NAT'L INST. OF STAND & TECH R.I.C.

A 111 DM 2^3354

NATIONAL INSTITUTE OF STANDARDS &

TECHNOLOGY

Research Information Center

Gaithersburg, MD 20899

IR 88-3791

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PUBLICATIONS

NBSIR 88-3792

Investigation into the Ashland

Oil Storage Tank Collapse

on January 2, 1988

Center for Building Technology

National Engineering Laboratory

Metallurgy Division

Institute for Materials Science and Engineering

6(Jr*EA\J Of *

S. DEPARTMENT OF COMMERCE

TIONAL BUREAU OF STANDARDS

Research Information Center

National Bureau of Standards

Gaithersburg, Maryland 20899

NBSIR 88-3792 aj

Q.O ( 00

J Investigation into the Ashland ^

Oil Storage Tank Collapse

on January 2, 1988

John L. Gross

Felix Y. Yokel

Richard N. Wright

A. Hunter Fanney

Center for Building Technology

National Engineering Laboratory

John H. Smith

George E. Hicho

T. Robert Shives

Metallurgy Division

Institute for Materials Science and Engineering

Issued June 1988

U.S. DEPARTMENT OF COMMERCE

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

ABSTRACT

On January 2, 1988, a four million gallon capacity oil storage tank at the

Ashland Petroleum Company Floreffe Terminal near West Elizabeth, Pennsylvania,

collapsed as it was being filled to capacity for the first time since it was

reconstructed at the site. The tank had been dismantled in Cleveland, Ohio,

after more than 40 years of service and reconstructed at the Floreffe site in

1986. The National Bureau of Standards (NBS) conducted an independent

investigation of the physical causes of the Ashland Tank collapse. Data were

obtained from NBS field observations, laboratory and analytical studies, from

the investigation of the Pennsylvania Tank Collapse Task Force appointed by

the Governor of Pennsylvania, and from the Battelle Columbus Division

investigation sponsored by the Ashland Petroleum Company. The cause of the

failure was determined to be brittle fracture, initiating from a flaw existing

prior to the reconstruction of the tank. Complete rupture of the tank shell

occurred because the steel of the shell was of inadequate toughness at the

operating temperature to prevent brittle fracture propagation. The steel did

not meet the fracture toughness requirements of the American Petroleum

Institute Standard 650, 1984, "Welded Steel Tanks for Oil Storage," which was

effective at the time of reconstruction of the tank. The collapse shows the

importance of using steel with sufficient fracture toughness to prevent

propagation of a brittle fracture in tanks whose sudden failure would mean

unacceptable human, environmental or economic losses.

Keywords: Brittle fracture; collapse; failure investigation; fracture

analysis; steel; tanks; welded steel tanks

iii
EXECUTIVE SUMMARY

Introduction

On January 2, 1988, a four million gallon capacity oil storage tank at the

Ashland Petroleum Company Floreffe Terminal near West Elizabeth, Pennsylvania,

collapsed as it was being filled to capacity for the first time since it was

reconstructed at the site. As a result of the collapse, about one million

gallons of oil spilled into the Monongahela River approximately 25 miles

upstream from Pittsburgh. The spill contaminated the water supply of many

municipalities that draw their water from the Monongahela and Ohio rivers.

The Ashland tank that failed was a welded steel tank that was reported to have

been in service near Cleveland, Ohio, for over 40 years. It was reported to

have been dismantled in 1986, reconstructed at the Ashland Floreffe Terminal

site, and placed in service in August 1987. The tank was 120 feet in

diameter and 48 feet high and had a column- supported shallow conical roof and

a flat bottom. The tank rested directly on a crushed limestone foundation and

a compacted soil fill.

Congressman Doug Walgren, the Fire Marshal of Allegheny County, and the

Governor of Pennsylvania requested the National Bureau of Standards (NBS) to

conduct an independent technical investigation into the cause of the collapse.

Data were obtained from: NBS field observations, laboratory and analytical

studies; the investigation of the Pennsylvania Tank Collapse Task Force

appointed by the Governor of Pennsylvania; and the Battelle Columbus Division

investigation sponsored by the Ashland Petroleum Company.

The objective of the NBS investigation was to determine the most probable

cause of failure of the tank. The investigation included: on-site

investigation and debris identification, subsurface exploration and foundation

evaluation, structural analysis, thermal analysis, materials characterization,

fracture analysis, review of procedures and methods used in dismantling,

transporting and reconstructing the tank, review of documentation related to

terminal operation, review of eyewitness accounts of the collapse, review of

applicable standards, and determination of conformance with them.

IV
The NBS investigation was limited to the technical aspects of the failure and

determination of properties , practices and procedures which may have

contributed to the failure. Issues specifically not addressed in this

investigation include assignment of responsibility for the failure,

performance of the spill containment system, and conditions of other oil

storage tanks.

Description of Conditions at Failure

The oil level in the reconstructed tank was reported never to have exceeded 37

ft prior to January 2, 1988 when the tank collapsed. On that morning Ashland

personnel began to fill the tank to 46 ft. The Ashland terminal operator

reported that at approximately 5 pm the level of oil in the tank was 45 ft

10 in. He then returned to the office when, within 5 minutes of taking the

reading, the tank collapsed.

The NBS team of investigators arrived at the site of the collapse on January

6, 1988. The site appeared to be undisturbed, except for removal of spilled

oil. The tank wall had split completely open and the once-cylindrical shell

had unwrapped until it was nearly straight. The tank shell came to rest on

top of a dike approximately 120 ft from its original location. The shell was

separated from the tank bottom which, except for two large tears in the lap

welded seams, remained relatively intact. The tank top was severely crumpled

with several large tears but was in one piece. The top remained attached to

the shell opposite the vertical split. Debris from the roof support system

lay scattered mostly within the dike area.

The flat surface of the fracture of the shell indicated that it was a result

of brittle fracture. Chevron- like markings were evident on the fracture

surface. These pointed to a flaw extending about two-thirds of the way

through the plate thickness and located about 8 ft above the base of the tank,

just below the horizontal weld between the first and second shell courses, and

just to the left (looking from the outside of the tank) of a vertical seam in

the second course. The chevron markings, pointing to the flaw from above and

below, are convincing evidence that the fracture of the shell initiated at

this flaw.

v
Analysis Results

NBS metallurgical investigations defined the properties of the base metal of

the shell steel (where it was unaffected by welding) , the weld deposits from

original construction and reconstruction, and the heat-affected zones of the

base metal adjacent to welds.

Battelle conducted analyses of the material in the vicinity of the flaw and

made results available to NBS. The flaw had a length of 0.74 in and was

located in the base metal at the top of the first course of the shell just

below the old weld deposit. The flaw lay near the interior surface of the

shell, in a vertical plane perpendicular to the shell surface. A carburized

layer on the surface of the flaw showed that it was created in a cutting

operation; this indicates the flaw existed prior to the reconstruction at

Floreffe. The base metal adjacent to the flaw showed a hardness substantially

higher than that of normal base metal indicating embrittlement caused by the

cutting that produced the flaw and by subsequent welding adjacent to the flaw.

Battelle made residual stress measurements in the base metal adjacent to a Tweld (intersection of a vertical and a horizontal weld) remote from the

fracture but similar to the T-weld at which the fracture initiated. The

maximum measured residual tensile circumferential stress was 15 kilo-pounds

(thousand pounds) per square inch (ksi)

.

The stress in the tank shell at the flaw at the time of failure was about 30

ksi based on the measured residual stress and the stress NBS calculated from

the level of oil in the tank. The upper limit would be the yield point of

the steel, about 34 ksi, at which the steel would undergo large, inelastic

deformations, and the lower limit about 15 ksi if the residual stress were

extraordinarily small. According to NBS analysis, the temperature of the

shell steel was about 32° F at the time of failure.

Prior experience with failures -- Liberty ships which split open in World War

II and welded tank failures in the 1940s -- includes many brittle fractures

initiated in steels similar to that of the Ashland tank at similar flaws and

stresses. Extensive laboratory studies conducted in the United States, United

Kingdom and Japan show that brittle fractures initiate at stresses well below

vi
the yield point in base metal embrittled by welding near flaws of dimensions

similar to that found in the Ashland tank. Stress in the Ashland tank at the

flaw where the fracture initiated approached the yield point. Nearby stresses

due to the oil in the tank were about half the yield point. These stresses

are consistent with initiation of brittle fracture in the embrittled zone and

continued fracture propagation beyond the embrittled zone because the shell

temperature was far below its ductile to brittle transition temperature as determined by drop-weight nil-ductility and Charpy-V-notch tests.

Fracture mechanics analysis showed that the fracture toughness of the base

metal exceeded the calculated stress intensity due to the flaw, the stresses

due to oil in the tank and residual stresses. However, the increased hardness

measurements in the heat-affected zone of the base metal adjacent to the flaw

indicate a marked reduction of the fracture toughness from that of the

unaffected base metal. This embrittlement allowed initiation of the fracture

at the stress intensity in the vicinity of the flaw at the time of fracture.

Conformance with Standards

The nationally recognized current standard for above-ground, atmospheric,

welded steel tanks is the American Petroleum Institute Standard 650, Welded

Steel Tanks for Oil Storage, Seventh Edition, Revision 1, 1984 (hereafter

denoted as API 650) . API 650 does not explicitly address the reconstruction

of existing tanks. It does, however, include provisions for the use of

"materials on hand which are not identified as complying with any listed

specification." In NBS ' s opinion, API 650 can be used for reconstruction of

a tank when documentation is not available for the steel. In these

circumstances, API 650 requires measurements of steel properties.

The steel of the tank shell did not conform to any of the steels allowed in

API 650. Its chemical and mechanical properties correspond to ASTM A 283

grade D steel. API 650 allows use of ASTM A 283 grade C steel, which has a lower tensile strength. API 650 specifies toughness requirements. For the 5°

F design temperature conditions of the Floreffe site, ASTM A 283 grade C steel

should be semi-killed (deoxidized) and possess a Charpy-V-notch toughness of

15 ft- lb. The Charpy-V-notch toughness for specimens taken from the tank was

6 ft-lb or less for temperatures below 40° F. Thus, the toughness of the

vii
material was seriously deficient.

Conclusions

The failure of the tank was caused bv a brittle fracture that initiated at a flaw in the tank shell about 8 ft above the base and lust below the horizontal

weld between the first and second courses of steel. The fracture propagated

to complete rupture of the tank shell because the steel lacked sufficient

toughness to arrest a propagating fracture at the temperature and stresses

existing in the tank shell.

Failure was reported to have occurred as the tank was being filled to capacity

for the first time at the Floreffe site. The temperature of the shell was

about 32° F, and the calculated stress at the initiating flaw about 13.7 ksi.

These conditions are fully consistent with initiation and complete propagation

of brittle fracture considering the flaw size, embrittlement of adjacent

steel, presence of residual stresses, and the low fracture toughness of the

shell steel at 32° F. The shell steel did not meet the fracture toughness

requirements of API 650.

The flaw at which the fracture initiated was determined bv Battelle to be

present prior to the reconstruction of the tank at the Floreffe site.

Deposits on the surface of the flaw indicate it was formed during a cutting

operation.

Although the weld quality did not conform to API 650, the weld quality was

not a contributing factor in the initiation or propagation of this fracture.

However, welding adjacent to the flaw contributed to embrittlement of the

metal

.

Foundation conditions were investigated to determine if instability or

irregular settlement of the foundation might have increased stresses in the

she 1 1 . No evidence was found of foundation instability or excessive

settlement

.

viii
Lessons Learned

This study focused on the tank that collapsed at the Floreffe Terminal on

January 2, 1988, and on the standards applicable to its dismantling and

reconstruction. In spite of this limited scope, lessons of general

significance have been learned from this failure, and relearned from earlier,

similar failures.

The catastrophic failure of the Ashland tank resulted from a flaw that might

have been, but was not, detected in tests for welding quality, and the use of

steel that did not possess sufficient fracture toughness to arrest a fracture

that initiated at a localized flaw. While there are many reasons for control

of weld quality, it is not feasible to guarantee the absence of localized

defects by control of fabrication and detailed inspection for defects during

fabrication. Neither is it feasible to guarantee the absence of localized

defects during service by recurrent inspections. Therefore, it is advisable

to use steel of adequate toughness to provide "leak before break" fail-safe

behavior. If a crack is initiated, due to a flaw or local damage in service,

in a tank with steel of adequate fracture toughness, the crack will arrest,

the tank will leak, normal spill-control measures will confine the spread of

contents, and the tank can be repaired, all without catastrophic consequences.

Available standards should be reviewed to identify those calling for

sufficient fracture toughness to prevent catastrophic brittle fracture at

design stresses and temperatures. This study found the provisions of API 650

(1984) to be sufficient for the structural steels in its Materials Group I.

Compliance with sufficient fracture toughness provisions should be required

for steel to be used in construction or reconstruction of tanks, or for

changes in service conditions.

There is concern for the safety of existing tanks whose catastrophic failure

would cause unacceptable human, environmental or economic losses. If

documentation exists to show that the steel of an existing tank meets

sufficient fracture toughness provisions, the concern will be alleviated. If

documentation shows inadequate fracture toughness, remedial actions should be

taken. Possible remedial actions include conversion to adequately high

temperature service, retirement, or installation of crack arresters.

IX
Test and assessment protocols should be developed for assessing the

safety of potentially hazardous existing tanks that lack

documentation for definition of the fracture toughness of their steel

fracture

adequate

x
ACKNOWLEDGEMENTS

Special appreciation is extended to Mr. Leonard C. Smith, NBS Chief

Metallographer for exceptional skill and care taken in the metallographic

evaluation and hardness tests conducted for this investigation. In

particular, the specific choice of areas in the microstructures to be examined

and the skill and judgement used in presenting the results were a key factor

in leading to the conclusions reached in this report.

Thanks also go to Dr. William Stone of the Structures Division of the Center

for Building Technology for performing the computer modeling, analysis and

post-processing of the symmetric model of the tank for the foundation

settlement analysis.

