A crack was found at Nozzle V during the dye 

penetrant inspection and weld repaired. 

4. The weld repair at Nozzle V induced a 

residual stress that would assist crack 

propagation. 

The geometry of the LOX tank section between 

Nozzles V and N was modeled using a thick plate 4.73 

inches thick by 7 inches wide to encompass both the 

distance between the Nozzles and the weldments on 

each side connecting the Nozzles to the tank. 

Conservatisms built into the geometry assumption are: 

1. The geometric model does not account for 

any stiffness of the surrounding tank or 

Nozzles. 

The stress distribution applied to the crack growth 

model was nonlinear in both the thickness and width 

directions to accurately describe the stress distributions 

of the tank pressure, thermal stress and residual 

stresses from both construction and repair welding. 

Conservatisms built into the stress profiles are: 
11 

1. The thermal stresses have a factor of ten 

applied to the heat transfer coefficient, 

increasing the thermal stress magnitude above 

what would be expected. 

2. The thermal stress is assumed constant 

between the Nozzles, when it decays rapidly 

from Nozzle V. 

3. The residual stresses from construction are 

considered to be 15% of the room temperature 

yield stress (due to post-construction stressrelief), when the literature suggests the 

residual stresses should be closer to zero. 

4. The residual stresses from the repair weld are 

tensile through the entire weld nugget and 

heat-affected zone and then drop to zero stress 

instead of a beneficial compressive stress. 

The mechanical material properties were taken from 

data generated during the construction and repair of the 

tank. The crack growth rate data was taken from the 

NASGRO material database to fit the high stress ratio 

data of 304 stainless steel at cryogenic temperature. 

Conservatisms in the material properties are: 

1. The crack growth rate curve is nearly a factor 

of two more conservative than the NASGRO 

standard curve and is an upper bound for all 

data. 

The initial crack size is assumed to be 0.5 inches, 

which corresponds to a crack that was detected and 

repaired in 1990. The NASA standard for damage 

tolerance inspection, NASA STD-5009 recommends 

an initial crack size of 0.1 inches in radius based on 

inspection method, in this case fluorescent penetrant, 

and geometry. Conservatisms in the initial crack size 

are: 

1. The initial crack size of 0.5 inches used in the 

analysis is five times that recommended by 

NASA STD-5009. 

A summary of the all the crack growth analyses 

performed based on different assumptions is presented 

in Table 6. 

The damage tolerance analysis performed using 

NASGRO 4.22 predicts that a one-half inch corner 

crack would fracture the ligament between Nozzle V 

towards Nozzle N in 44,721 pressurization cycles. The 

NASA STD-5003 requires pressure vessels that have a 

catastrophic failure mode, such as the LOX tank, have 

an inspection interval of one-fourth the total damage 

tolerance lifetime. Therefore, the NASA standards 

would recommend a maximum time between 

inspections of 10,000 cycles. 

Taking into account the current state of the tank, 

reducing the recommended inspection interval to 1,000 

pressurization cycles would increase the factor of 

safety on the damage tolerance analysis to forty. To 

understand what a recommended 1,000 cycle 

inspection program would provide as a factor of safety, 

parametric studies were performed to evaluate the 

sensitivity of the damage tolerance analysis to the 

material fracture toughness because of uncertainties in 

the welding process. Presuming the factor of four on 

lifetimes (i.e. the total life is 4,000 cycles), the fracture 

toughness of the weld could be as low as 

approximately fifty percent (~ 45 ksi in1/2) of the 

reported value. Including the conservatisms built into 

the base analysis predicting 44,721 cycles to failure, an 

inspection interval of 1,000 cycles is very 

conservative. 

Acknowledgements 

The authors would like to thank J. Brady, E. 

Glaessgen, I. Raju, J. Newman, and the NASA 

Langley Pressure Systems Committee for their 

valuable insight and guidance in these analyses. 

Reference 

[1] NASGRO Version 4.22, nasgro.swri.org, managed 

by Southwest Research Institute, San Antonio, TX. 

[2]Material & Welder’s Record, Southwest Welding & 

MFG. Co., Rocketdyne Contract R-943-C-F-972, 1959 

[3] ASME-BPVC, Section VIII, Division 1, Rules for 

Construction of Pressure Vessels – Division 1, 2004. 

