METALLURGICAL EVALUATION AND I

FRACTURE ANALYSIS OF A |

PNEUMATICALLY-BURST SEAMLESS f

STEEL COMPRESSED GAS ' CYLINDER li

Bruce W. Christ and John H Smith ;! ti;-

Mechanical Properties Section

Metallurgy Division

Institute for Materials Research

National Bureau of Standards

Washington, D C. 20234

Failure Analysis Report

April, 1977

Prepared for Office of Hazardous Materials Operations

Department of Transportation

Washington, D.C. 20590

Contract No. DOT AS-40034

U.S. DEPARTMENT OF COMMERCE, Juanita M. Krops, Secretary

Dr. Betsy Ancker-Johnson, Assistant Secretary for Science and Technology

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Acting Director
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NOTICE

The contents of this report reflect the views of the authors

who are responsible for the facts and the accuracy of the data

presented herein. The contents do not necessarily reflect the

official views or policy of the Department of Transportation.

This report does not constitute a standard, specification or

regulation.

This document is disseminated under the sponsorship of the

Department of Transportation in the interest of information

exchange. The United States Government assumes no liability

for its contents or use thereof.
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TABLE OF CONTENTS

2 .

3.

4.

5.

GENERAL INFORMATION

1.1 Reference

1.2 Background

1.3 Identification and Pescription

PURPOSE

COMPARISON OF TEST RESULTS WITH REQUIREMENTS OF

SPECIFICATION DOT 3AA

3.1 Chemical Analysis

3.2 Magnetic Particle Inspection i.

3.3 Tensile Properties

3.4 Wall Thickness

FRACTURE DESCRIPTION

4.1 Origin of Fracture *

4.2 Probable Fracture Path

I'

j, 4.3 Description of Fracture Mode .

RESULTS OF METALLOGRAPHIC EXAMINATION

J -I 5.1 Observations of Fracture Profiles .

5.2 Inclusions

5.2.1 Kind and Orientation

5.2.2 Inclusion Gontent *

5.3 Microhardness Measurements Near Gouge

5.4 Search For Evidence of Heating

5.5 Evidence of Decarburization

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TABLE OF CONTENTS (continued)

6. RESULTS OF TESTS NOT REQUIRED IN SPECIFICATION DOT 3AA

6.1 Quenching

6.2 Impact Tests

6.3 Notched Tensile Tests

6.4 Tensile Tests on Circumferential Specimens

7. STRESS ANALYSIS VIA FRACTURE MECHANICS

7.1 Background

7.2 Hoop Stress At Failure

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7.3 Estimation of Fracture Toughness

8. DISCUSSION

9. CONCLUSIONS

1 0

.

ACKNOWLEDGEMENTS

LIST OF TABLES

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Metallurgical Evaluation and Fracture Analysis of a Pneumatically-Burst Seamless Steel Compressed Gas Cylinder

1. GENERAL INFORMATION

1.1 Reference

U.S. Department of Transportation, Bureau of Hazardous Materials,

Office of Hazardous Materials Operations, Engineering Branch, 2100-2nd

Street, S.W., Washington, D.C. 20590. This report was prepared at the

request of Mr. Arthur J. Mallen, Engineering Branch Chief, in a letter

dated October 20, 1975.

1 . 2 Background

Sometime during July, 1975, a seamless steel gas cylinder in oxygen

service exploded while it was being refilled. The event occurred near

San Francisco, California. It was reported that the pneumatic burst

occurred at a pressure of about 2500 psi. It was also reported that the

bursting cylinder grazed the ankle of the operator on the filling platform,

but that no substantial injury or other damage occurred.

The burst gas cylinder was crated and shipped to the National Bureau

of Standards for examination, at the request of Mr. Arthur J. Mallen and

with the concurrence of the operator of the gas cylinder filling facility.

The crated cylinder arrived at NBS on or about September 12, 1975.

Mr. Mallen also supplied a copy of the test report prepared by the

disinterested inspector licensed by the Department of Transportation to

carry out inspections and testing during gas cylinder manufacture.

Four preliminary reports on this cylinder were sent to the Office of

Hazardous Materials Operations!.

1. 3 Identification and Description

The seamless steel gas cylinder is identified as follows:

Specification

Test Date

Date of Test Report

Serial Number

Identifying Symbol

Type

Model

Minimum Sidewall

Heat Code

Type of Steel

DOT 3AA 2265

12-74

January 8, 1975

T-0967

PST

H

8BC250

0. 207-inches

V18

Water Capacity

Size

"Intermediate Manganese," quenched

and tempered

95.3

pounds, minimum

8 9/16-inches inside diameter

51-inches length

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Specification DOT 3AA can be found in section 178.37 of Title 49

of the Code of Federal Regulations^

.

\ A picture of the as-received cylinder appears in Figure la, where

it is shown with an intact cylinder of the same size for comparison.

The two halves of the neck appear at the upper left center and lower

right center. An overall view including the bottom appears in Figure lb.

All the metal of the cylinder appeared to have remained together, despite

the bursting forces. However, the neck ring and valve were not found in

the shipment.

Identifying markings found on the cylinder appear in Figure 2. The

Department of Transportation stamping at the neck end is in Figure 2a.

The test date and other stamping appear in Figure 2b. The heat code,

V18, was stamped on the bottom of the cylinder. Figure 3.

2. PURPOSE

The purpose of this report, as indicated in the request letter, is

fourfold

:

• to evaluate compliance of the cylinder with those requirements

of Specification DOT 3AA which relate to chemistry, magnetic

particle inspection, tensile properties and wall thickness.

• to carry out failure analysis.

• to carry out certain tests not appearing in Specification DOT 3AA.

• to carry out a stress analysis via fracture mechanics.

3. COMPARISON OF TEST RESULTS WITH REQUIREMENTS OF SPECIFICATION DOT 3AA

3 . 1 Chemical Analysis

Chemical analysis was carried out on a specimen cut from the bottom

of cylinder T-0967. Results appear in Table I, along with a tabulation

of the specification requirements. It is clear that these results comply

with the chemistry requirements in section 178.37-5-(a) of specification

DOT 3AA for "intermediate manganese" steel.

There is no requirement in specification DOT 3AA on the concentration

of the residual elements cited in Table II. However, the concentrations

of residual elements fall within the limits defined in the American Iron

and Steel Institute Steel Products Manual^, a widely recognized

specification for the chemistry of steels.

Titanium is not included in either the DOT specification or the AISI

specification. The rather high concentration of titanium, 0.03 weight

percent, is notable, however. For this much titanium to be present, it

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making process to deoxidize the molten steel. Traditionally, silicon

and aluminum serve as deoxidizing agents. The use of titanium as a deoxidizer of commercial heats of' steel has been increasing in recent

years.

Additions of deoxidants lead to the formation of oxide and nitride

inclusions. Furthermore, sulfur combines with manganese to produce

manganese sulfide inclusions. Analysis of the inclusions in this steel

is reported in Section 5.2, where it is shown that manganese sulfide is

the predominant type of inclusion, as is the case for most steels.