A special thanks goes to Mr. F. Keith Mackley and Mr. Raymond J. Mele of the

National Engineering Laboratory (NEL) Graphics for their work in producing the

drawings and labeled photographs and for performing the layout for the report.

xi
-

.
TABLE OF CONTENTS

LIST OF FIGURES xv

LIST OF TABLES xviii

1.0 INTRODUCTION 1

1 . 1 Background 1 1.2 Objective and Scope of Investigation 2

1.3 Cooperation with Other Organizations 3

1.4 Organization of the Report 3

2.0 DESCRIPTION OF THE TANK 9

2.1 General Description 9

2.2 History of Operation 10

2.3 Dismantling of the Tank 11

2.4 Reconstruction of the Tank 12

3.0 DESCRIPTION OF THE COLLAPSE 19

3.1 Operation of Tank No. 1338 19

3.2 Conditions Preceding the Collapse 19

3.3 Witness Accounts of the Collapse 20

3.4 Observations of the Debris 21

3.5 Description of the Tank Wall Fracture 22

4.0 FOUNDATION STUDY 35

4.1 Objective of Foundation Study . 35

4.2 Description of Foundation and Site 35

4.3 Foundation Analysis 43

4.4 Summary 51

5.0 STRUCTURAL ANALYSIS 59

5.1 Stress Analysis of the Tank 59

5.2 Foundation Settlement Analysis 60

6.0 THERMAL ANALYSIS OF THE TANK WALL 67

6.1 General 67

6.2 Ambient Temperature and Wind Data 67

6.3 Temperature of the Oil in the Tank 68

6.4 Thermal Analysis 68

xiii
7.0 METALLURGICAL INVESTIGATION ................... 75

7.1 Introduction .................. 75

7.2 Materials Properties Characterization ............ 76

7.3 Metallographic and Hardness Evaluation ...... 88

7.4 Characterization of the Fracture Surface Defect ....... 91

7.5 Stress Analysis . . ............. 93

8.0 ANALYSIS OF TANK FAILURE 133

8.1 Scope of Tank Failure Analysis 133

8.2 Assessment of Fracture Initiation and Arrest 133

8.3 Fracture Analysis 136

8.4 Discussion of the Fracture Analysis Results 141

8.5 Evaluation of Fracture Propagation and Arrest . 144

9.0 CONFORMANCE WITH STANDARDS 151

9.1 General ........ ..... 151

9.2 Materials ........ 151

9.3 Shell Design 153

9.4 Inspection .... ........... 154

9.5 Testing 155

9.6 Foundation ........ 155

10.0 CONCLUSIONS 159

10.1 Cause of Tank Collapse 159

10.2 Lessons Learned 160

REFERENCES ......... . 162

APPENDIX 166

xiv
LIST OF FIGURES

1.1 Aerial View of the Ashland Tank Collapse, January 4, 1988 .... 5

1.2 Site of the Tank Collapse 6

1.3 Plan of the Ashland Floreffe Terminal

2.1 General Description of the 96,000 Barrel Tank 15

2.2 Plan View of Roof Support System 16

2.3 Column Base Resting Directly on Tank Bottom 17

2.4 Roof Girders Supported on Top of Columns 17

2.5 Roof Rafter Attachment to Tank Shell 18

3.1 Product Level in Tank No. 1338 23

3 . 2 Maximum and Minimum Daily Temperatures at

Allegheny County Airport 24

3.3 Aerial View of the Collapse Site 25

3.4 View of Collapsed Tank No. 1338 from Ground Level 26

3.5 View of Tank Bottom with Tank Wall and Roof in Background .... 27

3.6 Tear in Tank Bottom at Fillet Attaching First Shell

Course and Bottom Plate 28

3.7 Roof Framing Debris Scattered on Tank Bottom 29

3.8 Tank No. 1367 Dented in the Collapse of Tank No. 1338 30

3.9 Right Side of Fracture 31

3.10 Chevrons in Fracture Surface 32

3.11 Dark Area on Right Fracture Surface 33

3.12 Approximate Path of Fracture in Tank Wall 34

4.1 Cross Section of Tank Foundation 53

4.2 Topography of Foundation Surface Surveyed After

Removal of the Base Plate 54

4.3 Aerial Photograph of Tank Foundation Taken Prior to

Removal of the Base Plate 55

4.4 Location Map for Soil Exploration 56

xv
4.5 Schematic Drawinig of Soil Profile Developed by

Interpolation Between Borings . .......... 57

4.6 Assumed Shape of Normalized Settlement Profile of Tank ..... 58

5.1 Axisymmetric Analysis Model of Tank Wall 62

5.2 Circumferential Membrane Stress in the Tank Wall 63

5.3 Meridional Bending Stress in the First Two Shell Courses .... 64

5.4 Symmetric Model of Tank Wall for Foundation

Settlement Analysis 65

5.5 Circumferential Stresses in the Tank Wall at a Height of 8 ft due

to an Assumed Differential Settlement of 0.01 ft 66

6.1 Temperature of the Oil in Tank No. 1338 74

7.1 Location of Primary Vertical Fracture and NBS Test Samples . . . 107

7.2 Location of Test Specimens from First Shell Course ....... 108

7.3 Location of Test Specimens from Second Shell Course 109

7.4 Location of Test Specimens from Horizontal Weld ......... 110

7.5 Location of Test Specimens from Vertical Weld . Ill

7.6 Location of Samples for Chemical Analysis ... 112

7.7 Charpy-V-Notch Tests for First Shell Course - Transverse Orientation 113

7.8 Charpy-V-Notch Tests for First Shell Course - Longitudinal Orientation 114

7.9 Charpy-V-Notch Tests for Second Shell Course - Transverse Orientation. ........ .... 115

7.10 Charpy-V-Notch Tests for Second Shell Course - Longitudinal Orientation. 116

7.11 Single Edge-Notch Bend Specimen for Embrittlement Study 117

7.12a Embrittlement Zone at Notch Tip in Single-Edge

Notch Bend Test 118

7.12b Hardness of Embrittled Zone in Single-Edge

Notched Bend Specimens ..... 118

7.13 Results of Notched-Bend Test 119

xvi
7.14 Microstructure of Steel Plates 120

7.15 X-ray Radiograph of Horizontal Weld after Surface Grinding . . . 121

7.16 Microstructure and Hardness of Horizontal Weld Metal- 122

7.17 Microstructure of Horizontal Weld Metal 123

7.18 Defects in Horizontal Weld Metal 124

7.19 Defects in Horizontal Weld Metal after Grinding Welds 125

7.20 Defects in Vertical Weld Metal 126

7.21 Location of Fracture Surface Defect .... 127

7.22 Fracture Surface Defect . 128

7.23 Location of Fracture Surface Defect From Outside of Tank .... 129

7.24 Defect on Fracture Surface 130

7.25 Details of Sectioning to Determine Location of Defect 131

7.26 Transverse Section of Fracture Surface Defect Showing Exact

Location of Defect ..... 132

8.1 Fracture Analysis Diagram (FAD) 148

8.2 Calculated Stress - Defect Size Relationship for

Through-Wall Defect 149

9.1 Defects in a Horizontal Weld between the First and

Second Shell Courses ... 158

xvii
LIST OF TABLES

2.1 Shell Plate Thicknesses . ...... ........ . 14

4.1 Shell Displacements 52

6.1 Surface Weather Observations - January 2 , 1988 ......... 72

6.2 Average Tank Wall Temperature at Elevation 8 ft ........ . 73

7.1 Chemical Composition of Tank Steels and Weld Metal ....... 95

7.2 Tensile Properties of Plate Steels ......... . 96

i 7.3 S ummary of Charpy-V-Notch Test Results - First Shell Course - Transverse Orientation . ....... 97

7.4 Summary of Charpy-V-Notch Test Results - First Shell Course - Longitudinal Orientation .................... 98

7.5 Summary of Charpy-V-Notch Test Results - Second Shell Course - Transverse Orientation 99

7.6 Summary of Charpy-V-Notch Test Results - Second Shell Course - Longitudinal Orientation .................... 100

7.7 Ductile- to-Brittle Transition Temperatures Determined from

Charpy-V-Notch Tests ....... 101

7.8 Drop-Weight (NDT) Test Results ................. 102

7.9 Summary of Fracture Toughness Test Results 103

7.10 Summary of Notched-Bend Test Results 104

7.11 Summary of Hardness of Steel Plates ............... 105

7.12 Charpy-V-Notch Results from Weld Metal 106

8.1 Summary of Values Used in the Fracture Analysis . 146

8.2 Summary of Calculated Stress - Intensity Factors 147

9.1 Design Shell Plate Thicknesses ...... 157

xviii
1.0 INTRODUCTION

1 . 1 Background

On January 2, 1988, a four million gallon capacity oil storage tank at the

Ashland Petroleum Company Floreffe Terminal collapsed as it was being filled

(see Figure 1.1). The Ashland Floreffe Terminal is located near West

Elizabeth, Pennsylvania on the Monongahela River (see Figure 1.2). A plan of

the terminal indicating the location of the collapsed tank relative to Route

837 and the shoreline of the Monongahela River is shown in Figure 1.3. As a result of the collapse, an estimated one million gallons of oil spilled into

the Monongahela River approximately 25 miles upstream from Pittsburgh. The

spill contaminated the water supply of many municipalities that draw their

water from the Monongahela and Ohio Rivers.

The tank that failed was a reconstructed steel tank that had originally been

in service near Cleveland Ohio for over 40 years. It had recently been

dismantled and reconstructed at the Ashland Floreffe Terminal site near

Pittsburgh. The tank was 120 ft in diameter and 48 ft high and had a columnsupported shallow conical roof and a flat bottom. The tank rested directly on

a crushed limestone and compacted fill foundation.

The tank was being used to store No. 2 fuel oil and had been in service at its

new location since August of 1987. Oil was brought to the Ashland terminal by

barge on the Monongahela River and pumped into the storage tank. The failure

occurred as the tank was being filled to capacity for the first time.

Congressman Doug Walgren, the Fire Marshal of Allegheny County and the

Governor of Pennsylvania requested the National Bureau of Standards (NBS) to

conduct an independent technical investigation into the cause of the collapse.

A team of NBS investigators arrived at the site of the collapse on January 6,

1988. (Access to the site prior to this date was not possible due to the

presence of oil within the dike area surrounding the collapsed tank.) The

investigation was conducted under public law which gives NBS the authority to

conduct independent investigations into the causes of structural failures

which affect the general public and which have broad importance.

1
1.2 Objective and Scope of Investigation.

The objective of this investigation was to determine the most probable cause

of failure of the Ashland tank. The investigation included:

o On-site investigation and debris identification

o Sub- surface exploration and foundation evaluation

o Materials characterization

o Fracture analysis

o Review of procedures and methods used in dismantling, transporting and

reconstructing the tank

o Review of documentation related to terminal operation

o Review of eyewitness accounts of the collapse

o Review of applicable codes and standards and determination of

compliance with them.

The investigation was limited to the technical aspects of the failure and a determination of practices and procedures which may have been contributory to

the failure. Issues not addressed in this investigation include:

o Assignment of responsibility for the failure

o Performance of the spill containment system

o Recommendations for either new standards or changes to existing

standards to prevent future occurences of this type of failure.

2
1.3 Cooperation with Other Organizations

Following the collapse of the oil storage tank, the Governor of the

Commonwealth of Pennsylvania appointed a task force to investigate the

collapse and make recommendations to prevent further occurrences of similar

accidents. The Pennsylvania Tank Collapse Task Force (hereafter referred to

as the Pennsylvania Task Force) worked closely with NBS and shared information

related to the investigation. This information included: records provided to

the Pennsylvania Task Force by Ashland Petroleum Company (hereafter referred

to as Ashland) ; information obtained by staff investigators through personal

interviews; and test data obtained from soil samples taken under the direction

of the Pennsylvania Task Force. NBS also worked closely with the Advanced

Technology for Large Structural Systems (ATLSS) Research Center at Lehigh

University, which provided technical assistance to the Pennsylvania Task Force

on the structural and metallurgical aspects of the tank failure.

In addition, Battelle Columbus Division (hereafter referred to as Battelle)

,

contracted by Ashland to conduct an investigation into the cause of the

collapse, provided valuable metallurgical data and characterization of the

fracture surface . Under an agreement worked out among NBS , Ashland and

Congressman Walgren, both fracture surfaces were shipped to Battelle for

preservation and analysis. Thus, all information related to the fracture

surface was provided to NBS by Battelle. NBS investigators personally

observed the taking of samples by Battelle and the testing and fractography

conducted at the Battelle laboratory.

1 . 4 Organization of the Report

Chapter 2 contains a description of the tank and its history including

dismantling of the tank at its previous location and reconstruction at the

Ashland Floreffe Terminal.

Chapter 3 is a description of the collapse including activities and conditions

preceding the collapse, eyewitness accounts of the collapse, and observations

of the debris made during the initial site investigation.

3
Chapter 4 contains a description of the foundation and site and describes the

subsurface exploration and foundation analysis.

Chapter 5 presents the results of an analytical study to determine the

stresses in the tank.

Chapter 6 addresses the prediction of the temperature in the tank wall at the

time of collapse.

Chapter 7 presents the results of metallurgical tests conducted to

characterize both the base metal and weld metal. The tests include chemical

analyses, tensile tests, notch sensitivity tests, hardness tests, fracture

toughness tests and residual stress determination.

Chapter 8 presents several fracture analyses, based on the metallurgical test

results and tank shell stresses determined by analysis, to determine why the

fracture of the tank shell initiated and why it propagated to complete

rupture.

Chapter 9 addresses the issue of conformance with applicable codes and

standards, limited to those aspects relevant to structural performance.

Chapter 10 presents NBS's findings regarding the cause of failure and lessons

learned from the failure.

4
Figure 1.1 - Aerial View of the Ashland Tank Collapse, January 4, 1988.

(Photo Courtesy of Vince Musi -- Pittsburgh Press/SIPA)

5
6
Ashland Oil Co. - Floreffe Terminal

R.R Tracks

^-U.S. Gov't Harbor Lin*

Figure 1.3 - Plan of the Ashland Floreffe Terminal

7

2.0 DESCRIPTION OF THE TANK

2 . 1 General Description

A description of the tank and details of its history were compiled from a variety of sources. Since no drawings of the tank were available, the

physical description of the tank was reconstructed from examination and

measurements of the debris. In addition, no cutting or welding specifications

were available from the erector, therefore, the information which follows was

obtained by interviewing individuals with first-hand knowledge of the

procedures used in dismantling and reconstructing the tank. The Pennsylvania

Task Force conducted the interviews and gave NBS access to all of their

findings

.