[4] ASME-BPVC, Section IX, Qualification Standard 

for Welding and Brazing Procedures, Welders, 

Brazers, and Welding and Brazing Operators, 2004. 

[5] Source Book on Stainless Steels, Metals 

Engineering Institute, ASM International, Metals Park, 

OH, 1979. 

[6] Metals Handbook Tenth Ed., Volume 1: Properties 

and Selection: Irons, Steels, and High-Performance 

Alloys, ASM International, Metals Park, OH, 1990, pp. 

841-907. 

[7] ASME BPVC, Section II, Part A, Materials - 

Ferrous Material Specifications, 2004. 

[8] Thomas, Jr., R.D. and Messler, Jr., R. W. 

“Welding Type 347 Stainless Steel-An Interpretive 

Report,” Welding Research Council Bulletin 421, May 

1997. 

[9] Stainless Steels Types 321, 347 and 348, Allegheny 

Technologies Inc., Allegheny Ludlum Technical Data 

Blue Sheet, 2003. 

[10] Drawing No. D-9093: Nozzle N Blended Weld & 

Forging Proposals, Southwest Welding & Mfg. Co., 

dated 3-25-1959. 

[11] ANSI/AWS A5.4-92, Specification for Stainless 

Steel Electrodes for Shielded Metal Arc Welding, 

American Welding Society, 1992. 
12 

[12] Welding Handbook Ninth Ed., Volume 2: 

Welding Procecesss, Part 1, American Welding 

Society, Miami, FL, 2004. 

[13] Welding Handbook Eighth Ed., Volume 4: 

Materials and Applications – Part 2, American 

Welding Society, Miami, FL, 1998. 

[14] Welding Handbook Ninth Ed., Volume 1: 

Welding Science & Technology, American Welding 

Society, Miami, FL, 2004. 

[15] ASME-BPVC, Section II, Part C, MaterialsSpecifications for Welding Rods, Electrodes, 2004. 

[16] Welding Procedure Specification No. G-4274 # 1, 

dated 1-17-65. 

[17] Military Specification MIL-E-22200 / 2D 

Electrodes, Welding, Covered (Austenitic ChromiumNickel Steel), September 1986. 

[18] ANSI/AWS A5.9-93, Specification for Bare 

Stainless Steel Welding Electrodes and Rods, 

American Welding Society, 1993. 

[19] Pro/ENGINEER Version 2.0, www.ptc.com, The 

Product Development Company, 2005. 

[20] Pro/MECHANICA, www.ptc.com, The Product 

Development Company, 2005. 

[21] Patran, www.mscsoftware.com, MSC Software, 

2005. 

[22] Abaqus, www.abaqus.com, Dassault Systémes, 

2005. 

[23] Nastran, www.mscsoftware.com, MSC Software, 

2005 

[24] Kays, W. M. and Crawford, M.E, Convective 

Heat and Mass Transfer; McGraw-Hill: New York, 

USA, 1980. 

[25] Residual Stress Effects in Fatigue, ASTM STP 

776, ASTM, 1981. 

[26] Report of Welded Repair for LOX Run Tank, 

NASA Contract No. 18367C, 1991. 

[27] Nondestructive Evaluation Requirements for 

Fracture Control Programs, NASA STD 5009, 2003. 

[28] Damage Tolerant Design Handbook, Volume 1, 

December 1983. 

[29] J. M. Barsom and S. T. Rolfe, Fracture & Fatigue 

Control in Structures, Applications of Fracture 

Mechanics, 2nd ed., Prentice-Hall, Inc., 1987, pp. 159-

188. 

[30] Forman, R.G., and Henkener, J.A., “An 

Evaluation of the Fatigue Crack Growth and Fracture 

Toughness Properties of Beryllium-Copper Alloy 

CDA172,” NASA Technical Memorandum 102166, 

1990. 
13 

Figure 1. Schematic of the Liquid Oxygen (LOX) tank construction. Black lines indicate welds.
14 

Figure 2. Drawing showing residual stress orientation in a conventional welded plate. [7, 8]. 

Figure 3. Schematic showing residual stress distribution in a conventional welded plate (indicated directions are 

with respect to the weld) and level of residual stresses. [7, 8]. 
15 

Figure 4. Schematic showing weld modifications of the plugs to the LOX tank [26]. 
16 

Figure 5. Repair drawing for Nozzle V of the LOX tank [26]. 