3. 2 Magnetic Particle Inspection

Magnetic particle inspection was applied to the exterior and interior

surfaces of the as-received cylinder. No cracks were detected. Since

this inspection was applied to the burst cylinder, it is concluded that

the as-manufactured cylinder must have conformed to the requirements of

section 178. 37-ll-a-(6) of specification DOT 3AA2.

3 . 3 Tensile Properties

Four longitudinal tensile specimens were prepared in accord with

section 178 . 37-16- (b) of specification DOT 3AA2. Because the as- received cylinder was distorted as shown in Figure 1, material was flattened at room temperature in a press before machining the tensile

specimens. This material was taken as close as possible to the origin

of fracture. Test results appear in Table III. Cold work introduced

during catastrophic failure or flattening would have the tendency to

raise the strength of the material. Yield strength is usually more

sensitive to cold work than is tensile strength.

In section 178 . 37-17- (a) of specification DOT 3AA, it is required

that elongation in the tensile test be "...at least 20 percent for 2"

gauge length...". Results in Table III show that this requirement was met

by all four tensile specimens, and that the average elongation exceeded

the minimum required by 8.2 percent. Furthermore, ductility data which

appear in Table III compare favorably with results which appear in the

test report.

A notable feature of the tensile strengths in Table III is that they are about 7% lower than the tensile strengths appearing in the test report.

The reasons for this difference are not known.

3 . 4 Wall Thickness

Wall thickness measurements were made along the fracture surfaces of

the burst gas cylinder. Results are shown in Table IV, where the sidewall

thickness ranges from 0.220 to 0. 260- inches . Using the minimum measured

wall thickness, 0. 220- inches , the wall stress at test pressure was

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calculated from the formula in section 178. 37-10-(c) of specification

DOT 3AA,

S = 3775 (1.3(9.003)^ + 0.4 (8.563)^)/

((9.003)^ - (8.563)^) = 65,800 psi.

This wall stress slightly exceeds the allowable wall stress at test

pressure of 64,500 psi, which is determined by taking 67% of the minimum

tensile strength of the steel, 96,800 psi, as defined in section

178. 37-10-(b) of specification DOT 3AA. Hence, this requirement was

not met. However, this wall stress does not exceed the value of

66,900 psi, which results from taking 67% of the average tensile strength

reported in Table III. )'

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i A. FRACTURE DESCRIPTION

I 4.1 Origin of Fracture

I

! The pneumatically-burst cylinder was split by three longitudinal

I tears running from top to bottom, as is evident in Figure 1. Shear

: fracture was the principal fracture mode. The six longitudinal fracture

^ surfaces all of which were relatively unrusted in the as-received

I condition were scanned visually for evidence of fracture origins.

' Only one region was found which was suspected of being a fracture

; origin. The two sides of thi^ region are labeled a and a' in Figure lb

the region is about 10- inches below the neck.

The side of the origin labeled a' in Figure lb is shown at two

different angles in Figures 4a and 4b. The origin consists of:

i 1. A rusty and discolored region about 3-inches long, with elliptical

ends. This region penetrates about 91% of the wall thickness.

,2. A narrow, unrusted highly-ref lective band along the inside wall,

! spanning about 9% of the wall thickness.

3. A small step of perhaps 3/16-inch near the neck end.

, 4. A gouge beneath the step.

^ 5. Some lips of raised metal on the external surface.

6. Some rusted regions where paint was missing.

I! Figure 5 shows the side of the fracture origin labeled a in Figure lb.

The step and highly-ref lective band along the inside wall are evident.

Dimensions of the fracture origin are shown in the schematic diagram

* included in Figure 5.

I The fracture surface within the origin was fibrous, except on the

step. In the as-received condition, the step showed up as a smooth and

brightly-reflecting surface. The smoothness of the step contrasts with

the fibrous nature of the adjacent fracture surface in Figure 6, which

^ was obtained with the scanning electron microscope.

li Mating halves of the origin were cut out and placed together as shown

in Figure 7 to give an approximate picture of the configuration before burst.

Curvature of the pieces which probably developed during burst prevented precise matching. The lips of raised metal on the side which

I appeared in Figure 4b are evident in Figure 7. Here, they

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i have the appearance of having been flattened somewhat. The other side of

the fracture origin shows regions where it had apparently been gouged.

! To examine the fracture origin more closely, metallographic sections

I were cut from one side of the origin only. The other side was preserved

i intact. Specimens II, III and IV, as labeled in Figure 7, were cut to

j show the fracture profile at the origin. Specimen I near the origin was

, cut to show the fracture profile in the region of the fast-running crack.

A specimen including the step was cut for examination of the fracture

j origin in the scanning electron microscope.

I Figure 8 shows the fracture profile at Specimen IV. A gouge near the

fracture surface is evident in Figure 8a. The gouge is about 0.02-inches

deep. Flowed metal beneath the gouge can be seen in Figures 8b and 8c.

The most severely flowed metal seems to penetrate about 0. 0006- inches

below the bottom of the gouge. Away from the flowed metal, the microstructure has the usual appearance of a quenched and tempered steel.

j These observations a gouge, flowed metal beneath the gouge, and a typical quenched and tempered microstructure were found in Specimens

II and III, also, indicating a reasonably uniform disturbance throughout

I the region shown in Figure 7. A crack in the bottom of the gouge was

I detected in the fracture profile of Specimen II, as shown in Figure 9.

I Figure 10 shows rusted regions where paint is missing near the fracture

origin. Some of these rusted regions are probably due to abrasion of

paint during normal handling of the cylinder. Other rusted regions bear

a striking resemblence to the temporary identification markings usually

I placed on the cylinder at the time of manufacture. For example, it is

possible to discern "67 ) 114" in Figure 10. "67" is identical to the

i last two digits in the serial number T-0967 and "114" is identical

‘ to the tare weight noted for this cylinder on the Record of Hydrostatic

j Tests. The "114" can be seen in Figure 7 if it is turned upside down.

These markings are known to develop when paint applied at the time

of manufacture over temporary markings does not adhere at the location

of the markings. It appears purely coincidental that these markings show

ij up near the fracture origin.

I 4.2 Probable Fracture Path

I

j It is helpful to imagine the as-received cylinder folded back into

I its original shape in order to discuss the probable fracture path. It

I is likely that two cracks moved out from the origin marked at a and a'

J in Figure lb, one traveling toward the top and the other toward the

j bottom. Evidently, the crack which moved toward the top jumped the

I neck. A clean break is indicated by the relatively undamaged threads

i in the neck. Figure 11. After jumping the neck, this crack appeared

j to have traveled longitudinally toward the bottom and arrested in the

I 0.6- inch thick bottom. ,

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' reaching the bottom, Figure lb. One crack traveled radially in the

j bottom and arrested. A second crack traveled a short way circumferentially

around the bottom and arrested. And the third crack traveled circumferentially a bit, then turned upward and followed a longitudinal path to an arrest point in the relatively thick neck.