The Ashland tank had a nominal capacity of 96,000 barrels or four million

gallons. It was a welded steel tank with a flat bottom and shallow conical

roof (see Figure 2.1). The roof was 3/16-in plate with lap welded joints. It

was supported by a system of rafters and girders which in turn were supported

by columns resting on the tank bottom. The cylindrical shell was 120 ft in

diameter and 48 ft high. It consisted of six shell courses, each

approximately 8 ft high and 32 ft long, which were butt welded. The thickness

of each shell course was measured by NBS and results are shown in Table 2.1.

Corrosion of the tank wall occurred over the service life of the tank so the

original plate thicknesses can only be assumed; Table 2.1 shows the assumed

original plate thicknesses. The shell thicknesses shown in Figure 2.1 are the

assumed original plate thicknesses before corrosion. The tank bottom was 1/4-

in lap welded, steel plate which was attached to the first shell course by a double fillet weld (fillet welds on both sides of the shell)

.

A structural framework supported the shallow conical roof. The framework

consisted of 120 roof rafters spanning in the radial direction, two rings of

column- supported girders and the tank shell (see Figure 2.2). There were five

columns supporting the inner ring of girders and 10 columns supporting the

outer ring. The columns consisted of two channel sections welded together so that the flange of a 9- in channel was welded to the back of the web of a 12-

in channel. The columns rested directly on the tank bottom and the load was

9
distributed by the "H" shaped column base shown in Figure 2.3. The girders

comprising the rings consisted of 15-in channel sections which spanned between

the columns. They were welded to a cap plate on top of each column as shown

in Figure 2.4. The girders formed five- and ten- sided polygons for the inner

and outer support ring, respectively. The rafters, which were 8 -in channels,

usually were connected to the girders by means of clip angles riveted to the

girder (original construction) and welded to the rafter. The rafters were

attached to the tank shell by means of plates as shown in Figure 2.5.

2 .

2

History of Operation

The tank that collapsed had been recently moved from a site near Cleveland,

Ohio, and reconstructed at the Ashland Floreffe terminal. It replaced an

aging tank designated as No. 1338. The old tank was approximately 115 ft in

diameter and 40 ft high and had a capacity of 74,000 barrels. It was reported

to be a riveted tank and had a recent history of leaking. An inspection by

Ashland determined that the old tank would be too expensive to repair and it

needed to be replaced.

A 96,000 barrel tank that was out of service was located at a site called

Whiskey Island near Cleveland, Ohio. This tank, designated as No. 16, had

been built about 1940 and had most recently been used to store No. 6 oil, a viscous petroleum product which is generally maintained at a temperature of

approximately 150° F. The tank was covered with about 3 inches of foam

insulation in 1981. In March 1986, the Whiskey Island tank was inspected to

determine if it was in good enough condition to replace the aging 74,000

barrel tank at the Floreffe terminal. The tank was visually inspected and the

thicknesses of the first and sixth shell courses were measured ultrasonically

.

The inspection report indicated that the tank shell was not severely corroded

and could be re-used. The bottom plate, however, was found to be badly

corroded and the report recommended that the bottom plate be replaced with new

steel. The decision was therefore made by Ashland to scrap tank No. 1338 at

the Floreffe site and replace it with tank No. 16 from the Whiskey Island

site

.

10
2.3 Dismantling of the Tank

Dismantling of tank No. 16 at Whiskey Island began in May 1986, and took less

than four weeks. First the insulation was removed by hand along the weld

seams. Then, starting from the bottom, the tank was cut into sheets. The

cuts were made just to the right of the vertical welds, looking from the

outside of the tank, and above the horizontal welds. For the vertical cuts, a series of cuts about two feet long were made and an uncut section of about

three inches was left to hold the plates together. This remaining uncut

portion of the tank wall was known by the welders as a "tack" . For the

horizontal cuts, a series of cuts was made, each four feet long, and again a three- inch "tack" was left. At the intersection of a horizontal and a vertical cut, a "tack" of about six inches was left to hold the plates

together. After cutting the first shell course, the scaffolding was moved up

to the second course and the procedure was repeated. This continued until all

of the tank wall had been cut. Cuts were made manually with no guide using an

acetylene torch. As each sheet was cut, the shell course number was marked on

the inside of the plate and the edges were marked with an identifying seam

number so the tank could be reassembled.

Next, a crane was set in place for dismantling the roof plate. While the

piece to be removed was being hooked onto the crane, a welder made the final

cut. The piece of the roof was then swung free and lowered to the ground.

After the roof had been removed, a door opening was cut in the tank wall for

the crane to drive through. The crane was then used from the inside of the

tank to remove the rafters and girders. After the roof framing was removed,

the welders removed the wall of the tank beginning with the top shell course.

The piece to be removed was attached to the crane by two clamps spaced about

eight feet apart. Working from the scaffolding on the outside of the tank, a welder cut the remaining "tacks" and the plate was swung free, lowered and

stacked on the tank bottom.

The sheets cut from the tank still had insulation over most of their outer

surface. To remove the remaining insulation, the sheets were placed on the

ground, insulation side up, and a rubber- tired tractor with a scraper blade

was used to remove the insulation. After the insulation was removed, the

11
sheets were loaded onto a flatbed truck and hauled to the Floreffe site.

2 . 4 Reconstruction of the Tank

Before the dismantled tank from Whiskey Island could be erected, the old tank

that it was replacing had to be dismantled. This work began in June, 1986,

and took several weeks. The old tank was hauled away and scrapped. A new

foundation was built where the tank had been located. It consisted of two to

three feet of compacted fill covered by a soil stabilization fabric, and 12 to

18 in of crushed limestone (a complete description of the tank foundation in

given in Chapter 4) . The fill was placed in 6 to 8 inch lifts and compacted.

The foundation work was completed during the month of July 1986.

Reconstruction of the tank began in August, 1986. Work began by laying a new

steel bottom. Lugs were welded to the tank bottom for aligning the first

shell course and to hold the first course in place as it was being tackwelded. Once the first course was tack-welded, the vertical seams were welded

from the outside. Next the horizontal seams were welded from the outside.

Once the outside welds were complete, the root pass (first pass) was cut out

from the inside of the tank and the inside welds were made.

Edge preparation prior to welding consisted of hand flame cutting with an

acetylene torch to produce a bevel. It was very clear that the no attempt was

made to remove the old weld. It was not possible to determine whether the

flame cut edges had been ground prior to re-welding. It was reported that, at

the intersection of the vertical and horizontal welds or "T-welds", extra weld

passes were made (extra passes were observed at several T-welds during the

site investigation) . Inspection of the welds consisted of visual inspection

or liquid penetrant inspection in which fuel oil was sprayed on the inside of

the welds and the outside was checked for leaks. No radiographic inspection

was performed while welding was • in progress.

Radiographic inspection was conducted by an independent laboratory after the

tank had been completely welded. A total of 39 radiographs were taken; 26 in

vertical welds, 4 in horizontal welds and 9 in T-welds. Of the 39

12
radiographs, 22 were "rejected" in accordance with the criteria in ASME

Section VIII, Paragraph UW-51 [2.1]. The defects reported by the radiographer

include slag inclusion, lack of fusion and lack of penetration. The

radiographer reported to Ashland that the defects "appeared to be in the old

weld metal" which had not been removed prior to re-welding. No repairs were

made to the defective welds. Tank erection was completed in mid-November

1986.

On March 12, 1987, elevation readings were taken on the lip around the bottom

of the tank at the seams. By May 1987, all piping to the tank had been

completed and the tank was partially filled with water to test for leaks.

Water was supplied by the West Penn Water Company and the tank was filled

using a fire hose placed into the top of the tank. It took approximately two

days to fill the tank to the 5 ft- level. A gage on the side of the tank was

used to measure how much water was in the tank. The water was then allowed to

sit in the tank for several days and the tank was checked daily for leaks at

the base and side. On June 11, 1987, elevations were again taken around the

base of the tank. No leaks were observed and the water was emptied from the

tank. In August 1987, Tank No. 1338 was put into service storing No. 2 fuel

oil. Settlement readings were again recorded on August 25, 1987, when Ashland

inventory records indicated 10 ft - 3 1/2 in of oil in the tank.

13
Table 2.1 - Shell Plate Thicknesses

Plate Thickness (in)

Measured by NBS

(to nearest 1/16)

Original

(assumed)

Tank Bottom 1/4 1/4

Shell Course 1 7/8 7/8

Shell Course 2 11/16 3/4

Shell Course 3 1/2 5/8

Shell Course 4 3/8 1/2

Shell Course 5 1/4 3/8

Shell Course 6 3/16 1/4

Tank Roof 1/8 3/16

14
3/16"

Roof £

15
Column

16
Figure 2.3 - Column Base Resting Directly on Tank Bottom

Figure 2.4 - Roof Girders Supported on Top of Columns

17
3/16" £ roof

18
3.0 DESCRIPTION OF THE COLLAPSE

3 . 1 Operation of Tank No. 1338

Tank No. 1338 was put into service at Ashland's Floreffe Terminal in August,

1987, when No. 2 fuel oil was first added to the tank. The first entry in

Ashland's Daily Inventory Report was on August 25, 1987, at which time the

product level was 10 ft 3-1/2 in. Fuel oil was brought to the terminal by

barge on the Monongahela River. The oil was pumped from the storage tank into

tank trucks for local distribution. The level of oil varied from day to day

as oil was added and removed. On the days that the Floreffe terminal was

operating, the product level in each storage tank was recorded for inventory

purposes. The product level for Tank No. 1338 is plotted in Figure 3.1 for

the period beginning August 25, 1987, and ending January, 1, 1988, the day

before the collapse. It can be seen that the product level had never exceeded

the 37-ft mark prior to the day the tank collapsed.

The daily high and low temperature recorded at the Allegheny County Airport,

which is approximately seven miles from the Floreffe Terminal (see map in

Figure 1.2), is shown in Figure 3.2. The lowest temperature experienced

during the history of operation of the reconstructed Tank No. 1338 was 12° F,

which occurred on January 2, 1988, the day of collapse.

3 . 2 Conditions Preceding the Collapse

At about 4:30 am on January 2, 1988, two barges containing No. 2 fuel oil

arrived by river tug at the Ashland terminal dock. An 8,500-barrel barge was

the first to be unloaded into Tank No. 1338. Ashland personnel determined

that there was more oil on the barges than could be put into Tank No. 1338.

It was decided to fill the tank to the 46-ft mark -- six inches below the

level at which the fire foam lines entered the tank. The 11,000-barrel barge

containing No. 2 fuel oil began unloading into Tank No. 1338 at about 11:30

am. The pumping rate was approximately 2000 barrels per hour which, for a 120-ft diameter tank, translates to a fill rate of approximately 1 inch every

5 minutes. The level of oil was read every half hour using the gage on the

19
side of the tank. At about 4:00 pm, gage readings began to be taken every ten

minutes. At approximately 5:00 pm the Ashland terminal operator recorded the

level at 45 ft - 10 in. He determined that it would take approximately 10

more minutes of loading to reach the 46-ft level. He then returned to the

office and, within 5 minutes after taking the reading, the tank collapsed.

At the time of collapse, it was still daylight (dusk), the weather report at

the Allegheny County Airport indicated clear skies and the temperature was

26° F. The level of oil in the tank at the time of collapse was approximately

46 ft - 0 in.

3 . 3 Witness Accounts of the Collapse

Descriptions of the collapse of Tank No. 1338 were given by four witnesses.

All of the witnesses, except the Ashland operator, were more than 500 yards

from the tank at the time of collpase. None of these witnesses was looking

at the tank at the instant of collapse. Rather, their attention was drawn to

the tank by the sound of the collapse which was described as a "loud rumbling

noise” similar to "thunder". Accounts indicated that the sound lasted between

3 seconds and 30 seconds. Witnesses saw the tank fall toward the north and a wave of "foam" or "white smoke" spill over Route 837. They also reported what

was variously described as "green mist," "a green cloud" or "chemical cloud"

which formed and then dissipated.

The Ashland operator who was standing next to the tank just minutes prior to

the collapse had not observed any unusual circumstances such as leaking

product or noises coming from the tank.

The accounts of the collapse can be summarized as follows:

o There were no warnings or evidence that a failure was imminent,

o The loud rumbling sound created by the collapsing tank lasted at least

several seconds

.

o There was no report of an explosion or fire.

20
3.4 Observations of the Debris

The NBS team of investigators arrived at the site of the collapse on January

6, 1988. Prior to that date, oil which had spilled from the tank flooded the

dike area (see Figure 1.1) and prevented safe access. By January 6, 1988,

most of the oil had been removed from the dike area of Tank No. 1338 as well

as from the adjacent dike areas. The site appeared to be otherwise

undisturbed at the time NBS investigators arrived.

As can be seen from the aerial view of the collapse site shown in Figure 3.3

(3.3), the tank wall had split completely open and the once-cylindrical shell

had unwrapped until it was nearly straight. The tank shell came to rest on

top of a dike approximately 120 ft and generally north of its original

location. A description of the fractured tank wall is given in the next

section. A composite photograph of the collapsed tank from within the dike

area is shown in Figure 3.4.

The wall was completely separated from the tank bottom. Figure 3.5 shows the

tank bottom in the foreground and the tank wall and portions of the roof in

the background. The bottom tore at the toe of the inside fillet weld

attaching the first shell course to the bottom plate leaving a 3-inch "lip"

attached to the tank wall (see Figure 3.6). This tear was ductile as evidenced by the approximately 45° failure plane. The tank bottom remained

relatively intact except for two large tears in the lap welded seams. The

edge of the bottom plate where the tank wall tore away had a "scalloped"

appearance as shown in Figure 3.5.

The tank roof was severely crumpled with several large tears but otherwise was

in one piece. The tank wall had torn away from the roof as it "unwrapped"

except for a portion which remained attached to the wall opposite the vertical

split. The tear was generally at the weld between the roof plate and the top

angle and was also generally ductile.