Figure 6. Pressurization profile of the LOX tank. 
17 

Nozzle D 

Nozzle V 

Nozzle M Nozzle N 

Nozzle A Nozzle B 

Support Bracket 

Plug 

Figure 7. Pro-Engineer model of the LOX tank. 
18 

Figure 8. Von-Mises stresses for the inside of the top of the LOX tank modeled using Mechanica.

(a)

Nozzle D

Nozzle N

Nozzle V

 (b)

Nozzle D

Nozzle N

Nozzle V
19 

(c)

Nozzle D

Nozzle N

Nozzle V

Figure 9. Patran models with (a) course, (b) medium, and (c) fine mesh refinement of the upper portion of the LOX 

tank. 

(a)

Maximum 

Equivalent 

Stress

(Nozzle N)

40.86 ksi.

(b)

Maximum 

Equivalent 

Stress

(Nozzle N)

45.24 ksi.
20 

(c)

Maximum 

Equivalent 

Stress

(Nozzle N)

45.66 ksi.

Figure 10. Von-Mises stress plots from Abaqus with (a) course, (b) medium, and (c) fine mesh refinement of the 

upper portion of the LOX tank. 

Figure 11. Comparison of Mechanica and Abaqus mesh details. 

(a)

Corner

Hole of Nozzle N

 (b)

Corner

Hole of Nozzle N

Figure 12. Comparison of (a) Mechanica and (b) Abaqus geomtry details at Nozzle N. 
21 

Convective Heat Transfer Coefficient 

in LOX Run Tank Nozzle "V"

-0.0200

0.0000

0.0200

0.0400

0.0600

0.0800

0.1000

0.1200

0.1400

0.1600

0.1800

0.00 50.00 100.00 150.00 200.00 250.00 300.00

Time (seconds)

BTU/Ft^2 - F - s)

Figure 13. Heat transfer coefficient versus time during LOX tank operation at Nozzle V. 

10x Convective Heat Transfer Coefficient 

in LOX Run Tank Nozzle "V"

-0.2000

0.0000

0.2000

0.4000

0.6000

0.8000

1.0000

1.2000

1.4000

1.6000

1.8000

0.00 50.00 100.00 150.00 200.00 250.00 300.00

Time (seconds)

BTU/Ft^2 - F - s)

Figure 14. Conservative estimate of the heat transfer coefficient versus time at Nozzle V used for the Nastran 

analyses. 
22 

Figure 15. Nozzle V temperature gradient, Ro, at 62 seconds. 

Figure 16. Nozzle V von Mises thermal stress gradient at 62 seconds. 
23 

Figure 17. Nozzle V temperature gradient, Ro, at 245 seconds. 

Figure 18. Nozzle V von Mises thermal stress gradient at 245 seconds. 
24 

 

CL 

Nozzle V

CL 

Nozzle N 

Crack 

Growth 

Direction 

Welds connecting 

Nozzles to Tank 

Weld repair 

4.73” 

12” 

Nozzle ID to ID 

7.0” 

3.75” 

4.75” 

Figure 19. Diagram of the geometry assumed for the crack growth model (Not to Scale). 

Figure 20. Corner crack in a thick plate model used to simulate cracking in a nonlinear stress field from Nozzle V to 

Nozzle N. 
25 

Figure 21. Principal pressure stress used for the damage tolerance analysis. 

Figure 22. Thermal stress cycle of Nozzle V in the hoop direction (z) at 245 seconds. 

Region in DT 

analysis 

Nozzle V ID 

Repair Weld 

Area 
26 

Crack 

Growth 

Direction x 

y 

 

Length (inches)

02468

Stress (ksi)

0

10

20

30

40

Thermal

Repair Residual

 Figure 23. Residual stress profiles in the vicinity of Nozzle V. Thickness, y (inches) 012345

Stress (ksi)

-10

0

40302010

Construction ResidualRepair Residual
27 

∆K (ksi in1/2)

10 100 20 40 60 80 200

10-7

10-6

10-5

10-4

10-3

10-2

NASGRO R=0.1 (304SS, -322F)

NASGRO R=0.5 (304SS, -322F)

347SS Weld CL, RT

347SS HAZ, RT

347SS 0.05" CL, RT

NASGRO Curve Fit Used

Figure 24. Fatigue crack growth rate data and curve-fit used for the NASGRO damage tolerance analyses. 
28 

Crack Size at Failure 

K > KIe 

ac = 2.54” 

Nc = 44,721 cycles 

Drawn to Scale 

ai = 0.5 inches 

Figure 25. Graphical representation of the damage tolerance analysis for a 1⁄2 inch crack to failure. 