4 . 3 Description of Fracture Mode

Ductile fracture seemed to be the only fracture mode for the

longitudinal tears. This observation applies to the fracture origin,

as well as to the regions of the fast-running crack. Ductile fracture

was evidenced by the fibrous appearance of the fracture surface

Figures 5, 6 and 11 as well as by wall thinning near 'the fracture

surface. Wall thinning is evident in Figure 12, where the throughthickness fracture profile of Specimen I is shown. Measurements made

on Figure 12 show that the wall thickness at the fracture surface is

about 7% less than the wall thickness about 0.3-inches away from the

fracture surface. Ductile fracture was also evidenced along the

longitudinal tears by shear lips at about 45° to the longitudinal

fracture plane. Figure 13.

There is no obvious explanation for the smooth surface of the step

between the pair of unlinked, part-through cracks. The crack arrest

point which appears under the step in Figure 4b may have developed when

the step formed.

5. RESULTS OF METALLOGRAPHIC EXAMINATION

5. 1 Observations of Fracture Profiles

Fracture profiles were examined in the region of the fracture origin

using Specimens II, III and IV, as indicated in Figure 7. The objective

was to look for differences between the features of the fracture profile

along the rusty and discolored region which spanned 91% of the wall

thickness, and along the unrusted, highly-reflective band spanning 9% of

the wall thickness, as indicated in Figures 4 and 5. Typical results

appear in Figure 14.

Figure 14a shows the full thickness fracture profile of Specimen IV.

Figures 14b and 14c are close-ups of the regions indicated by the upper

and lower arrows, respectively, in Figure 14a. In general, the fracture

profile near the upper arrow showed rounded recesses and rounded dimples,

which were coated with corrosion product. In the region indicated by the

lower arrow, the fracture profile showed sharp recesses and pointed

dimples. Similar observations were made on the full thickness fracture

profiles of Specimens II and III. These results support the view that

a part-through crack existed at the fracture origin. Evidently, the

part-through crack had been exposed to a corrosive medium, but it is not

possible to establish the time of exposure accurately.

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I in Figure 14b. The orientation of these voids suggests that they formed

I where inclusions of the type shown on face C of Figure 15 (See Section

I 5.2.1) separated from the steel matrix. The corrosion product in the voids,

i and also the rounded contours of the voids, suggest that they were connected

I to the fracture surface. The connection is not evident in Figure 14b,

I probably because of the way the metallographic section was cut. Perhaps

I these and other voids formed in response to local stresses which caused

I some shape distortion when the part-through crack formed.

I 5.2 Inclusions

^ 5.2.1 Kind and Orientation

From metallographic observations in three directions, Figure 15,

it was concluded that the kinds of inclusions present included manganese

sulfide, silicates and titanium nitride. Intact inclusions elongated in

the longitudinal direction are probably manganese sulfide inclusions.

Fragmented inclusions strung out in the longitudinal direction are probably alumina inclusions. And those inclusions which appear as squares or rectangles are probably titanium nitride inclusions.

The orientation of inclusions in the gas cylinder wall is developed

by the fabricating process. For example, the manganese sulfide inclusions

are highly ductile at forging and drawing temperatures. Consequently,

these inclusions are elongated into continuous platelets, with the

longest axis in the longitudinal direction as is evident in Figures 15b

and 15c. The second-longest axis of the platelet is in the circumferential

direction, as is evident in Figures 15b and 15d. On the other hand,

alumina and titanium dioxide inclusions are not ductile at forging and

drawing temperatures, and consequently these inclusions are fragmented

and strung out with the principal axis in the longitudinal direction,

Figures 15b and 15c. Moreover, these more brittle inclusions are not

readily flattened into platelets.

The orientation of all these kinds of inclusions has a strong

influence on the direction in which cracks can run most readily. It

is common for cracks to run along the interface between an inclusion

and the steel matrix because the fracture energy along this interface

is lower than the fracture energy of the steel matrix itself. Consequently,

in a structure with predominantly longitudinal inclusions, longitudinal

crack propagation will require less energy than circumferential crack

propagation, and failure via longitudinal crack propagation will be

assisted. Of course, the primary reason for longitudinal crack propagation

in a cylindrical pressure vessel is the fact that the hoop stress is twice

as large as the axial stress.

Circumferential crack propagation will be assisted by inclusions which

are slightly extended in the circumferential direction and also by

localized stresses associated with shape distortion. Measurements of the

orientation dependence of the energy for crack propagation appear in

Section 6.2

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5.2.2 Inclusion Content

A quantitative estimate of the inclusion content was carried out as specified in ASTM Designation E 45-74 "Standard Recommended Practice

For Determining the Inclusion Content of Steel^" Method A of this

recommended practice was used on face A, Figure 15b, on a single specimen

which was about 1/2 x 1- inches. Counts were made by two observers.

Manganese sulfide and alumina inclusions were the principal types found.

For this steel, titanium nitride inclusions were counted under the alumina

classification. Manganese sulfide and alumina inclusions are designated

as Type A and Type B inclusions, respectively, on the Jernkontoret chart

cited in ASTM E 45-74. Inclusions are classified as thin or heavy. As

shown in the following tabulation, only thin sulfide inclusions were

found, whereas both thin and heavy alumina inclusions were found.

INCLUSION CONTENT RATED BY ASTM E 45-74 - METHOD A

(Based on worst field in 0.25 sq. inch specimen)

Manganese Sulfide Alumina

(Type A) (Type B)

Thin Heavy Thin Heavy

First Observer 3.60 — 1.25 1.25

Second Observer 3.67 — 1.67 —

Average 3.63 — 1.46 1.25

The ASTM scale of inclusion rating numbers ranges from 1 to 5, with

1 representing the lowest inclusion content. A rating number of 3 to 4

is not unusual for a steel with the concentrations of sulfur, phosphorus

and oxygen shown in Tables I and II. Of particular interest in Table I is the observation that the sulfur concentration is the maximum permitted

by specification DOT 3AA. The inclusion contents for classifications other

than manganese sulfide and alumina Inclusions, namely, silicate (Type C)

and globular (Type D) , were so low that they did not register on the ASTM

scale

.

5 . 3 Microhardness Measurements Near Gouge

In Figure 8, evidence of a 0.02- inch deep gouge and evidence of

flowed metal beneath the gouge was shown for the fracture profile of

Specimen IV. Two Knoop microhardness traverses were made on this

metallographic specimen, along the lines labeled A and B in Figure 8a.

The purpose of making these hardness measurements was to search for

evidence of strain hardened metal beneath the gouge. To facilitate

accurate measurement, hardness was measured using the Knoop microhardness

technique. The range of Knoop hardness numbers measured, 257 to 313,

corresponds roughly to Rockwell hardness numbers on the C-scale of 21

to 30.

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Figure 16. The mid-thickness of the wall occurs at a deptli of about

I 0.12- inches. It can be seen in Figure 16 that microliardness at depths

I between 0.01 - 0.03 below the surface as indicated in Figure 8a , is

i greater for traverse A than for traverse B. Below 0.03 inches, micro-

’ hardness is about the same along both traverses. These results are interpreted as evidence of strain hardened metal beneath the gouge.