Debris from the roof support system lay scattered mostly within the dike area

as can be seen in Figure 3.3. A closer view of some of the debris on the tank

bottom is shown in Figure 3.7. The roof framing members were severely twisted

21
and many of the welds had failed.

Tank No. 1367, which was empty at the time of the collapse, was severely

dented on the side facing Tank No. 1338. Tank No. 1367 is partially in view

at the left in Figure 3.4 and is shown in Figure 3.8. An oil stain was

evident on this tank and on Tank No. 1339 extending as high as the roof line

which was approximately 48 ft for both tanks.

3 . 5 Description of the Tank Wall Fracture

One side of the fracture in the tank wall came to rest inside the dike area

near Tank No. 1366. It is shown in Figure 3.9. The fracture surface was flat

and perpendicular to the tank wall indicating that it was a brittle fracture.

Chevrons were evident on the fracture surface (see Figure 3.10). The chevrons

pointed to a dark area extending about two- thirds of the way through the plate

thickness (see Figure 3.11). The location of this dark area is just below the

horizontal weld, between the first and second shell courses, and just to the

left (looking from the outside of the tank) of a vertical seam in the second

course. The fracture was believed to have initiated at this point and

extended to the top and bottom of the tank wall. A diagram showing the path

of the fracture is shown in Figure 3.12. Note that the fracture neither

started nor propagated in the welds.

22
LEVEL

OF

OIL

IN

TANK

00

ULJ y—

<

Q

(U) 110 do 13A31

23

Figure

3.1

-

Product

Level

in

Tank

No.

1338
MAXIMUM

AND

MINIMUM

DAILY

TEMPERATURES

AT

ALLEGHENY

CO.

AIRPORT

T T 7

CO

CO

LD

<

Q

co

LO

C\J

co

(do) 3dfllVd3dlAI31

24

Figure

3.2

-

Maximum

and

Minimum

Daily

Temperatures

at

Allegheny

County

Airport
ftactura sur

fracture

Tank #1367

Tanl< # 1339

Figure 3.3 - Aerial View of the Collapse Site

(Photo by Pennsylvania Task Force)

25
Tank

#1367

auj| t|cnei/\|

9U(| lpVB|/\|

26

Figure

3.4

-

View

of

Collapsed

Tank

No.

1338

from

Ground

Level
Figure 3,5 - View of Tank Bottom with Tank Wall and Roof in Background

27
Figure 3 .

6

Tear in Tank Bottom at Inside Fillet Attaching First Shell

Course and Bottom Plate

28
29
Figure 3.8 - Tank No. 1367 Dented in the Collapse of Tank No. 1338

30
Figure 3 .

9

Right Side of Fracture

31
Figure 3.10 - Chevrons in Fracture Surface

32
Figure 3.11 - Dark Area on Right Fracture Surface

33
34

Figure

3.12

-

Approximate

Path

of

Fracture

in

Tank

Wall
4.0 FOUNDATION STUDY

4 . 1 Objective of Foundation Study

The observations of the debris of Tank No. 1338 described in Section 3.3

indicate that the tank shell had split and was separated from the tank bottom

plate by a bottom-plate failure adjacent to the weld between the bottom plate

and the shell. There were also tears in some lap welded seams of the bottom

plate which appear to have originated near the seam between the shell and the

bottom plate. The objective of the foundation study was to determine whether

a foundation stability failure or excessive foundation settlements could have

caused a failure of the bottom plate and its separation from the shell, and

whether shell distortions introduced by non-planar foundation settlements

along the tank perimeter could have contributed to the failure of the shell.

4 . 2 Description of Foundation and Site

4.2.1 Sources and Scope of Information

Information for the foundation study was derived from data provided by

Ashland, including construction drawings, records of field density tests taken

during the construction of a raised compacted fill base, and elevation

readings along the rim of the tank taken during partial filling of the tank,

and data provided by the Pennsylvania Task force, including topographic data

for the site and elevations of the foundation surface taken in a field survey

after removal of the bottom plate of Tank No. 1338, and data from a subsurface

exploration conducted after the tank collapse. The site was also visually

inspected by the NBS investigation team prior to, and during the subsurface

exploration.

The soil exploration and laboratory testing data used in the foundation

analysis were provided to NBS by the Pennsylvania Task Force. Boring logs,

oedometer tests from boring B-4 and B-5, and data from unconsolidated

undrained triaxial compression tests are presented in the Appendix. All other

data can be obtained fron the Pennsylvania Task Force.

35
Since there was concern that drilling through the tank foundation might cause

groundwater contamination by residual fuel oil from the spill, the soil

exploration methods had to be modified to prevent any infiltration of fuel oil

into the ground water. As a consequence, the borings through the tank

foundation itself were not carried to bedrock. However, two borings through

dikes in the vicinity of the tank foundation were carried into bedrock and

thus provide more complete information on the stratigraphy of the

unconsolidated deposits.

4.2.2 Foundation Description

Figure 4.1 shows a cross section of the tank foundation taken from the

construction plans provided by Ashland. The bottom plate of the tank was

supported by a crushed limestone pad varying in thickness from 1 ft at the

center to 1.5 ft at the tank perimeter. The top surface of the crushed stone

pad was to be graded to slope from the center of the tank to its perimeter,

with the center 0.5 ft higher than the perimeter. The crushed stone pad in

turn was supported by a compacted fill base which was constructed above the

natural grade to a height of approximately three feet.

A geotextile (reinforced plastic) blanket was inserted between the crushed

stone pad and the compacted fill base. The raised foundation had a diameter of

130 ft at the top and had sides which, according to the plans, sloped at a ratio of 3:1 toward the natural grade. The distance from the tank perimeter

to the upper edge of the slope was to be 5 ft.

4.2.3 Foundation and Site Conditions

Past Loading History

Prior to the reconstruction of Tank No. 1338, the site supported a tank 115 ft

in diameter and 40 ft high. The center of the foundation of this old tank was

approximately 27 ft from the center of Tank No. 1338. Thus a substantial

portion of the foundation of Tank No. 1338 was pre-loaded by the previous

tank. There is no evidence that the previous tank experienced any foundation

problems

.

36
Construction Records

Records are available from the construction of the compacted fill base. Solar

Testing Laboratories, Inc. of Pittsburgh issued six inspection reports between

July 2 and July 25, 1986 during the construction of the compacted fill. The

reports indicate that the soil was to be compacted to a Standard Proctor (ASTM

D 698) density of 95%. A soil sample taken from the borrow area prior to the

construction of the fill was described as a "reddish-brown sandy silt with

clay and rock fragments" with a maximum (ASTM D 698) dry density of 117.2 pcf

and an optimum moisture content of 12.4%. The inspection reports indicate

that the original grade prior to the placement of fill had some wet spots.

The muddy areas were to be skimmed and the site was to be proof-rolled prior

to placement of the fill. There is no evidence indicating that the inspection

laboratory confirmed that this was done.

The fill was placed in four 8-inch layers and nuclear density tests were

performed at several locations for each layer to determine dry density. Most

of the teste points met or exceeded the specified density. However, some test

points were somewhat deficient. Corrective actions were recommended by the

inspection laboratory. While the foundation contractor stated that corrective

actions were taken, there is no documented evidence that the deficient areas

were re-tested to confirm that the in-place density of the fill is adequate.

The available records, discussed above, indicate that the fill probably was

constructed in accordance with specifications.

Foundation Surface

The foundation surface was visually inspected by the NBS team after removal of

the bottom plate of the ruptured tank, and level readings were taken to

document its topography. The level readings are shown in Figure 4.2. These

can be compared with Figure 4.3 which shows part of an aerial photograph taken

by the Task Force after the collapse. It is important to recognize that, at

least in part, the deviations of the level readings from the planned

foundation profile shown in Figure 4.1 are attributable to disturbances that

occurred during the collapse.

37
It is apparent from the elevation readings that there was no recognizable

overall planar tilt of the tank foundation. There are, however, three areas,

marked in figure 4.2 and 4.3 as A, B and C, where the topography shows

localized depressions resembling the effects of local stability failures (edge

failures) . In all three of these areas bottom plate failures other than the

separation of the bottom plate from the shell of the tank occurred. The

aerial photograph shown in Figure 4.3 also shows some dark areas where oil was

ponding on the bottom plate. Visual inspection of the foundation pad after

removal of the bottom plate indicated some localized depressions at these

locations. These depressions could conceivably be indicative of localized

settlements which may have preceded the tank failure. In addition to the

depressions at points A, B, and C there is some variation in the elevations

around the tank perimeter. While these variations are not necessarily an

indication that non-planar shell settlements of a similar magnitude took place

prior to the collapse, records of level readings taken during partial filling

of the tank after its construction give an indication that there were some

non-planar shell settlements. These will be discussed later in this chapter.

Because no documents on the as-built foundation-pad profile are available, no

definitive conclusions can be drawn about possible overall residual

settlements from the difference between the level readings and the planned

foundation pad profile which specified a difference of 0.5 ft between the

elevation at the center of the tank and that at the rim. However, it

appears that the difference between the elevation at the center of the tank

and those at the rim is typically 0.3 ft, and thus may have decreased by

approximately 0.2 ft from that specified.

An examination of the area surrounding the elevated foundation pad gave no

indication of any heave such as that resulting from a rotational slide

opposite areas A, B, and C.

In areas A and B some soft, clayey silt was encountered near the surface of

the depressions. Test pits dug to explore the extent of the soft material

showed that soft material only occurred in small, localized pockets. It is,

therefore, reasoned that the pockets of soft material are attributable to

disturbance of the compacted fill and mixing with oil as a result of oil

38
spills. It was also observed that the surface of the fill under the geotextile

blanket was covered with a thin film of oil.

Compacted Fill

Data on the condition of the compacted fill are available from sixteen shallow

borings (M1-M16) , three compaction tests from bag samples (Dl, D2, D3) , and

three undisturbed (Shelby Tube) samples (SI, S2, S3) which were subsequently

tested in the laboratory (refer to location map in Figure 4.4). The bag

samples were combined and re -constituted in the laboratory to two dry

densities (111.9 pcf @ w = 16.8%; and 116.4 pcf @ w = 14.3%; where w = water

content) in order to perform triaxial compression tests for the range of

densities encountered in the compacted fill.

The shallow borings ranged in depth from 1.4 ft to 1.85 ft below the top

surface of the crushed stone pad. The borings were advanced by continuous

Standard Penetration Test (SPT) sampling (ASTM D 1586)

,

using a rope and

cathead assembly and a donut hammer. Blowcounts ranged from 14 to 30 blows

per ft and tended to be higher near the bottom of the borings . The water

content of the samples retrieved ranged from 12.9 to 18.1 percent. In the

laboratory the samples from the shallow borings were identified as "Brown

Sandy Lean Clay (CL)" with plasticity indices ranging from 10.4 to 13.7%.

Bag samples taken from the field density tests were identified as "Brown

Clayey Sand" (SC)

.

Compaction tests (ASTM D 698) performed on these samples

indicate a maximum dry density of approximately 117 pcf at an optimum moisture

content of 14%. This compares with a reported maximum dry density of 117,2

pcf at an optimum moisture content of 12.4% determined from a sample taken at

the borrow site prior to the construction of the fill.

Dry densities determined from the undisturbed samples (SI, S2, S3) ranged from

106 to 119 pcf and water contents from 14% to 22%. The data indicate that

five out of the six samples tested had dry densities in excess of the

specified dry density of 111.3 pcf.

39
Results of the two unconsolidated undrained triaxial compression test of the

remolded samples indicate the following strength values:

112 pcf density sample: c a = 1.3 ksf; <£ a = 14.5°

116 pcf density sample: c a = 2.0 ksf; 0 a = 26.0 s

where, c a = apparent cohesion

<f> a = apparent angle of friction for total stress envelope.

These samples were not saturated (83% to 86% saturation) and, therefore, the

tests do not provide a measure of the undrained shear strength in a saturated

condition. It is reasonable to assume that, at the time of the tank failure,

the compacted fill was also not saturated since it is elevated above the

surrounding area.

It is concluded from the above information that the compacted fill was

constructed in accordance with the specifications, and had a wet unit weight

of approximately 134 pcf and a shear strength under prevailing loading (refer

to 4.3.2) and drainage conditions at the time of collapse of approximately 2

ksf.

Subsurface Conditions

The site was explored by five deep borings (B1 through B5)

.

In addition, four

undisturbed samples (S4, S 5 , S5A, and S6) were taken at a shallow depth

immediately below the compacted fill. The boring and sampling locations are

shown in Figure 4.4.

The borings were advanced by continuous SPT sampling (ASTM D 1586)

.

The rope

and cathead method was used to drop a 140- lb donut hammer from a height of 30

inches to advance a 2 -inch O.D. 1 3/8 inch I.D. split spoon sampler without a plastic liner. The sampler was followed by a 3.5 inch I.D., 7,5 inch O.D.

hollow-stem auger which was used to support the sides of the drill hole. Rock

coring was performed with an NX double core barrel sampler which took 3 -inch

diameter rock cores. The recorded data include blow count for 6-inch depth

increments, soil type by visual-manual identification, sample recovery in

40
inches, pocket penetrometer readings on the split spoon samples, and Shelby

Tube and rock core recovery. The undisturbed and split- spoon samples were

preserved for laboratory testing.

To prevent groundwater contamination all boreholes were filled with cement

grout after completion of the boring. Special precautions were taken in the

area of the tank foundation (B3, B4, and B5) . For these borings a 16-inch

diameter hole was first drilled to a depth of 4 to 6 ft. Then a 12- inch O.D.

casing was placed in the drillhole and driven 6 inches into the bottom of the

16-inch hole. The space between the casing and the borehole was subsequently

filled with cement grout. Then a boring was advanced from the bottom of the

casing, using continuous sampling and a hollow-stem auger as described

previously . At the end of each working day, the drill rod and the hollowstem auger were withdrawn regardless of the depth to which the boring advanced

during the day and the borehole was filled with cement grout. As a result of

the time restraints imposed by this procedure, the borings located in the area

or immediate vicinity of the tank foundation had to be terminated in soft

material and the unconsolidated deposits were not explored to their full

depth

.