0

0.5

1

1.5

2

2.5

3

0 10,000 20,000 30,000 40,000 50,000

Remaining Life (Pressure Cycles)

Crack Size (inches)

Figure 26. Crack size versus remaining life generated from the damage tolerance analysis of a 1⁄2 inch crack to 

failure. 
29 

0

10,000

20,000

30,000

40,000

50,000

30 40 50 60 70 80 90

Fracture Toughness, KIC (ksi in1/2)

Cycles to Failure

0

0.5

1

1.5

2

2.5

3

Critical Crack Depth, ac (inches)

Fracture Toughness

Critical crack depth

Figure 27. Cycles to failure and critical crack depth versus fracture toughness generated from the damage tolerance 

analysis of a 1⁄2 inch crack to failure. 
30 

Table 1. Typical tensile properties of annealed 347 stainless steels [5] 

Product 

Form 

Temp 

°F 

UTS, 

ksi 

YS, 

ksi 

Elong, 

% 

Reduct. 

Area, % 

E, 

Msi 

Charpy-V, 

ft-lb 

75 94 37 52 ---- 120 347 sheet, plate -320 198 61 47 ---- 66 

75 97.4 49.3 57 76 28.3 347 bar -320 214 62.2 43 60 30.3 

Table 2. Qualifying Mechanical Properties of 347 Weld Metal at -320°F

WELD METAL TEST RESULTS - rec'd via telecon from W. Vetter on 31 March 1959

Project 00218

Elongation Reduction

UTS YS (in 2 in.) Area

psi psi % % 

Weld Metal Sample 3 Ambient 81,090 48,450 42.9 40.3

Sample 4 Ambient 81,040 47,220 32.0 31.1

Sample 1 -320°F 173,470 48,890 36.6 29.9

Sample 2 -320°F 132,620 50,000 19.9 17.1

Table 3. Qualifying Mechanical Properties of 347 Base Material and Weld Metal at -320°F 

MATERIAL REORTS FOR 8000 GALLON LOX SPHERE

test results for low temperature tensile and Charpy Impact

Elongation Reduction Charpy

UTS YS (in 2 in.) Area Impact

Component Heat No. Slab No. psi psi % % ft-lbs

Weld metal 173,700 --- 25.0 25.3 18.5, 18.5, 16.75

169,500 --- 32.0 30.1

(retest of above) 173,476 48,894 36.6 29.9 not required

Plate: 336438 1 166,200 75,900 26.0 55.5 57, 56, 57

4-5/8" segments 336438 2 156,900 68,800 25.0 58.7 48, 49, 50

336439 1 170,000 72,700 27.0 56.4 49, 46, 48

336439 2 196,300 73,400 28.0 53.2 50, 49, 54

336440 1 190,800 70,600 28.0 53.2 51, 49, 56

336440 2 210,300 79,400 32.0 50.5 45, 44, 43

336441 1 154,500 72,700 30.0 58.2 41, 45, 45

336441 2 204,600 87,200 33.0 51.4 36, 35, 35

Flanges 40242 194,187 44,914 55.4 59.1 77, 75, 73

(forgings for 78501 225,500 --- 32.3 39.7 57, 55, 48

manhole & nozzles) 78505 210,200 --- 33.5 36.9 58, 40, 49

(Retest) 78501 205,641 48,225 45.7 51.0 not required

(Retest) 78505 213,845 40,427 44.0 51.5 not required

78579 193,840 37,681 36.0 26.3 45, 36, 49

4007 207,911 34,956 42.9 48.5 39, 41, 44

205,641 43,676 45.7 42.8 39, 41, 44

(Retest) 4007 207,461 46,769 43.6 47.9 not required
31 

Table 4. Chemical Analysis of the 347 Stainless Steel LOX Run Tank Sphere Segments 