They confirm that the metal had flowed when it was gouged.

j 5.4 Search For Evidence of Heating

I In general, etched specimens of this gas cylinder steel developed

; a rather dark appearance, as shown in Figures 8 and 12. This type of

i etching is characteristic of a quenched and tempered steel, which is

, the heat treatment reportedly given to this cylinder at the time of

manufacture. Had the cylinder been heated in the region from which

these specimens were cut the region shown in Figure 7 some evidence would probably appear in the etched microstructure. For

example, had an arc strike occurred or had weld metal been deposited

I in the fracture origin, a detectable change in the microstructure would

j have developed. Since such a change was not evident, it is inferred

that heating of the cylinder wall to temperatures associated with arc strikes or welding did not occur.

; 5.5 Evidence of Decarburization

Incidental to the observation of the fracture profile is the

observation of a decarburized layer along the inside cylinder wall,

i The layer shows up as the white region along the inside wall in

' Figures 8 and 12. Specimens I, II, III and IV each showed evidence

I of decarburization the layer was about 0.002 to 0.004 inches deep.

This decarburized layer does not seem to be related to the fracture of

the gas cylinder.

[ 6. RESULTS OF TESTS NOT REQUIRED IN SPECIFICATION DOT 3AA

6 . 1 Quenching

Because of the widely-recognized susceptibility of intermediate

manganese steel to quench cracking, some laboratory quenching tests were carried out in an attempt to evaluate the quench cracking behavior

of the steel used in gas cylinder T-0967. Test specimen dimensions

were about 1 1/2 x 3/4 x 1/4 - inches. Four specimens were austenitized

in a salt bath at 1600°F for about 40 minutes. One specimen was quenched

in oil at about 140°F, one in oil at room temperature, one in boiling

water and one in ice water. The as-quenched specimens were examined

visually for cracks none were found. To check further, the specimen

receiving the most severe quench was sectioned for metallographic

examination. No cracks were found in this section. It was concluded

that quench cracking was not a likely source of defects, at least for

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these small size specimens. Furthermore, results of the magnetic particle

inspection reported in Section 3.2 indicated that quench cracks had not

formed at the time of manufacture, when the whole gas cylinder was

quenched. '

I 6.2 Impact Tests

I

I Impact tests were carried out to evaluate the toughness of the steel

i’ in gas cylinder T-0967. Toughness is basically a measure of the energy

I absorbed during fracture. High values of absorbed energy generally

I indicate a low susceptibility to brittle fracture.

\

j Charpy test specimens were prepared according to ASTM Specification

I E 23-72.5 The cylinder wall was not thick enough to allow machining fullI size Charpy test specimens so half-size specimens were prepared. All

j specimens were machined from material which was flattened at room

j temperature in a press. Flattening was necessary because of the

' distorted condition of the as-received cylinder. Flattening also helped

overcome the natural curvature of the cylinder. Specimen dimensions were

'* 0.197 X 0.395 X 2.165 - inches. Notch depth was 0.089 - inches

] 0.010 - inches deeper than the standard ASTM notch depth. This slightly

j; deeper notch resulted from a machining error.

I

i

Charpy test results appear in Figures 17 and 18. Absorbed energy

for longitudinal and circumferential crack propagation appears in

Figures 17a and 17b, respectively. Lateral expansion for longitudinal

and circumferential specimen orientations is plotted in Figures 18a and

18b, respectively. The notch orientation and crack propagation direction

are shown pictorially on each figure.

Results in Figures 17 and 18 show a transition from higher to lower

values of toughness, starting at about -40°F. At and above 0°F, these

results show that there is less resistance to longitudinal crack

propagation than to circumferential crack propagation. The absorbed

energy data show an increase in resistance by a factor of 3.6, and the

lateral expansion data show an increase in resistance by a factor of 3.3.

Hence, the manner in which gas cylinder T-0967 failed three fulllength longitudinal tears is consistent with these results. As

noted in Section 5.2.1, the primary reason for longitudinal crack propagation in a cylindrical pressure vessel is the fact that the hoop stress is

twice as large as the axial stress.

In Table V, the energy absorbed by Charpy test specimens is

compared with the energy requirements of the proposed ISO specification

for gas cylinders. Comparison is made with the ISO specification simply

because it is the only known seamless steel gas cylinder specification

with Charpy test requirements. These requirements are exceeded substantially

in the case of longitudinal specimen orientation - circumferential crack

propagation. However, the energy requirement of the specification is not

met in the case of circumferential specimen orientation - longitudinal crack

propagation. These comparisons emphasize the importance of specifying

notch orientation in Charpy testing.

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j 6 . 3 Notched Tensile Tests

Tensile tests were carried out on notched specimens to evaluate

the influence of a notch on ductility. Three test specimens were

machined from material which was flattened at room temperature in a press. Flattening was necessary to overcome the distorted condition of

the as-received cylinder.

Test specimens were of the same dimensions as those for the

j tensile specimens described in Section 3.3. However, each specimen

: had notches machined into the sides at the center of the gage length,

i as indicated in the footnote to Table V. A notch in a tensile specimen

I of this size constrains the specimen from deforming in the width

'direction. Such constraint results in multiaxial stresses, as is the

. case in a pressurized gas cylinder. This stress state is exaggerated in

i the presence of a part-through crack.

Test results appear in Table VI. Comparing ductility in Table VI

I with ductility in Table III shows a significant decrease for the notched

j specimens average reduction in area decreased from 48.8 to 9.7%.

j Another significant influence of the notch is evident when comparing

I strengths in Tables VI and III. The average notched tensile strength is

i 113,700 psi, as compared to the average unnotched tensile strength of

3 99,800 psi. These results are consistent with known influences of notches

j on uniaxial tensile test results, namely, a significant loss of ductility

I and an apparent increase in tensile strength. Since all specimens were

) flattened, it is unlikely that the process of flattening influenced the

li trends in these comparisons.

t_6.4 Tensile Tests on Circumferential Specimens

bl Tensile tests were carried out on specimens so oriented that their

• long axis was in the circumferential direction, rather than in the longitudinal direction, as is usual. The objective of these tests was to determine

whether or not there was any influence of test specimen orientation on

I tensile properties. All specimens were machined from material which was

I flattened at room temperature in a press. Flattening was necessary to

1 overcome the natural curvature of the cylinder as well as to overcome

I’ the distorted condition of the as-received cylinder. Dimensions of

these test specimens were the same as those of the longitudinal specimens

described in Section 3.3.

The average tensile and yield strengths in Table VII compare favorably

with the average tensile and yield strengths in Table III. However,

reduction in area and elongation show decreases of about 30 and 40%,

respectively. Hence, there seems to be no orientation dependence for

strength, but a substantial dependence for ductility. This reduction in

ductility is attributable in part to the inclusions elongated in the

longitudinal direction, Figure 15.

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The circumferential ductility shown in Table VII, 20% elongation,

jcompares favorably with the circumferential ductility required in the

flattening test, section 178.37-17(a) of specification DOT 3AA, which

lis 14-18%, as calculated from anvil separation, cylinder diameter and

iwall thickness.