Laboratory tests included soil classification in accordance with ASTM D 2487,

and oedometer and unconsolidated undrained triaxial compression tests of

undisturbed samples. Most of the undisturbed samples were not saturated, and

thus the undrained-unconsolidated tests did not provide definitive information

on the undrained shear strength. However, an attempt was made to saturate

samples from borings S4 and S5 by applying a back pressure in excess of 11 ksf

before performing unconsolidated undrained triaxial tests. A set of

consolidated undrained tests was also performed on an undisturbed sample taken

from boring S6

.

In order to assess the subsurface conditions to the full depth of the

unconsolidated deposits it is necessary to interpolate between borings B1 and

B2 which were carried to bedrock. Boring B1 is 200 ft from the centerline of

the tank in a Southeasterly direction and was carried 7 ft into bedrock to

elevation 719. Boring B2 is approximately 215 ft from the center of the tank

in a Northwesterly direction and was carried 10 ft into bedrock to elvation

41
726. Between these two borings are borings B3 , B4 , and B5 . Boring B5 is in

the center of the tank and extends to a depth of approximately 24 ft below the

compacted fill. Boring B3 is on the tank perimeter in a Northwesterly

direction from the tank center and extends to a depth of approximately 18 ft

below the compacted fill. Boring B4 is approximately 85 ft from the

centerline of the tank in a Southeasterly direction, it is approximately 25 ft

outside the perimeter of the tank, and it extends to a depth of 36 ft into the

natural ground. While not extending to bedrock, boring B4 traverses all the

compressible layers and extends approximately 6 ft into a layer of dense silty

sands and sandy silts.

Comparison between adjacent borings indicates that the stratification of the

compressible layers is approximately horizontal. Figure 4.5 shows an

idealized schematic soil profile based on the available information. The

upper soil layer extending approximately to elevation 757 is identified as a stiff moist silty clay. Several undisturbed samples were taken from this

layer, providing information on compressibility and undrained shear strength.

The undrained shear strength (su) given in figure 4.5 is based on the lowest

strength recorded for a sample from boring S4 which was saturated in the

laboratory (this is a lower limit, because it is unlikely that the soil was

saturated at the time of collapse) . This upper layer is underlain by sandy

silty clays and sandy clayey silts to approximate elevation 745. Part of this

second layer, below approximate elevation 751, is saturated. Below this

second layer is a layer of saturated loose silty sands and sandy silts. Below

approximate elevation 740 is a layer of dense clayey silty sands with rock

fragments which extends to bedrock. It can be seen from the recorded

plasticity indices (P.I.) that the boundaries between the layers above the

dense clayey silty sands are artificial and predicated on the soil

classification method used. Actually, the deposits become gradually more

coarse-grained with depth, their clay fraction gradually decreases with depth,

and there is a gradual transition from clays to silts to silty sands.

42
4.3 Foundation Analysis

4.3.1 Failure Limit States

On the basis of past performance of oil tanks with similar foundations,

several possible foundation- induced failure limit states have been identified.

These are: stability failures which include base shear failures (bearing

capacity failure or planar tilt) and edge shear failures (local stability

failure); and settlement- induced failures which include bottom plate or bottom

plate connection failure induced by excessive differential settlements and

excessive stresses induced in the shell by non-planar shell settlements along

the tank perimeter (shell distortion) . In addition to the possibility that

the overall tank failure may have been triggered by foundation instability or

excessive settlements, NBS investigated the possibility that settlementinduced shell distortions could have contributed to the tank failure by

increasing the circumferential tensile stress in the shell at the location

where the brittle fracture originated. It is important to consider the fact

that prior to the erection of Tank No. 1338, the site supported a somewhat

smaller tank which exerted foundation loads of approximately 80 percent of the

load acting at the time of collapse. No foundation failures were reported for

this old tank.

4.3.2 Stability

Duncan and D'Orazio (1984) [4.1] studied available documented information on

steel tank foundation shear failures and developed empirically-based criteria

for foundation stability. The safety factors against base and edge shear

failure calculated herein are based on these criteria.

Using an oil depth in the tank of 46 ft and a specific gravity of the oil of

0.87, the oil pressure on the bottom of the tank at the time of failure was

approximately 2.5 ksf. If it is assumed that the bottom plate was 1/4 inch

thick, the roof plate was 1/8 inch thick, the average thickness of the shell

was 1/2 inch, and an allowance is made for the columns and beams supporting

the roof, the total weight of the tank is calculated to be approximately 630

kip. If, for the purpose of base stability calculations, it is assumed that

43
the tank weight is evenly distributed over the base, it would contribute

approximately 0.06 ksf to the pressure at the base of the tank, for a total

base pressure of 2.56 ksf. If it is assumed that the crushed stone pad had a unit weight of 140 pcf and the compacted fill had a unit weight of 134 pcf,

the pressure at the base of the fill was approximately 3 ksf, spread over the

equivalent diameter ,De = 124 ft.

It is assumed that a shear failure would not extend below the surface of the

dense silty sand at elevation 740. Thus the thickness of the soft material,

T, is approximately 28 ft. Therefore,

De/T = 124/28 - 4.43 and Nc = 6.1 (Reference [4.1])

where, Nc = dimensionless bearing-capacity factor for base shear.

No triaxial test information for the U2 undisturbed samples, which were below

the groundwater table at elevation 751, is available. It is assumed that the

undrained shear strength of these saturated deposits is lower than that of the

partially saturated deposits above them after their saturation in the

laboratory. Using an estimated average undrained shear strength of 1 ksf for

the entire depth of the soft layer, the factor of safety against base shear,

Fb , can be calculated:

Fb = (1) (6 . l)/3 = 2.03

This estimate is conservative, because it is based on the lowest undrained

shear strength obtained for a saturated soil condition. At the time the

borings were taken the soil was only partially saturated above elevation 751,

and it is unlikely that at the time of the collapse the upper soil layer was

saturated. The estimated minimum factor of safety against base failure in

excess of 2 corroborates the findings from the level survey of the foundation

surface which indicate the absence of any recognizable planar tilt.

In order to calculate the factor of safety against edge shear failure it is

necessary to make an assumption with regard to the size of the area that

failed, or else find the area with the lowest factor of safety. A failure

44
associated with areas A, B, and C outlined in Figure 4.2 would involve a segment of the tank bottom of approximately 20 ft width in a radial direction

(refer to foundation surface topography, figure 4.2). The possibility of a slope failure involving an even .smaller bottom plate segment (essentially a failure of the compacted fill slope) is considered unlikely because the shell

would bridge over such a small perimeter segment and thus transfer the load in

case of a foundation movement.

The following expression suggested by Duncan, et al [4.1] for bearing-capacity

factors for edge failures was .derived from values suggested by Bjerrum and

Overland [4.2]:

Ne = 5.2 + B/D

where, Ne = dimensionless bearing capacity factor for edge shear

B = radial width of failed bottom segment.

D = diameter of tank

Safety factors against edge shear failure were calculated in accordance with

methods recommended in Reference 4.1, assuming an undrained shear strength of

2 ksf for the compacted fill, an angle of internal friction of 40° for the

crushed stone pad, and an undrained shear strength of 1.2 ksf for the material

underlying the compacted fill. As recommended in Reference 4.1 the shear

resistance of the pad and compacted fill was considered in the calculations,

and a bearing pressure of 2.8 ksf at the base of the crushed stone pad was

assumed. For these assumptions it was found that the factor of safety against

edge failure is 2.2 or larger.

4.2.3 Settlements

Differential Settlements

Steel tanks can tolerate large differential settlements when compared to other

structures. However, there are records of tanks which ruptured as a result of

excessive differential settlements. D'Orazio and Duncan (1987) [4.3]

summarize available records on settlement characteristics and tolerances of

45
steel tanks available before December 1986. On the basis of these records,

methods for predicting settlement profiles and tolerances are recommended.

The paper addresses differential settlements in the radial direction only,

which are associated with failure of the bottom plate, or its connection to

the shell.

Bottom-plate failures could also include failure of the fillet welds

connecting individual plates similar to that observed between points A and B

in figure 4.3.

The shape of the settlement profile of steel tanks depends on the depth and

stiffness of the underlying soft soil layer, and can be predicted on the basis

of two parameters: the D/T Ratio, which in this case is 4.43; and the factor

of safety against stability failure (base or edge) which in this case is

larger than 2. In accordance with Reference 4.3, the shape of the predicted

settlement profile falls within the limits shown in Figure 4.6 (profile-shape

B in Reference 4.3). The effect of subsurface characteristics on profile

shapes suggested in Reference 4.3, which is based on empirical data, was also

corroborated by some finite-element studies by the authors of ref. [4.3]. For

the profile-shape limits shown in Figure 4.6 the following differential

settlement tolerance is recommended in reference 4.3 to prevent bottom plate

rupture

(Pc-Pe) /D = 0.015

where, p c = settlement under center of tank

p e = settlement under edge of tank.

The settlement itself has two components: a settlement caused by the

distortion in the foundation soil due to shear strains and a consolidation

settlement. The settlement caused by the distortion in the foundation soil

occurs immediately since the shear distortion is not associated with volume

changes and thus requires no pore water displacement. The consolidation

settlement is time dependent because it requires volume change in the subsoil

and thus displacement of pore water. In the case of saturated clay, only very

small consolidation settlements could occur during the time frame of the

46
loading of the tank prior to failure. For the conditions at the site,

however, part of the consolidation settlement could have occurred, because: 1)

the silty clays and silts above elevation 751 were not fully saturated (about

83 to 85% saturation) and volume change could occur by the compression of the

air space in the pores; and 2) soil permeability is increasing with depth, and

the clayey silts and silty sands below elevation 750 could have experienced

some instantaneous volume changes. An upper limit of the consolidation

settlement is therefore calculated and compared with the failure criterion.

The instantaneous shear- strain induced settlement can be calculated by the

following equation ([4.3]):

Pi = (I

pqDe )/2E

where: = instantaneous settlement

q = foundation pressure = 3 ksf

D e = diameter of loaded area = 124 ft

E = average value for Young's modulus of the soil

Ip = dimensionless influence factor, depending on D e /T and r/R (radial

distance from center divided by radius)

.

A lower limit for Young's modulus is estimated to be 640 s u = 770 ksi [4.3].

Influence factor Ip is estimated to be 0.15 in the center of the tank and 0.08

at the perimeter. Thus the instantaneous shear-strain induced settlement is

of the order of 0.04 ft. in the center and 0.02 ft on the perimeter of the

tank.

An upper limit for the consolidation settlement calculated from the oedometer

test results (assuming that all the pore water was displaced prior to the

failure) is 0.3 ft at the center of the tank and 0.16 ft at the tank perimeter

(actual settlements conservatively estimated to have occurred within the time

frame of the failure are on the order of 0.18 ft in the center and 0.1 ft at

the perimeter). Thus using the upper limit:

(Pc'Pe)/D “ 0.0013 « 0.015

47
Thus the differential settlement in the radial direction at the time of the

tank failure is estimated to have been more than an order of magnitude smaller

than the differential settlement that could have caused a rupture in the

bottom plates of the tank or the fillet weld connecting the bottom plates.

Non-Planar Shell Settlements

Unlike differential settlements in the radial direction, non-planar shell

settlements cannot be predicted on the basis of the stiffness and depth of

compressible soil layers, even though some estimates could conceivably be

derived from changes in soil characteristics between adjacent borings. For

this reason information is usually obtained by monitoring shell elevations

around the tank perimeter when the tank is first filled with water to test its

integrity

.

There are three possible reasons for non-planar shell settlements: 1)

differential foundation settlements during the filling of the tank or after

the tank is filled; 2) irregularities in the foundation pad; and 3)

irregularities in the tank geometry. Differential foundation settlements tend

to increase gradually while the tank is filled and may increase with time

after the tank is filled. Irregularities in the foundation pad and tank

geometry, on the other hand, could cause initial seating displacements which

do not increase when additional fluid is added.

Two questions need to be addressed in this analysis: 1) Could shell

distortions have caused a structural failure of the shell? 2) Could shell

distortions have contributed to the shell fracture by increasing the

circumferential tensile stress at the flaw where failure originated?

Records of shell displacements are available for an initial filling with 5 ft

of water on June 11, 1987, corresponding to a foundation live load of 0.312

ksf, and a subsequent filling with 10 ft of oil on August 25, 1987

corresponding to a foundation live load of 0.542 ksf. The foundation live load

at the time of the failure was 2.5 ksf or 4.6 times the maximum load for which

data on non-planar shell settlements are available.

48
Table 4.1 shows level readings, recorded to the nearest 0.01 ft, which were

taken at 12 points on the perimeter of the tank, representing twelve 30° tank

segments with a distance of 31.59 ft along the arc between measuring points.

Another reading was taken between points 11 and 13. The reading is not shown

in the table because there was no shell distortion between points 11 and 13.

The readings taken on 3/12 are for the empty tank, those taken on 6/11 are for

a 5-ft water level, and those taken on 8/25 for a 10-ft oil level. Column 5

shows elevation changes for the water filling and column 6 for the oil

filling. Column 7 shows the estimated out-of -plane settlements for the oil

filling. The estimated out-of -plane settlements were calculated asstiming a planar settlement of 0.02 ft and no planar tilt. Hypothetical out-of -plane

settlements at the time of the failure are listed in column 8. These

settlements were calculated by linear extrapolation of the out-of-plane

settlements calculated from the data for the 10-ft oil filling. They are

thought to represent an estimated upper bound, because the time-element

required for consolidation settlements was not taken into consideration. It

is also important to realize that large errors in the estimate of these

hypothetical settlements may exist because the precision of the recorded

settlements is only 0.01 ft and initial seating settlements are not

considered. Thus the out-of-plane settlements in column 8 are at best a very

rough estimate.

There are criteria for non-planar shell settlements of floating-roof tanks

[4.4, 4.5, 4.6]. These settlements cause out-of-plane deformations in the

shell which in turn distort the geometry at the top of the tank, causing

problems with the floating roof. The criteria are intended to keep the

deformations of the top of the tank within acceptable tolerances and are not

related to stresses induced in the shell. A criterion related to shell

stresses, which would be applicable to conical-roof tanks, was proposed by

Marr et al.(1982) [4.7]. The latter criterion was derived from "beam theory"

and thus assumes that no horizontal deformations occur in the shell.