Chemistry Analysis for Sphere Segments of LOX Vessel (Inner Sphere), wt %

Manufacturer ASTM Spec. Heat No. C Mn P S Si Cr Ni Cb Ta Cb+Ta

Allegheny 347 336438 0.044 1.69 0.024 0.011 0.55 17.65 9.46 0.66 0.058 0.71

" " 336439 0.049 1.61 0.025 0.01 0.64 17.69 9.16 0.63 0.049 0.67

" " 336440 0.05 0.92 0.026 0.013 0.58 17.92 9.66 0.67 0.047 0.717

" " 336441 0.044 1.07 0.029 0.013 0.57 17.92 9.66 0.81 0.045 0.855

Table 5. Weld Procedure Certification for Repair Welds of the LOX Run Tank 

Weld Procedure Certification No. G-4274 # 1: 1-17-65 

Base Metal: SA-240-T347 

Standards: Military Specification MIL E-22200 2A 

Electrode: E308 ELC-16 

250°F max interpass temp. Temperature to be checked by temple sticks 1⁄2” from area where weld 

bead is deposited. All weld slag will be completely removed before subsequent layers are deposited. 

All defects will be ground out to sound metal. Inspection: after gouging and grinding to the root of 

the outside weld, the groove preparation 1⁄2" on each side of the groove will be dye penetrant inspected. 

Dye penetrant inspection will be required at approx. 1 1⁄2” of weld, 3” of weld and final welded joint. 

This procedure is applicable to the repair welding of the shell penetrations for approx. 4 1⁄2” thick LOX 

storage tank. 

 1/8” 145 26 

 5/32” 165 26 

Table 6. Summary of Crack Growth Analyses 

Analysis Initial Crack 

Size (inches) 

Fracture Toughness, 

KIC (ksi in1/2) 

Repair Residual 

Stress (ksi) 

Cycles to Failure 

Dye Penetrant Crack 0.1 83 35 131,923 

Repair Crack 0.5 83 35 44,721 

Flow stress repair resid. 0.5 83 62.5 20,647 

0.5 40 35 0 

0.5 50 35 9,764 

0.5 60 35 23,748 

Low Fracture toughness 

welds 

0.5 70 35 35,396 
REPORT DOCUMENTATION PAGE Form Approved

OMB No. 0704-0188

2. REPORT TYPE 

Technical Memorandum

 4. TITLE AND SUBTITLE

Damage Tolerance Analysis of a Pressurized Liquid Oxygen Tank

5a. CONTRACT NUMBER

 6. AUTHOR(S)

Forth, Scott C.; Harvin, Stephen F.; Gregory, Peyton B.; Thompson,

Jon E.; and Hoffman, Eric K.

 7. PERFORMING ORGANIZATION NAME(S) AND ADDRESS(ES)

NASA Langley Research Center

Hampton, VA 23681-2199

 9. SPONSORING/MONITORING AGENCY NAME(S) AND ADDRESS(ES)

National Aeronautics and Space Administration

Washington, DC 20546-0001

 8. PERFORMING ORGANIZATION

 REPORT NUMBER

L-19212

10. SPONSOR/MONITOR'S ACRONYM(S)

NASA

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An electronic version can be found at http://ntrs.nasa.gov

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Unclassified - Unlimited

Subject Category 39

Availability: NASA CASI (301) 621-0390

19a. NAME OF RESPONSIBLE PERSON

STI Help Desk (email: help@sti.nasa.gov)

14. ABSTRACT

A damage tolerance assessment was conducted of an 8,000 gallon pressurized Liquid Oxygen (LOX) tank. The LOX tank is

constructed of a stainless steel pressure vessel enclosed by a thermal-insulating vacuum jacket. The vessel is pressurized to

2,250 psi with gaseous nitrogen resulting in both thermal and pressure stresses on the tank wall. Finite element analyses were

performed on the tank to characterize the stresses from operation. Engineering material data was found from both the

construction of the tank and the technical literature. An initial damage state was assumed based on records of a nondestructive

inspection performed on the tank. The damage tolerance analyses were conducted using the NASGRO computer code. This

paper contains the assumptions, and justifications, made for the input parameters to the damage tolerance analyses and the

results of the damage tolerance analyses with a discussion on the operational safety of the LOX tank.

15. SUBJECT TERMS

Analysis; Damage; Liquid; Oxygen; Pressurized; Tank; Tolerance

18. NUMBER

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 PAGES

36

19b. TELEPHONE NUMBER (Include area code)

(301) 621-0390

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5f. WORK UNIT NUMBER

732759.07.11

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NASA/TM-2006-214274

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01- 02 - 2006