!

j 7 . STRESS ANALYSIS VIA FRACTURE MECHANICS

17.1 Background

i Fracture mechanics is a field of stress analysis which evaluates

i the influence of cracks on the stress which a structure can safely

'support. The development of this field during the last thirty years

resulted from the widespread recognition that most engineering structures

! may contain cracks which could lead to failure at applied stresses well

below the yield stress. The utility of fracture mechanics lies in its

I ability to predict critical crack sizes in terms of applied stress and

resistance to crack propagation. A structure with cracks below the

critical size would remain intact, whereas a structure with larger cracks

would fail

.

I Fracture mechanics measures the ability of a material to resist

crack propagation in terms of "fracture toughness," signified by "K^ "

for plane strain conditions, and "K " for plane stress, and with units

of psi - finches . Typical values o^ frac ture toughness for steels range

between 50,000 and 400,000 psi - ^/inches . Typical aluminum alloys

' exhibit fracture toughnesses in the range, 20,000 - 100,000 psi - finches.

Fracture toughness can be measured on precracked test specimens in an ordinary tensile testing machine. ASTM Standard Method E 399 - 72&

describes plane strain fracture toughness testing of metallic materials.

The application of fracture mechanics to pressure vessels facilitates

the prediction of stability, leaking, or catastrophic fracture in terms of

I fracture toughness, crack size and applied stress. The application of

fracture mechanics to thin-walled cylindrical shells made from ductile metals,

I e.g., pipe, gas cylinders, is presently in the early stages of development . 7-9

As regards gas cylinder T-0967, it was reported that the cylinder

burst during a filling operation at a pressure of about 2500 psi. (See

Section 1.2). This pressure is significantly less than the design burst

pressure, estimated as 5100 psi by multiplying the marked filling pressure

by the appropriate safety factor, 2.25. Results from this investigation

|i! show that the probable cause of this discrepancy between the design and actual

I burst pressures was a pair of closely-spaced, part-through cracks, as shown in Figures 4 and 5. The overall length is 3 inches, and the

i depth is 0. 232- inches , which is about 91% of the wall thickness. One

I crack is 2 1/2- inches long, and the other is 1/2- inch long. The step,

* which is about 3/16-inches high, appears in Figure 6. The non-fibrous

fracture surface of the step suggests that it did not form when the

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i fibrous fracture surfaces of the part-through cracks and the fast-running

cracks formed. Hence, it is assumed that the step formed during the

j catastrophic rupture of the gas cylinder.

j The cracks have a definite elliptical shape at the end away from the

I step. This shape is a widely-recognized characteristic of part-through

I cracks.lO Judging from the depth of the cracks, it is likely that the

i cracks were sharp at the bottom rather than blunt. Furthermore, metallographic

evidence from the part-through crack surfaces indicates that the fracture

j was ductile. The fracture in the region of the fast-running crack which

: developed at temperatures near or somewhat above room temperature, was also

j ductile. (See Section 4.3). The observed ductile mode of fracture is

! consistent with the impact data in Section 6.2, which show that this steel

I is on the upper shelf at temperatures above 0°F.

j 7.2 Hoop Stress At Failure

I The following fracture mechanics evaluation of gas cylinder T-0967

I is based on the empirical analysis of Maxey^ . This analysis was developed

I for pipe with yield strengths from 40,.000 to 70,000 psi, but it can be

' applied to other tubular products which are approximately in this yield

i strength range, including gas cylinders. As a starting point, Maxey®

j (and others)',^ point out that stress magnification will occur at the ends

I of a crack. Consequently, plastic flow and fracture can occur at nominal

applied wall stresses lower than the yield strength. In the case of a longitudinal crack in the sidewall of a gas cylinder, stress magnification

is due to redistribution of the hoop stress around the crack and to bending

1 moments caused by bulging near the crack.

The stress magnification factor due to bulging of tubular products

containing a through-wall, axial defect, M^, was developed by Folias"^

and is given by

2 4 = (1 + 1.255 - 0.0135

T RT ^2^2

) 1/2 (la)

For the case of a part-through axial crack, the stress magnification factor,

Mp, is estimated by Maxey^ asMp = (1 - d/M^t)/(l - d/t) (lb)

In the above equations.

d = depth of a part-through crack

2c = length of a through-wall crack

R = cylinder radius

t = wall thickness

2c' = equivalent crack length in the case of a part-through crack

with non-rectangular ends

Mrp' is calculated using c' in equation la.

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The stress magnification factor shows up in the equations used to

calculate the hoop stress at failure. Equations have been defined by

MaxeyS for two types of fractures, 'classified as "flow stress controlled"

(plastic behavior) and "toughness controlled" (elastic-plastic behavior).

Equations for both types of behavior appear below.

In the case of a through-wall, axial crack, the equation for

toughness-controlled behavior, is as follows:

Jin sec j(M^o^/op= K^-n/(8co^) = 12ETf (C^/A^) / (8c a^) (2a)

For a through-wall defect, the equation for flow stress controlled

behavior is

* a„ = o ,/Hr (2b)

In the case of a part-through axial crack, the hoop stress at failure

for toughness-controlled behavior, Op, is given by

Jin sec 2PPrc (M^a /aJ = K^n/(8c'o^) = 12Eir(C /A )/(8c'a^)

f VC f

whereas for flow stress controlled behavior, the equation is

* —

(3a)

Or = a/Mp. (3b)

In the above equations 0^, Op = hoop stress at failure for toughness

controlled behavior,

ap. Op = hoop stress at failure for flow stress controlled behavior

0

^ = the flow stress of the steel (assumed by Maxey and others

to be yield strength + 10,000 psi)

K = plane stress fracture toughness of material

= impact energy per unit area of a Charpy test specimen (half

size or full size)

E = Young’s modulus (30 x 10 psi)

2c or 2c ’= the length of the crack

Mp = stress magnification factor given by Equation la

Mp = stress magnification factor given by Equation lb

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j Stress analysis cannot be easily carried out for the case of the two

i closely-spaced , but separate cracks shown in Figures 4 and 5. Hence, to

estimate upper and lower bounds for' the hoop stress at failure, calculations

were made for the actual 2 1/2-inch long crack and for an assumed 3-inch long

''crack. Because of the depth of the part-through crack, 91% of the

wall thickness, determination of c’ rather than c for calculating was considered an unnecessary refinement. Results

appear in Table VIII, where the first column shows that the stress

magnification factor is larger for the part-through crack than for the

full through crack. The second column shows a factor based on the Charpy

j test results which is used to classify the fracture as toughness controlled

jor flow stress controlled. Maxey indicates that values of this factor less

I than 4, as appearing in the table, mean that the fracture is toughness

I controlled. Thus, Equations 2a and 3a were used to calculate the hoop

'stresses at failure, which appear in the third column. Values of the

I parameter, Ma/o^, which appear in the fourth column, also show that the

'fracture was slightly toughness controlled.