Therefore the shell stiffness, and consequently the induced circumferential

stresses are overestimated; as a result the criterion is overly conservative

with respect to circumferential stresses. The criterion can, however, be

considered a lower bound for estimating the magnitude of potentially harmful

distortions with respect to induced circumferential stresses. The criterion

49
ignores potential effects from induced vertical stresses resulting from

horizontal bending of the shell out of its plane. There are no documented

cases where shell distortion caused structural failure of the shell. This may

indicate that vertical shell distortions, while causing out-of-plane

deformations, do not tend to overstress the shell. Reference 4.7 notes that

three tanks which experienced shell distortions greater than those permitted

in the proposed criterion did not exhibit shell damage. The following

criterion for maximum tolerable out-of-plane settlements has been proposed in

4.7:

AS! < ll(<7

y

)(i 2 )/[(E>(H)]

where

,

Asi = Si - 0.5(s^_^ + s i+l) = distortion of tank wall at point i

s-l = out-of-plane settlement at point i

Oy - yield stress of steel

2 = length of shell between observation points i and i + 1

E = Young's Modulus of steel

H = height of tank

The criterion is to prevent yielding of the shell by circumferential stresses.

For a ductile steel, actual rupture would occur at considerably larger

deformations. Assuming a steel yield stress of 34 ksi, a length between gage

points of 31.59 ft and a tank height of 48 ft, the allowable limiting value of

As^ according to this criterion can be calculated as, :

As£ = 11(34) (31 . 59 2 )/[ (29000) (48) ] = 0.268 ft.

The highest value of As^ for the 10 ft oil level occurs at point 7 and is

calculated at:

As t = 0.01 - 0.5(0-0.01) = 0.015 ft

if this value is linearly extrapolated to the 46 ft oil level at the time of

failure, it results in a As^ value of 0.069 which is only approximately 1/4 of

the lower bound for settlement distortions, which could induce circumferential

yield stresses in the tank shell.

50
Thus it is unlikely that non planar she'll settlements could have induced

critical circumferential stresses in the tank shell.

Possible effects of the shell distortion on circumferential stresses at the

location of the flaw are approximately assessed in Section 5.2 by a finite

element study of the effect of an induced vertical sinusoidal shell

deformation at the foundation level. It is concluded that the out-of -plane

shell distortions did not significantly affect the magnitude of the

circumferential stress at the locetion of the flaw.

4 . 3 Summary

An analysis of available data on subsurface conditions, foundation loads at

the time of failure, and foundation displacements indicates that it is

unlikely that the base of the tank ruptured as a result of instability or

excessive differential settlements. Based on an analysis of the observed

effects of a 10 ft oil level to estimate shell distortions at the time of the

failure, it is concluded that shell distortions caused by foundation

displacements did not contribute to the tank failure.

51
Table 4.1’ - Shell Displacements

(all dimensions in ft)

Point/Date 3/12 6/11 8/25 PI P2 s 2 Sf

1 96.81 96.79 96.70 .02 .02 0 0

2 96.89 96.87 96.87 .02 .02 0 0

3 96.95 96.93 96.93 .02 .02 0 0

4 96.94 96.93 96.92 .02 ' .02 0 0

5 96.95 96.93 96.92 .02 .03 -.01 - .046

6 96.98 96.96 96.95 .02 .03 -.01 - .046

7 96.91 96.91 96.90 .00 .01 + .01 + . 046

8 96.95 96.94 96.93 .01 .02 0 0

9 97.04 97.03 97.02 .01 .02 0 0

10 97.09 97.09 97.08 0 .01 + .01 + . 046

11 96.98 96.98 96.96 0 .02 0 0

13 96.91 96.91 96.90 0 .01 + .01 + .046

PI = total settlement on 6/11/87

p 2 = total settlement on 8/25/87

S 2 = out-of-plane settlement on 8/25/87

Sf = hypothetical out-of-plane

settlement at failure

2

52
53

Figure

4.1

-

Cross

Section

of

Tank

Foundation
N

2.2

2.2

2.4

Elevations above El. 770, in ft.

Figure 4.2 - Topography of

the Base Plate

Foundation Surface

. (Provided by the

Surveyed After Removal of

Pennsylvania Task Force)

54
Figure 4.3 Aerial Photograph of Tank Foundation Taken Prior to Removal of

the Base Plate. (Provided by the Pennsylvania Task Force)

55
o4—* CM

05

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03a

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«

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03

03 a

o

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in

0303C

o

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o

56

Figure

4.4

-

Location

Map

for

Soil

Exploration.

(From

information

provided

by

the

Pennsylvania

Task

Force)
CNJ

Ol 0 ^ CM

Q.

CD

CM7O

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03

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in>N

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57

Figure

4.5-

Schematic

Drawing

of

Soil

Profile

Developed

by

Interpolation

Between

Borings.

(Based

on

subsurface

data

supplied

by

the

Pennsylvania

Task

Force)
P/Pc

Figure 4.6

0 0.2 0.4 0.6 0.8 1.0

r/R

r=radial distance

R=radius of tank

p=settlement

pc =settlement in center of tank

Assumed Shape of Normalized Settlement Profile of Tank.

(Profile-shape B from D'Orazio et al., 1987)

58
5.0 STRUCTURAL ANALYSIS

5 . 1 Stress Analysis of the Tank

A structural analysis was performed to determine the stresses in the tank at

the time of collapse. The cylindrical wall of the tank was modeled (see

Figure 5.1) and analyzed by the finite element method for the loading produced

by the static head of the stored product. Pre- and post-processing of the

data were done using PATRAN [5.1]. An axisymmetric analysis was performed

using 4-node quadrilateral elements [5.2]. The following assumptions were

made in modeling the tank:

o The plate thicknesses were assumed to be the nominal plate thicknesses

(shown in Figure 2.1) less an average corrosion of 1/16 in.

o The shell courses were assumed to be 8 ft - 0 in high.

o The welds between successive shell courses were modeled as smooth

transitions as shown in Figure 5.1.

o The tank bottom and roof were not included in the model.

o The base of the tank was modeled as fixed in the radial direction but free

to rotate.

o The top of the tank was modeled as free in the radial direction and free to

rotate

.

o The height of the stored product was assumed to be 46 ft - 0 in.

o The specific gravity of the stored product was assumed to be 0.87 for a density of 54.3 pcf.

Two elements were used through the shell thickness and the aspect ratio of the

elements was no greater than 1:1.5. This produced a model of the full height

of the tank consisting of 5913 nodes and 3940 elements.

59
The circumferential membrane stress (the normal stress in the circumferential

direction reported at the tank wall centerline) is plotted in Figure 5.2. The

circumferential membrane stress is zero at the tank base (the base is radially

restrained) and reaches a maximum of 15.1 ksi in the first shell course at a height of approximately 40 in. The circumferential membrane stress at the

juncture of the first and second shell courses is approximately 13.7 ksi.

The effect of a rotational restraint at the base of the tank was also

investigated. The analysis described above was repeated except that the base

was restrained against both translation and rotation- The meridional bending

stress (vertical stress at the outer surface) in the first two shell courses

is shown in Figure 5.3. The rotational restraint produces a significant

meridional bending stress at the base of the tank but the effects of the

restraint diminish rapidly with increasing height. At the juncture of the

first and second shell courses (8 ft) , there is essentially no difference in

the meridional bending stress between the rotationally restrained and

rotationally free cases. In addition, the meridional bending stress at this

elevation, resulting from the discontinuity in plate thickness, is shown to be

approximately zero. Thus, the stress at the juncture of the first and second

shell courses is essentially the circumferential membrane stress and is

independent of the base fixity condition.

5 . 2 Foundation Settlement Analysis

To evaluate the stresses in the tank resulting from differential foundation

settlement, a finite element analysis was made for an assumed displacement

profile. The following assumptions were made in modeling the tank:

o The plate thicknesses were assumed to be the nominal plate thicknesses

(shown in Figure 2.1) less an average corrosion of 1/16 in.

o The shell courses were assumed to be 8 ft - 0 in high.

o The welds between successive shell courses were not modeled.

60
o The tank bottom and roof were not included in the model.

One half of the tank wall was modeled using 8 -node quadrilateral shell

elements. Each element was uniform in size and measured 24 in high by 23.52

in wide. This resulted in 24 elements along the height of the tank and 96

around one half of the circumference for a total of 2304 elements. A diagram

of the finite element model is shown in Figure 5.4.

Symmetry was accounted for in the boundary conditions along the tank

centerline. The top of the tank wall was restrained in the radial direction

only to account for the restraint provided by the roof. The base of the tank

was restrained radially and subjected to an assumed displacement profile in

the vertical direction, vq(0), given by

This displacement profile is symmetric about the tank centerline, has a maximum displacement of 0.12 in (0.01 ft), and has a zero slope at 0° (tank

centerline) and 30° (extent of localized differential settlement) . This

profile was selected to represent a differential displacement of 0.01 ft in a 60° arc corresponding to level readings taken at 12 locations around the tank

base

.

The maximum meridional membrane stress was approximately 3 ksi. Of more

interest is the circumferential stress at the juncture of the first and second

shell courses (height above tank bottom of the flaw) . The membrane

circumferential stress, plotted as a function of d , is shown in Figure 5.5.

As can be seen, the maximum tensile stress is approximately 120 psi and the

maximum compressive stress is approximately 200 psi. The stresses are seen to

decay to zero within 90°. From this analysis, it is seen that locallized

differential settlements on the order of 0.04 ft (approximately 0.5 in) would

produce tensile membrane stresses in the vicinity of the flaw of only 500 psi.

Thus, it is unlikely that any differential foundation settlements contributed

significantly to the tensile stress normal to the flaw.

vq(0) = 0.06 cos(60) + 0.06

vq(0) = 0-0

for 0° < 9 < 30°

for 30° < 9 < 180°

61
Figure 5.1 - Axi symmetric Analysis Model of Tank Wall

62
TANK

HEIGHT

(ft)

CIRCUMFERENTIAL STRESS VS. TANK HEIGHT

Figure 5.2 - Circumferential Membrane Stress in the Tank Wall

63
MERIDIONAL STRESS VS. TANK HEIGHT

Figure 5.3 - Meridional Bending Stress in the First Two Shell Courses

64
65
CIRCUMFERENTIAL

MEMBRANE

STRESS

(psi)

6 (degrees)

Figure 5,5 - Circumferential Stresses in the Tank Wall at a Height of 8 ft

due to an Assumed Differential Settlement of 0.01 ft

66
6.0 THERMAL ANALYSIS OF THE TANK WALL

6 . 1 General

The fracture toughness of carbon steel, such as that used in this tank, is

dependent on temperature. Therefore, to determine the fracture toughness of

the steel at the time of collapse, it is necessary to estimate the temperature

of the tank wall in the vicinity of the flaw. This information will be used

in the fracture analysis in Chapter 8.

6 . 2 Ambient Temperature and Wind Data

Surface weather observations for January 2, 1988, the day of the collapse, are

presented in Table 6.1. These data were recorded at the Allegheny County

Airport (see map in Figure 1.2) located approximately seven miles from the

Ashland Floreffe terminal. At the approximate time of collapse (16:57), the

recorded temperature was 26° F. As indicated in Table 6.1, the temperature

was 26° F for the preceeding two readings as well. Due to the proximity of

the Ashland terminal to the Allegheny Airport, it was assumed that the ambient

air temperature at the time of collapse was 26° F.

The wind speed recorded at the Allegheny County Airport at the approximate

time of collapse was out of the south-west at 8 knots or slightly more than 9

mph. To translate this information to the tank collapse site, the difference

in surface terrain between that at the airport where the data were recorded

and that at the collapse site must be considered. The variation in wind speed

with terrain roughness is well documented [6.5]; at equal elevations, the mean

wind speed will be less over the rougher terrain. Additionally, there is a variation in wind speed with height above the ground. The relationship between

mean wind speed, height above ground and roughness of the terrain is [6.5],

u(z)

u(z

r

)

in

‘0 z in —Z

0

67
where

,

u = mean wind speed

z = height above ground

z r = reference height

z o = roughness length

Roughness lengths for various terrain categories have been compiled [6.5]. If

the airport is categorized as "open” terrain ( zq = .230 ft) and the Ashland

terminal as "sparsely built up" terrain (zq = .989 ft), then for a wind speed

of 9 mph recorded at height of 20 ft (assumed) , the mean wind speed at a height above ground of 8 ft (top of first shell course) is calculated to be

between 4 and 5 mph.

6 . 3 Temperature of the Oil in the Tank

The temperature of the oil was recorded by Ashland personnel at the time the

product level readings were taken. A graph showing the temperature of the

fuel oil in Tank No. 1338 for the period August 25, 1987 through January 1,

1988 is presented in Figure 6.1. The temperature of the oil in Tank No. 1338

on the day before the collapse (the last date on which the temperature was

recorded) was 46° F. The temperature of the oil pumped into Tank No. 1338 on

the day of collapse was reported to be 46° F as well. Assuming that the

recorded temperature of the oil, both in the tank and on the barges, was

representative of the stored product (i.e., no stratification had occurred),

and that complete mixing occurred as a result of the introduction of the fuel

oil into Tank No. 1338, a reasonable estimate of the oil temperature at the

time of collapse is 46° F.

6 . 4 Thermal Analysis

The steady- state heat transfer through the wall of the storage tank may be

expressed as

68
- oil - A

^ 1 Ax 1 h.A

+

kA

+

h A

l o where

Toil = temperature of the oil within the storage tank, °C = surrounding ambient temperature, °C

9

hj_ = inside surface heat transfer coefficient, W/m^*°C

hQ = outside surface heat transfer coefficient, W/m^^C

k = thermal conductivity of the storage tank wall, W/m°C

Ax = thickness of the* storage tank wall, m

A = heat transfer area, nrIn accordance with Reference 6.1, the inside surface heat transfer

coefficient, h^, due to free convection may be evaluated using,

Nu = .59 I Gr _ Pr

J

,25

y.f L y.f fJ

where

Nuy,f = local Nusselt number evaluated at the interior film

temperature, (T^ + T0 i ) /2 where Tj_ is the inside surface

temperature, and at the juncture of the first and second shell

courses (dimensionless)

,

Gry ( f = local Grashof number evaluated at the interior film

temperature and at the juncture of the first and second shell

courses (dimensionless) , and

Prf = Prandtl number evaluated at the interior film temperature

(dimensionless)

.