In the case of the part-through crack, failure means that the

- ligament beneath the part through crack ruptured, and a full- through

.crack formed. At this point, results in Table VIII suggest that the

1 cylinder simply leaked, because the ligament ruptured at a hoop stress of

i 12,200 to 20,200 psi, which is less than the stress required for the fullI through crack to run catastrophically, namely, 38,100 to 54,300 psi. Continued

filling of the leaking cylinder to the allowable filling pressure would have

.developed hoop stresses in this range. For example, the hoop stress

I estimated from the reported gas pressure at burst,

!,j Og = PR/t = 2500 X 4.28/0.256 = 41,800 psi, falls in this stress

|j range.

||7.3 Estimation of Fracture Toughness

The plane stress fracture toughness of the quenched and tempered

i steel in gas cylinder T-0967 can be estimated from Maxey's^ empirical

i correlation

:! K = /12 E C /A (4)

C VC

Ij where the terms have the meanings defined in Section 7.2. The estimated

i fracture toughness is orientation dependent, due to the orientation

I dependence of the Charpy values shown in Table V. Fracture toughness

1 values appear in Table IX. These values compare favorably with other

I plane stress values reported for quenched and tempered steels of

J similar yield strength. 12

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8. DISCUSSION

Results in Section 3 demonstrate that the material in gas cylinder

T-0967 conformed to the requirements of specification DOT 3AA, in matters

of chemistry, resistance to crack formation during quencliing, strength

and ductility. One notable feature in the chemistry is that the sulfur

concentration was at the maximum allowable, leading to the maximum content

of manganese sulfide inclusions, which are known to lower crack propagation

resistance. Another notable feature in the chemistry is the relatively

high titanium content, which suggests the possibility of deoxidation with

titanium rather than with the more traditional elements, aluminum and

silicon. Magnetic particle inspection did not reveal any cracks in the

as-received cylinder. Furthermore, it was found that quench cracks could

not be developed in small test specimens which were given a severe quench

in ice brine. Ductility exceeded specification requirements significantly.

Minimum tensile strength in Table III and minimum wall thickness in

Table IV are such that the calculated hoop stress at test pressure,

65,800 psi, exceeded by about 2 percent, the allowable hoop stress,

.67 x 96,800 = 64,500 psi, as defined in section 178.37-10(b) of

specification DOT 3AA. Hence, this strength requirement was not met.

This observation is consistent with the hydrostatic test report, which

indicates a rather high permanent expansion for gas cylinder T-0967, as compared with other cylinders in the same heat treatment lot. As a matter

of general interest, it is noted that the calculated hoop stress at test

pressure, does not exceed the value of . 67 x 99,800 = 66,900 psi, which

is obtained by taking 67 percent of the average tensile strength reported

in Table III.

The burst of gas cylinder T-0967 during filling seems traceable to

the presence of two unlinked, but closely-spaced part- through cracks

which were located about 10-inches below the neck. These cracks penetrated

about 91% of the wall thickness. They appeared to be related to a surface

disturbance which was evidenced by a gouge, lips of raised metal and strain

hardened metal close to the surface. The part-through cracks, the gouge,

and the lips of raised metal suggest that the cylinder received an impact

of mechanical origin. Furthermore, the rust appearance of the fracture

surfaces of the part-through cracks and the rust on the damaged exterior

surface near the cracks suggest that the cracks were present at the time

the cylinder was being filled. That the reported burst pressure was much

lower than the design burst pressure also suggests that the defects were

present at the time of filling.

From a fracture mechanics analysis of the cylinder containing a part-through crack, it was inferred that the ligament beneath the partthrough cracks ruptlired during filling at a hoop stress between 12,200

and 20,200 psi. It was also inferred that the cylinder leaked while

filling continued, until a critical hoop stress between 38,100 and

54,300 psi was reached, at which stress the full-through crack ran catastrophically and ultimately formed three full-length longitudinal

tears. The uncertainty in the hoop stresses at which rupture occurred
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arises because the step in the part-through crack makes precise analysis

difficult. Nevertheless, the hoop stress estimated from the reported gas

pressure at burst, 41,000 psl, falls in this stress range. Protection

against catastrophic crack propagation can be designed into a gas cylinder

by developing sufficient fracture toughness to resist extension of the

largest expected defects at operating pressures. The observed low strength

of gas cylinder T-0967 would not have affected its catastrophic failure

via longitudinal tears unfavorably. In fact, a trend toward lower

strength usually signifies increasing fracture toughness, and thus,

increasing resistance to fracture.

Ductile fracture seemed to be the only mode of fracture for the

longitudinal tears. This observation applies to the fracture origin,

as well as to the region of the fast-running crack. There is no obvious

explanation for the smooth surface of the step between the pair of

unlinked, part- through cracks.

Results of impact tests on Charpy specimens showed that longitudinal

crack propagation required significantly less energy than did circumferential

crack propagation. This result is attributable to the rather high concentration of sulfide inclusions elongated in the longitudinal direction. The

absorbed energy for circumferential crack propagation exceeded the requirement in a proposed ISO specification. However, no requirement was defined

in that specification for longitudinal crack propagation, which was the

principal crack propagation direction in the pneumatically-burst gas

cylinder. This behavior, and the observation of a lower energy for longitudinal crack propagation, demonstrate the importance of recognizing the

orientation-dependence of mechanical properties.

Notched tensile specimens showed a substantial loss in reduction in

area, as compared to smooth tensile specimens. This observation suggests

that notch sensitivity is an important characteristic of this quenched and

tempered carbon-manganese steel.

Tensile test results on circumferential specimens were similar to

results on longitudinal specimens in the matter of yield and tensile

strengths. Ductility of the circumferential specimens was significantly

lower, but still compared favorably with the circumferential ductility

required in the whole-cylinder flattening test. These observations

demonstrate once again the importance of recognizing the orientationdependence of mechanical properties.

9. CONCLUSIONS

1. The material in gas cylinder T-0967 conformed to the requirements

of specification DOT 3AA in matters of chemistry, resistance to

crack formation during quenching, and ductility.

2. The material in gas cylinder T-0967 did not conform to the requirements of specification DOT 3AA in matters of tensile strength and

wall thickness. Minimum tensile strength and minimum wall thickness

17
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were such that the calculated hoop stress at test pressure exceeded

the allowable hoop stress by about 2.0%. This strength consideration

would not cause an unfavorable effect on the fracture toughness of

the material.

3. The burst of gas cylinder T-0967 during filling seems traceable

to the presence of two unlinked, but closely-spaced part-through

cracks which were located about 10-inches below the neck. The

cracks probably resulted from an impact of mechanical origin and

probably were present at the time the cylinder was being filled.

4. Results of impact tests on Charpy specimens showed that longitudinal

crack propagation required significantly less energy than did

circumferential crack propagation.

5. Results of tensile tests showed that longitudinal test specimens

exhibited significantly greater ductility than did circumferential

test specimens.

6. Results of tests on notched tensile specimens indicate that notch

sensitivity is an important characteristic of quenched and tempered

carbon-manganese steel.