The inside surface heat transfer coefficient is computed from the relationship

where

h. =

i

Nu _ k_

y. £ f

kf = thermal conductivity of the oil evaluated at the interior film

temperature, W/m°C, and

y = height of the first shell course, m. 69
The required thermal properties were taken from Spiers [6.2]. The outside

heat transfer coefficient is computed using the Jurges correlation presented

in McAdams [6.3] ,

hQ = 5.7 + 3.8V

where

V = wind speed in m/s

.

The thermal conductivity of the steel wall is taken as 54 W/m*°C which

corresponds to a 0.5% carbon steel [6.4], The inside and outside surface

temperatures are computed as

and

T. = T .,

- i oil h.

i

T

o = T, h

Because the properties used to compute the inside surface heat transfer

coefficient are evaluated at the interior film temperature which in turn is

dependent on the inside surface temperature, the solution requires an

iterative approach. An initial guess is made to approximate the inside

surface temperature. The inside heat transfer coefficient, heat transfer, and

accompanying inside and outside surface temperatures are then computed. The

interior film temperature is computed from the resulting inside surface

temperature. The computations are repeated using the computed inside surface

temperature. This procedure is repeated until the interior film temperature

converges to within 0.01°C.

The heat transferred through the tank wall was computed using the values for

the temperature of the oil in the tank, ambient air temperature, and wind

speed determined in the previous sections and converted to the appropriate

units. It was found that the temperature gradient through the tank wall at a height of 8 ft was almost negligible, being less than one half a degree.

Therefore, the average of the outside and inside wall temperatures is

reported. For an oil temperature of 46° F, an outside ambient temperature of

26° F and a wind speed of 5 mph, the average wall temperature of the tank, at

70
a height of 8 ft, was approximately 32° F.

To illustrate the effects of the oil temperature and wind speed on the

temperature of the tank wall, heat transfer calculations were made for oil

temperatures ranging from 35 to 55° F and wind speeds from 0 to 10 mph.

Results are given in Table 6.2.

71
Table 6.1 - Surface Weather Observations - January 2, 1988

(Allegheny Co. Airport)

Time Temperature

(° F)

Wind Speed

(knots)

Wind Direction

(deg from North)

00:50 18 7 290

01:50 16 7 290

02:50 16 10 290

03:50 14 7 270

04:50 13 6 270

05:50 13 5 270

06:50 12 4 300

07:45 12 6 300

08:45 14 4 270

09:50 15 6 180

10:50 17 9 170

11:50 20 5 230

12:50 22 0 000

13:50 24 6 230

14:50 26 6 240

15:50 26 8 220

16:57 26 8 230

17:51 24 5 210

18:50 24 6 230

19:55 23 5 190

20:55 22 6 200

21:55 21 6 200

22:55 21 6 200

23:55 20 5 180

72
Table 6.2 -Average Tank Wall Temperature at Elevation 8 ft

Wind Speed

(mph)

Oil Temperature (° F)

35° 0 O 45° 50° 55°

0 30.2 33.0 36.0 39.3 43.0

1 29.7 32.2 34.9 37.9 41.2

2 29.3 31.5 34.0 36.7 39.8

3 29.0 31.0 33.3 35.8 38.6

4 28.7 30.6 32.7 35.0 37.6

5 28.5 30.2 32.2 34.3 36.8

6 28.3 29.9 31.7 33.8 36.0

7 28.1 29.7 31.4 33.2 35.4

8 28.0 29.4 31.0 32.8 34.8

9 27.9 29.2 30.7 32.4 34.3

10 27.8 29.0 30.5 32.1 33.9

73
TEMPERATURE

OF

OIL

IN

TANK

CO

(do) 110 dO 3dmva3dlAI31

74

Figure

6.1

Temperature

of

the

Oil

in

Tank

No.

1338
7.0 METALLURGICAL INVESTIGATION

7 . 1 Introduction

The NBS metallurgical investigation consisted of examination of the failed

tank at the field site of the failure, selection of samples from the tank to

be used for characterizing material properties, identification of possible

failure modes and causes, detailed examination of the fracture surfaces, and

failure analysis based on fracture mechanics principles. In addition, a detailed fractographic analysis of the fracture surface and the defect found

on the fracture surface was conducted by Battelle. The results were made

available to NBS and are included in this report.

As described in detail in Chapter 2 of this report, the tank was constructed

of steel plates that were nominally 8 ft wide and 32 ft long. The tank

consisted of 6 shell courses each approximately 8 ft high. The bottom, or

first, shell course was nominally 7/8 in thick, the next higher or second

course was 1/8 in thinner i.e. 3/4 in thick, and each subsequent shell course

was nominally 1/8 in thinner than the next lower shell course. The highest or

sixth course was 1/4 in thick. A primary fracture was found that had

propagated vertically through all six shell courses and split the tank from

top to bottom as shown in Figure 7.1. Visual observation at the field site

showed that this fracture was flat and perpendicular to the plate surface over

its entire length (see Figure 3.10) with no evidence of tearing or shear

fracture. Chevron markings on the fracture surface pointed to a defect on the

fracture surface that was identified as the point of initiation of the

fracture. This defect was near the top of the first shell course just below

the horizontal weld between the first and second shell courses and just below

the vertical weld between two plates in the second shell course.

As shown in Figure 7.1, test specimens were taken from the first and second

shell courses to evaluate the typical properties of the steel plates in the

tank. Test specimens were taken from typical regions of vertical, horizontal,

and T weld sections to evaluate the welds in the tank. The laboratory tests

conducted to evaluate the properties of the tank steel were tensile tests to

determine strength and ductility, various fracture toughness and notch

75
sensitivity tests to determine the susceptibility to fracture initiation,

propagation, and arrest, and metallography and hardness tests.

7 . 2 Materials Properties Characterization

7.2.1 Location of Test Specimens

Samples were taken by Ashland from the failed tank to evaluate the properties

of the steel from which the tank was constructed. Test samples approximately

2 ft by 3 ft size were taken for the NBS from the steel plate in the first and

second shell courses of the tank near the vertical fracture surface as shown

in Figure 7.1. Test sample 5E was a base metal sample from the first shell

course of the tank approximately 4 ft from the vertical fracture. Test sample

3E was a base metal sample from the second shell course of the tank

approximately 4 ft from the vertical fracture. Test sample 4E included the

horizontal weld between the first and second shell courses of the tank. Test

sample 9 was a T weld at the junction of two plates in the second shell course

and a plate in the first shell course of the tank approximately 60 ft from the

vertical fracture. The identity and location of specific material properties

test specimens taken from each of the test samples are shown in Figures 7.2,

7.3, 7.4 and 7.5. For the plates in the first and second course of the tank,

the rolling direction was parallel to the long (32 ft.) dimension of the

plate. The orientation of the plate (relative to the rolling direction) was

maintained and taken into account for all specimens used for mechanical

properties tests and for fracture toughness tests.

7.2.2 Chemical Composition

To identify the plate steel used in the tank, samples for chemical analysis

were taken from the first and second shell course plates. The chemical

analysis was performed by emission spectroscopy following ASTM Standard E-415-

85 [7.1]. Quantitative analysis was performed to determine the amount of C,

P, S, Mn, Si, Ni, Cr, V, Ti, Mo, Cu, Co, and A1 in each sample.

Chemical analyses were performed on three locations in the base metal specimen

from the first course (sample 5-3). As shown in the micrograph in Figure 7.6,

76
the first shell course base metal showed substantial evidence of "banding" due

to chemical segregation in the steel plate. Therefore, chemical analyses were

taken in three locations as shown in Figure 7.6 to determine the extent of the

chemical segregation. The chemical analyses were performed four times at each

location. The results (average of four determinations) are shown in Table 7.1 as samples 5-3-1, 5-3-2, and 5-3-3 corresponding to the locations shown in the

micrograph in Figure 7.6. Chemical analyses were performed on one location in

the base metal specimen from the second shell course (sample 3-2)

.

The sample

was analyzed four times. The results of the chemical analyses are shown in

Table 7.1 as sample 3-2.

The tank was probably constructed of steel made to the ASTM standard A- 10- 39

for Mild Steel Plates [7.2]. This ASTM Standard has been replaced by Standard

A-283-84 [7.3] for Low and Intermediate Tensile Strength Carbon Steel Plates,

Shapes, and Bars. The very limited specifications for chemical composition of

the steel are identical in both Standards A- 10- 39 and in A-283-84. The only

chemical composition limits specified in either standard are that the maximum

phosphorus content should be less than 0.040 percent, the maximum sulfur

content should be less than 0.050 percent, and. if copper is specified, the

minimum copper content should be greater than 0.20 percent. No specific

carbon content or amount of any alloying element is specified in either of the

referenced standards.

Table 7.1, shows that the steel plates from both the first shell course

(sample 5-3) and from the second shell course (sample 3-2) satisfy the

chemical composition requirements of both the ASTM A-10-39 and A-283-84

Standards. Steel made to the ASTM Standard A-283-84 (Grade C strength level)

is one of the types of steels permitted for construction of oil storage tanks

by the current version of the American Petroleum Institute Standard 650,

Welded Steel Tanks for Oil Storage [7.4],

The steel plates from this tank are identified as plain carbon steels typical

of common structural steels. No additional alloying elements were present in

amounts above the levels normally considered as residual levels. The ratio of

the manganese to carbon in these steel plates was less than three. To reduce

the tendency for brittle fracture, current steels for welded construction are

77
specified to have a ratio of manganese- to -carbon of not less than three [7.5].

The significance of this manganese to carbon ratio was not known or used at

the time this tank was originally constructed.

The very low silicon content (less than 0.001 percent) in sample 3-2 from the

second shell course indicates that this steel plate probably was produced as a "rimmed" steel. The slightly higher, silicon content (0.026 percent) of the

steel plate from the first shell course (sample 5-3) is still quite low for

steels of this type and indicates that this steel plate also probably was

produced as a "rimmed" steel. The present API 650 permits ASTM A- 283 (Grade

C) steel to be used as a Group I steel for oil tank construction, but requires

that the steel be produced as a semi-killed steel when fracture toughness

criteria apply. "Rimmed" steel is not allowed and therefore the steel in this

tank does not meet the requirements of the present API 650.

Chemical analyses were performed at 8 locations within the specimen taken from

a section of the horizontal weld (test sample 4E) as shown in Figure 7.6. The

results of these chemical analyses are shown in Table 7.1 identified as samples 1 through 8. Sample location 1 was in the base metal from the second

shell course. The chemical composition at this location is consistent, within

the expected normal variability in composition, with the chemical composition

of the same base metal plate (sample 3-2) at a different location. Sample

locations 2, 3, and 4 are in the base metal from the first shell course. The

chemical compositions at these three locations (2, 3 , and 4) are consistent

with the chemical compositions determined in corresponding locations in the

base metal sample from the first shell course (test sample 5-3). Although the

base metal from the first shell course showed significant banding in the

microstructure (Figure 7.6), no evidence of significant segregation was shown

by the chemical analyses

.

The chemical composition was determined for selected areas in the horizontal

weld (test sample 4E) . Areas of "new" weld deposited when the tank was

reassembled at the Floreffe site were identified by metallographic examination

of the weld metal and could be distinguished from the "old " welds deposited

when the tank was originally constructed near Cleveland.

78
.Locations 6 and 8 in the weld metal sample (Figure 7.6) have been identified

to be in the "new" weld metal (the weld made when the tank was reassembled at

the Floreffe site)

.

The chemical composition of the weld metal at locations 6

and 8 are considered to be equivalent within the expected range of variability

for chemical composition of weld metal. Locations 5 and 7 in the weld metal

sample (Figure 7.6) have been identified to be in the "old" weld metal that

was deposited when the tank was originally fabricated at the site near

Cleveland. The chemical composition at locations 5 and 7 are considered to be

equivalent within the normally expected range of variability for chemical

composition of weld metal.

The chemical composition of the "new" weld metal is significantly different

from the chemical composition of the "old" weld metal. This indicates that

the "new" welds most likely were made with a different type welding electrode

and/or by using a different welding process than the "old" welds. The

significant difference in chemical composition is that the "new" weld metal

had a nominal chemical composition 0.90 per cent manganese and 0.40 per cent

silicon whereas the "old" weld metal had a nominal chemical composition of

0.50 percent manganese and 0.16 per cent silicon. No attempt was made to

identify the specific type of welding electrode used from the chemical

composition.

7.2.3 Tensile Tests

Standard, round (0.500 in diameter) tensile test specimens made according to

ASTM Standard E-8-86 [7.6] were taken in the longitudinal (parallel to the

plate rolling direction) and transverse (perpendicular to the plate rolling

direction) orientations from test sample 5E (the first shell course) and from

test sample 3E (the second shell course) to evaluate the strength and

ductility of the steel in the tank at room temperature. The results of these

tests are summarized in Table 7.2. All tests were conducted according to the

procedures specified in ASTM Standard E-8-86. At least three specimens were

tested from each of the two shell courses in each of the two specimen

orientations. No significant differences were found in the tensile strength,

yield strength or elongation between specimens taken from the first shell

course and specimens taken from the second shell course. No significant

79
effect of orientation relative to the plate rolling direction was found for

the properties determined from the tensile tests.

Because there was no significant effect of specimen orientation and shell

course plate found in the tensile tests results, the average value from all

test specimens will be used for the calculations shown in later sections of

this report. The value used for the yield strength is 34,000 psi and the

value used for the tensile strength is 67,100 psi.