18
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10. ACKNOWLEDGEMENTS

Mr. C. H. Brady obtained documentation photographs and conducted

the metallographic work. Magnetic particle testing was done by Mr

Mr. T. R. Shives. Mr. L. Smith, Mr. T. P. Royston and Mr. W. A. Willard

assisted in mechanical testing and in the preparation of the manuscript.

Mr. G. Hicho provided assistance with the inclusion counting, and

Mr. Hicho and Dr. C. G. Interrante assisted with the fracture mechanics

analysis. Mr. A. Mallen and Mr. P. Seay contributed several helpful

discussion during the course of this work.

19

REFERENCES

Preliminary Reports on Pneumatically Burst Gas Cylinder From NBS

Mechanical Properties Section to Office of Hazardous Materials

Operations - Department of Transportation.

a. September 26, 1975. "Thickness Measurements of Seamless Steel

Gas Cylinder T-0967," memo report from B. W. Christ to P. Seay,

OHMO-DOT

.

b. December 10, 1975. "Results of Tensile Tests on 3AA Seamless

Cylinder No. T-0967," memo report from J. H. Smith to

A. J. Mallen, OHMO-DOT.

c. January 22, 1976. "Evaluation of Steel Gas Cylinder T-0967,"

memo report from B. W. Christ to A. J. Mallen, OHMO-DOT.

d. June 18, 1976. "Preliminary Report on Pneumatically Burst

Gas Gylinder T-0967," memo report from B. W. Ghrist to

A. J. Mallen, OHMO-DOT.

Specification DOT 3AA Gode of Federal Regulations, Title 49.

Transportation, p. 418, 1975. U.S. Government Printing Office,

Washington, D.C. 20402.

American Iron and Steel Institute (AISI) Steel Products Manual.

a. Garbon Steel: Semifinished for Forging; Hot Rolled and Cold

Finished Bars; Hot Rolled Deformed Concrete Reinforcing Bars.

b. Alloy Steel: Semifinished; Hot Rolled and Cold Finished Bars.

ASTM E 45-74, "Standard Recommended Practice For Determining the

Inclusion Content of Steel," American Society for Testing and

Materials, 1916 Race Street, Philadelphia, Pennsylvania 19103.

ASTM E 23-72, "Standard Methods For Notched Bar Impact Testing of

Metallic Materials," American Society For Testing and Materials,

1916 Race Street, Philadelphia, Pennsylvania 19103.

ASTM E 399-72, "Standard Method of Test For Plane Strain Fracture

Toughness of Metallic Materials," American Society For Testing

and Materials, 1916 Race Street, Philadelphia, Pennsylvania 19103.

D. Broek: "Fracture of Structures - Pressure Vessels and Piping,"

Chapter 15.2 of Elementary Engineering Fracture Mechanics , Noordhoff

International Publishing, Leyden, The Netherlands, 1974.

W. A. Maxey: "Fracture Initiation, Propagation and Arrest," pp. J-1

to J-31 in Fifth Symposium on Line Pipe Research, Pipeline Research

Committee of American Gas Association, Houston, November, 1974.

20
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''• ' '-'<-01^’^ .

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, /a nM'-i>iuLB^*^‘*'moKPL7M ," • 3«unef-

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\:.x^j).^^cu o' , ^ !-','.c’ v'ji.niaiii'nT” ;»1 ^nal

I j;.-i *,i ... .J y:i<.,,.i<-y.t.. omm WP'O-t'.’Viio,

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• 'i’ • -i f .!( I T

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-.uH . ,:/ cit-3tjr4

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... XI.. I hi'H xjMl -V*»T to'; v .''si nr. -j i , a i ni \

"

.<(. I'i*': .0 I-, ».v; v>no‘)M .fii'riqi^bUidT ooo« dJ^.r'

'j-.u-aoM yni:a*l *io3 .i.-tT lo bo.fluM t ^^r^A. 7'.;.. 'J 'Hoi Sn.>8 n«-4 .' 0asnH;ji.io‘<'

.^1} .',('! 8 tbiiv i «/ ft i».'rttH fi 't li

i

.' dtb I. r j' an ... ssut^j 8 fao.t --»i

: '1 >ow •i.teeo’/ r» i«„:,^-.,-iT - 'iaTutO!jn;ja ..To •• J •.’•' >1^033' .•:

• '' * “iV ^..-tX t;>cgX .ti;a.rjb»i'||. t< '

. 'i»JqBJID

, '^laXaadu*'' oaT »,a3hy»il - XEno,. 'fi/i'xaiJnl-.

‘I'i '’,.'*-wytxA o';'-.'iJ' 1j./t'r o .A .V; 9nli9<ifVi 'vrtXJ no- ’cjm ;<cwjw^t; '’••

'<'*-T. 03

vfW * ,wo' 3«$x>^ liVoX .u.iDoiMiA ^.0 i. • - III/-. }b^ 'Hih/mi'I
9. REFERENCES (continued)

F. Erdogan: "DuctiJe Fracture Theories For Pressurized Pipes and

Containers," The International Journal of Pressure Vessels and

Piping, Vol. 4, pp. 253 - 283, October, 1976.

10. The Surface Crack : Physical Problems and Computational Solutions,

J. R. Swedlow, Editor. The American Society of Mechanical Engineers

345 East 47th Street, New York 10017. 1972.

11. E. S. Folias: "The Stresses in a Cylindrical Shell Containing an Axial Crack," ARE 64-174, Aerospace Research Laboratories (1964).

12. A. M. Sullivan and J. Stoop: "Some Fracture Mechanics Relationships

for Thin Sheet Materials," NRL Report 7650, December 21, 1973.

Strength of Materials Branch, Metallurgy Division, Naval Research

Laboratory, Washington, D.C.

21 -
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r'jd(T?.->DwU' Jfi'M ",6ifil'3i;)aH "iod' »oiNv>*4»« r«,vfcv* ,«4»i9jy>ci ., %“K aXici- ; >dt.M > .- fiJanViJa . .r,.n .jwt^filrfafcW ,yia.i»i,70<liiu3 ,A J
LIST OF FIGURE CAPTIONS

[la. Overall view of pneumatically burst gas cylinder, showing both halves

I of the neck. An intact cylinder is included for comparison.

!lb. Overall view of pneumatically burst gas cylinder, showing the bottom,

i Regions a and a’ indicate the fracture origin.

! 2a. View of Department of Transportation stampings on neck of gas

j cylinder T-0967. Note the crack arrest in this half of the neck.

i 2b. View of identifying* stampings on neck half of gas cylinder T-0967

opposite to that shown in Figure 2a.

'3. Markings on the bottom of cylinder T-0967, showing the heat code, V18.

! 4a. View of fracture surface showing the side of the fracture origin

labeled a’ in Figure lb. The elliptical ends of the fracture origin

and the. bright region along the inside wall are characteristic

features of a part-through crack.

' 4b. View of the same side of the fracture origin shown in Figure 3a. Note

the step in the fracture surface and the lips of raised metal on the

external surface.

5. View of the side of the fracture origin labeled a in Figure lb, with

I a schematic diagram showing dimensions.

6. The stepped region of the fracture origin, contrasting the smoothness

of the step with the fibrous nature of the adjacent fracture surface.

7. Mating halves of the fracture origin placed together to give an

I approximate picture of the configuration before burst.

8. Fracture profile of Specimen IV in Figure 7. (a) Gouge in outside

wall at fracture profile. (b) and (c) Evidence of flowed metal

beneath gouge.

' 9. Fracture profile of Specimen II in Figure 7, showing a gouge with a crack in the bottom.

i

, 10. Rusted regions on the external surface near the fracture origin.

1

j; 11. Fracture surface near neck of cylinder, showing fibrous fracture and

I relatively undamaged threads.

i 12. Fracture profile of Specimen I in Figure 7, showing wall thinning

[ near the fracture surface.

22
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13. Cross-section through wall in region of fast-running crack, showing

shear lip.

14. Fracture profile of Specimen IV in Figure 7. (a) Full- thickness

profile. (b) Corrosion product and rounded recesses in rusty and

discolored region of fracture origin. (c) Sharp recesses and

pointed dimples in the bright region along the inside wall below the

fracture origin.

15. Inclusions (a) Schematic representation and nomenclature (b) Inclusions

as seen on Face A in Figure 15a. (c) Inclusions as seen on Face B in Figure 15a. (d) Inclusions as seen on Face C in Figure 15a.

16. Knoop hardness versus distance below surface for the fracture

profile of Specimen IV in Figure 7.

17a. Absorbed energy versus temperature for longitudinal crack propagation.

17b. Absorbed energy versus temperature for circumferential crack

propagation.

18a. Lateral expansion versus temperature for longitudinal crack

propagation.

18b. Lateral expansion versus temperature for circumferential crack

propagation.

•I

; I

i,

'

23
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LIST OF TABLES

I Chemical Composition v

II Residual Elements

III Tensile Test Results On Longitudinal Test Specimens

IV Wall Thickness Measurements

V Comparison of Charpy Impact Test Results on Gas Cylinder T-0967

With Requirements of Proposed ISO Specification

VI Notched Tensile Test Results on Longitudinal Test Specimens

VII Tensile Test Results On Circumferential Test Specimens

VIII Results of Stress Analysis For Axial, Part-Through Cracks

IX Plane Stress Fracture Toughness, Of Quenched And Tempered

Carbon-Manganese Steel Inferred From Half-Size Charpy

Specimens

24 -
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CHEMICAL

COMPOSITION i 1 o

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TABLE II

RESIDUAL ELEMENTS

\

Element

Results* From Gas

Cylinder T-0967

(Weight Percent)

AISI Maximum

Limits of Residual

Elements (Reference 3)

(Weight Percent)

Copper 0.17 0.35

Nickel 0.13 0.25

Chromium 0.13 0.13

Molybdenum 0.02 0.06

Titanium 0.03 "h

Aluminum 0.03 +

Nitrogen 0.013 +

Oxygen 0.015 +

* Chemical analysis carried out by Chicago Spectro Laboratory,

Chicago, Illinois.

+ Unspecified.
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TENSILE

TEST

RESULTS

ON

LONGITUDINAL

TEST

SPECIMENS Ii

,)

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TABLE IV

Wall Thickness Measurements

Connection Ribbon Section

Bottom Longitudinal Crack

Arrest Side (B)

Top Longitudinal Crack

Arrest Side (C')

Near top 0.230 inches Near top 0.239 inches

0.220 0.228

0.224 0.230

0.227 0.225

Near bottom 0.241 0.242

Near bottom 0.241

Section Connected to Bottom

Bottom Longitudinal Crack Bottom Circumferential Crack

Arrest Side (B') Arrest Side (With Gouge (A')

Near top 0.230 inches Near top 0.249 inches

0.229 0.258 just above

0.229 0.257 just below

0.234 0.252

Near bottom 0.239 0.251

0.253

0.254

Near bottom 0.260

Section With Top Grack Arrest

Side Opposite Gouge (A) Top Longitudinal Crack

Arrest Side (C)

Near top 0.248 inches Near top 0.235 inches

0.255 0.227

0.257 just above origin 0.228

0.255 just below origin

0.253

0.253

0.249

0.243

Near bottom 0.250

Near bottom

0.235

0.238

0.240

0.245

Typical Bottom Thickness Measurements: 0.659, 0.654, 0.637-inches.
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TABLE VI

NOTCHED TENSILE TEST RESULTS ON

LONGITUDINAL TEST SPECIMENS*

Specimen

Number

Percent

Reduction

in Area

Tensile

Strength

(psi)

NL-1 6.0 116,700

NL-3 8.0 111,400

NL-4 15.0 113,100

Average 9.7 113,700

* Nominal Test Specimen Dimensions:

Gage Length

Gage Width

Gage Thickness

Notch Depth

Root Radius of Notch

Notch Angle

2.0 - inches

0.75 - inches

0.25 - inches

0.375 - inches

0.001 - inches, max.

60°

TENSILE

TEST

RESULTS

ON

CIRCUMFERENTIAL

TEST

SPECIMENS

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OF

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ANALYSIS

FOR

AXIAL,

PART-THROUGH

CRACKS 11 1 '!1!II1iI h

IN i

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TABLE IX

PLANE STRESS FRACTURE TOUGHNESS OF

QUENCHED AND TEMPERED CARBON-MANGANESE

STEEL* INFERRED FROM HALF-SIZE CHARPY SPECIMENS

Crack

Propagation

Direction

Charpy

Energy

(f t-lb/ in^)

Kc = (12E

psi - /inches

Longitudinal 100 190,000

Circumferential 372 366,000

* Tensile Strength = 99,800 psi
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Figure la. Overall view of pneumatically burst gas cylinder,

showing both halves of the neck. An intact

cylinder is included for comparison.

Figure lb Overall view of pneumatically burst gas cylinder

showing the bottom. Regions a and a' indicate

the fracture origin.
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View of Department of Transportation stampings

on gas cylinder T-0967. Note the crack arrest

in this half of the neck. X 1

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Figure 3. Markings on Che bottom of cylinder T-0967,

showing the heat code, V18. X 1
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Figure Aa. View of fracture surface showing the side of

the fracture origin labeled a’ in Figure lb..

The elliptical ends of the fracture origin

and the bright region along the inside wall

are characteristic features of a part-through

crack.

Figure 4b. View of the same side of the fracture origin

shown in Figure 4a. Note the step in the

fracture surface and the lips of raised metal

on the external surface.

0.232

inches

(OOJ 7&~o7Qg

Figure

5.

View

of

the

side

of

the

fracture

origin

labeled

a

in

Figure

lb,

with

a

schematic

diagram

showing

dimensions.

Figure 6. The stepped region of the fracture origin,

contrasting the smoothness of the step

with the fibrous nature of the adjacent

fracture surface. ^ ^

Figure 7. Mating halves of the fracture origin placed

together to give an approximate picture of

the configuration before burst. X 1
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Figure 12. Fracture profile of Specimen I in Figure 7 showing wall thinning near the fracture

surface. X 10

Figure 13. Cross-section through wall in region of

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