The tensile properties determined for the steel from both the first and second

shell course of the tank meet the requirements specified in current ASTM

Standard A-283-84 for Grade D and of the previous ASTM Standard A-10-39 in use

at the time the tank was originally built. The current ASTM Standard A-283-84

specifies that for Grade D material, the tensile strength should be between

60,000 and 72,000 psi and the yield strength should be greater than 33,000

psi. However, API 650 only allows ASTM A- 283 Grade C steel which has an

allowed tensile strength range of 55,000 to 65,000 psi. Therefore, the steel

in this tank does not meet the requirements for a ASTM A-283-84 Grade C steel

that is permitted by API 650. The average elongation of 34.4 percent

determined from these tensile tests exceeds the minimum of 23 per cent

required by the current ASTM Standard A-283-84.

7.2.4 Charpy-V-Notch Tests

Charpy-V-notch tests were conducted on material from the first shell course

(test sample 5E) and from the second shell course (test sample 3E) . Standard,

full size (10 x 10 mm) Charpy-V-notch specimens were made and tested according

to ASTM Standard E-23-86 [7.7]. Specimens were taken from each plate in the

longitudinal orientation (with the axis of the specimen parallel to the

rolling direction of the plate) and in the transverse direction (with the axis

of the specimen perpendicular to the rolling direction of the plate) . All

specimens were notched in the through thickness direction, that is with the

axis of the notch perpendicular to the plate surface. At least two specimens

from each plate in each orientation were tested at each of the selected test

temperatures. The tests were conducted at nominal temperatures of 20° F,

40° F, 60° F, 80° F, 120° F and 160° F to determine the full transition

80
temperature curve for the steel plates. For each test specimen, the energy

absorbed, lateral expansion, and percent shear fracture appearance were

determined.

The results of the Charpy-V-notch tests for each plate and each orientation

are summarized in Tables 7.3, 7.4, 7.5 and 7.6 and the values of the energy

absorbed are shown Figures 7.7, 7.8, 7.9, and 7.10. For test temperature

below 60° F all specimens exhibited low property values as determined by the

Charpy-V-notch specimens. At all temperature of 55° F and below in each of

the plates tested, the energy absorbed was less than 10 ft-lbs, the lateral

expansion was 15 mils or less, and the amount of shear fracture was less than

30 per cent This indicates substantial tendency for brittle fracture at

temperatures of 55° F and below. The energy absorbed and lateral expansion

were essentially the same at 20° F as at 40° F indicating that the steel was

operating in the "lower shelf" region as determined by the Charpy-V-notch

tests. The steel plates did not exhibit significant toughness, as measured by

the Charpy-V-notch tests until temperatures of 100° F and above. Specimens

exhibited greater than 50 per cent shear fracture appearance only above a temperature of 95° F.

The plates from the first shell course (test sample 5E) and from the second

shell course (test sample 3E) do not show significant differences in

properties determined from the Charpy-V-notch tests at temperatures below

60° F. At higher temperatures, the plate from the first shell course (test

sample 5E) appears to be slightly tougher than the plate from the second shell

course (test samples 3E)

.

At temperatures below 60° F, the orientation of the

test specimens with respect to the rolling direction of the plate does not

significantly affect the properties determined by these tests. At higher

temperatures, the specimens oriented in the longitudinal direction are

slightly tougher in both of the test plates.

The ductile- to-brittle transition temperature can be determined from the

Charpy-V-notch tests results. A different value of the transition temperature

is estimated depending on the choice of parameter used to define the

transition temperature [7.8]. The transition temperature can be defined as

(1) the temperature at which the energy absorbed exceeds 15 ft-lbs, (2) the

81
temperature at which the lateral expansion exceeds 15 mils, or (3) the

temperature at which the shear fracture appearance is greater than 50 percent.

Each of these measures of the transition temperature has been found to

correlate with service experience or to predict failure for specific types of

steels or structures. The most commonly used transition temperature criterion

for low strength steels is the 15 ft- lbs absorbed energy. The 15 ft- lb

criterion is used in API 650 to specify the transition temperature of steels.

A summary of the transition temperature for steel plates from the first and

second shell courses is shown in Table 7.7. It should be noted that the

transition temperature for the steel plates in this tank was well above the

operating temperature of the tank by all of the criteria used to define

transition temperature. This means that the tank was definitely operating at

a temperature at which the steel is expected to be brittle.

7.2.5 Drop-Weight Nil-Ductility (NDT) Tests

Drop-weight tests were conducted to determine the nil-ductility transition

(NDT) temperature for specimens taken from the first shell course (sample 5E)

and from the second shell course (sample 3E)

.

The specimens were prepared and

the tests were conducted according to ASTM Standard E-208-85 [7.9]. Standard

Type P-2 specimens (3/4 in thick by 2 in wide by 5 in long) were prepared from

the first and second shell course plates. In addition, subsize Type P-3

specimens (5/8 in thick by 2 in wide by 5 in long) were tested from the first

shell course plates to determine if specimen size had a significant effect on

the NDT temperature. Tests were conducted at temperatures in the range of

40° F to 80° F to determine the nil-ductility transition (NDT) temperature.

The results of the test are shown in Table 7.8 for all plates tested. The NDT

temperature for the plate from the first shell course (test sample 5E) was

+60° F for the standard Type P-2 specimens and was +50° F for the subsized

Type P-3 specimens. This difference in NDT temperature between the two

different specimen sizes is not considered to be significant. For purpose of

analysis the NDT temperature of the first shell course plate will be taken as +60° F. The NDT temperature for second shell course plate (test sample 3E)

was determined to be +70° F.

82
The NDT temperature is defined as the maximum temperature where a standard

drop-weight specimen breaks. This represents the temperature at which fracture

in the structure can be initiated by a small flaw (length less than one inch)

when stresses are at yield point levels. This temperature is significant

because at temperatures below the NDT temperature, fracture initiation from

relatively small flaws is possible and fracture arrest is unlikely to occur

once the fracture is initiated. The NDT temperature is one measure of the

temperature below which the structure behaves in a substantially brittle

manner. The operating temperature of the tank at the time of failure was

estimated to be approximately 32° F. Therefore, these test show that the tank

was operating at a temperature substantially below the NDT temperature of the

steel plates in the tank.

7.2.6 Fracture Toughness Tests

The fracture toughness was determined for the steel plate from the first

shell course (test sample 5E) and from the second shell course (test sample

3E) . The fracture toughness measured from these samples is representative of

the fracture toughness of the base metal in the tank. The fracture toughness

was determined using the ASTM E-813-87 "Standard Test Method for Jjc> a Measurement of Fracture Toughness" [7.10]. Standard IT size compact tension

test specimens were used to determine the fracture toughness. The single

specimen, compliance test method option was used for determining the Jjc. All

specimens were in the LT orientation as defined by ASTM Standard E- 399 -83

[7.11], that is the specimen was notched so that the crack propagated

transverse to the rolling direction of the plate. This orientation matches

the direction of crack propagation of the primary fracture in the ruptured

tank. The specimens from the first shell course plate (test sample 5E) were

nominally 0.750 in thick with side grooves to a depth of nominally 0.075 in on

each side of the specimen. The specimens from the second shell course plate

(test sample 3E) were nominally 0.625 in thick with sides grooves nominally

0.075 in deep on each side of the specimen.

The fracture toughness was determined over the temperature range of 20° F to

105° F. At least two test specimens from each plate were tested at each

temperature from 20° F to ambient temperature (about 70° F) . The fracture

83
toughness test results are summarized in Table 7.9.

The fracture toughness was determined according to the ASTM Standard test

method E-813-87 instead of according to the ASTM Standard test method E-399-83

for Plane-Strain Fracture Toughness (Kj^) of Metallic Materials because

preliminary calculations based on the estimated Kjq and the yield strength of

the steel plates indicated that test specimens sufficiently thick to meet the

validity requirements of E-399-83 could not be obtained from the available

steel plates. The ASTM E-813-87 test method determines the property which

is a measure of the toughness of the material at the onset of crack extension.

The requirements for the provisional value of Jq to be considered a valid Jjq

are specified in the ASTM Standard E 813-87. The requirements are:

1) The specimen thickness B is > 25 Jq/cry,

2) Initial specimen ligament, b 0 is > 25 Jq/a

y

>

3) The slope of the power law regression line, dJ/da, evaluated at Aaq is

less than ay,

4) No specimen demonstrates brittle cleavage fracture at the applicable

test temperature and rate,

5) None of the nine physical measurements of the crack size differ by

more than 7% from the average physical crack size,

6) Neither of the two near surface crack extensions measurements or

surfaces of plane sided specimens differs from the crack extension at

the center by more than .02W.

7) The difference between the crack extension predicted by elastic

compliance (or other method) at the last unloading and the average

physical crack extension, Aap , does not exceed 0.15 Aap for crack

extensions less than Aap(max) and 0.15 Aap(max) thereafter. E^ must

not differ from E by more than 10%.

84
The total J integral is made up of two components, the elastic part identified

as J e ]_, and the plastic part identified as Jpi- The elastic part is defined

as

[K (i )]

2 [l-u2

] Jel =

E

where

K (i) =

f(aQ/W) is the compliance calibration for the specimen used

JpL =

* apl a>

(Bfj) (bo)

where

Bjj = net specimen thickness

bQ = uncracked ligament, and

r) = ( 2 + 0.522 b0 )/W

In the tests conducted here, the Jjc tests did not give completely valid

values at the lowest test temperatures (20° F, 40° F and 70° F) because there

was little or no physical crack growth prior to failure of the specimens and

the specimens failed almost completely by cleavage fracture. In these cases,

as described in reference [7.12], a Jc value is obtained by taking the J

integral at the point of crack initiation or instability to be the elastic

component (Jel) as described above.

O _ For specimens tested at 20° F, 40 F, and 70° F, the J versus crack growth

curve did not show significant stable crack growth. This was confirmed after

the test by cooling the specimen to liquid nitrogen temperature and then

breaking the specimen open to examine the appearance of the fracture surface

for evidence of stable crack growth prior to final fracture. No significant

stable crack growth was found and the fracture surface was entirely brittle in

appearance for all specimens tested at 20° F, 40° F, and 70° F.

Pi

(B Bn W)H f(aoA0

85
The only completely valid Jjc that met all of the requirements shown above was

obtained for the specimen tested at 105° F. This was the only specimen that

showed sufficient stable crack growth prior to final fracture of the specimen.

The measured values of Jjc or were used to estimate the plane strain

fracture toughness (Kjq) or the plane stress fracture toughness (K^) of the

steel plates using the following relationship from reference [7.13]:

(Kc) 2 = (Jc) (E)

where

E = Young ' s modulus

The estimated values of Kc are reported in Table 7.9 and are used in the

fracture analysis shown later in this report.

The fracture toughness of the first shell course plate (test sample 5E) was

not significantly different from the fracture toughness of the plate from the

second shell course (test sample 3E) . The main interest of this study is

focused on the toughness of the plate from the first shell course because that

is where the fracture is believed to have originated. The fracture toughness

in the plates from the first shell course was about 48 ksiJTn at 40° F and was

47 ksiJTn at 20° F. At 70° F, the fracture toughness was slightly higher,

about 51 ksi4 in.

7.2.7 Notched Bend Tests

A region of hardened metal with a maximum hardness on the Rockwell C scale

(HRC) of 40 and a depth of 0.050 in was found around the defect on the

fracture surface that was believed to be the origin of the fracture [7.14],

This region was expected to have a toughness less than the toughness of the

normal steel plates in the tank. To verify this, special tests were conducted

to simulate the effect of a defect located in heat affected or embrittled

region. Single edge notched bend specimens were prepared from the first shell

course plate steel. The specimens were nominally 3/4 in by 1 inch in cross

section and 10 in long. A notch to simulate a defect was put in the center of

the specimen as shown in Figure 7.11. The notch was made by electrodischarge

86
machining. The notch was 0.100 in deep, 0.010 in wide with an 0.005 in root

radius. Some of the specimens were embrittled to simulate a heat affected

zone by heating the surface of the specimen with gas- tungsten- arc (GTA)

welding torch without melting the surface of the metal. This was done both

before and after notching. As seen in Figure 7.12, heating with the welding

torch produced an altered microstructure and hardened zone similar to the heat

affected region found around the fracture surface defect described below.

The specimens were tested in three-point bending. The loading rate was held

constant at 0.01 in/minute. Knife edges were attached to the mouth of the

notch to monitor the notch-mouth- opening displacement. An environmental

chamber was used to maintain the test temperature at 40° F. A load versus

notch-mouth opening displacement curve was determined for each test specimen.

One specimen with only a notch (without heating to cause embrittlement) was

tested at room temperature (70° F) . This specimen did not fail in a brittle

manner but instead became fully plastic .and did not fracture.

The results of additional specimens tested at 40° F are summarized in Table

7.10. A typical load versus notch-mouth-opening displacement curve for a notched only and notched plus heated specimen are shown in Figure 7.13. For

each specimen, the maximum load at final fracture, the load at the onset of

plastic deformation (yield load) , and the notch-mouth- opening displacement at

maximum load were determined. The area under the load versus displacement

curve up to the point of fracture (maximum load) was calculated as a measure

of the energy (expressed in ft- lbs) to initiate fracture in the specimen. A

distinct brittle fracture occurred for each of the specimens tested at 40° F.

The specimen that was notched only, to simulate a defect in unaltered base

metal, had a maximum load at the onset of fracture about 20 percent higher

than the specimens that had a notch located in an embrittled and hardened

region to simulate a defect in a heat affected region. The notch-mouthopening displacement was more than twice as large for the specimen that only

had a notch as for the specimens that were notched and heated. The total

energy required to initiate fracture, taken as the area under the loaddisplacement curve, was more than twice as large (21.7 ft- lbs) for the notched

only specimen as for the notched plus embrittled specimens (9.0 and 9.7 ft87
lbs) . No significant difference was found between the specimens that were

heated first and then notched and the specimens that were first notched and

then heated.

These results show that it is much easier to initiate fracture at a defect in

the steel base metal plate used in this tank if the defect is in a region that

has been hardened and embrittled by heating. These results indicate that the

toughness of the hardened and embrittled region may be less than one half the

toughness of the normal steel plates.

7 . 3 Metallographic and Hardness Evaluation

7.3.1 Base Metal Evaluation

Additional characterization of the steel plates from the first and second

shell courses was conducted by hardness tests and metallographic examination.

The results of hardness tests on several samples are summarized in Table 7.11.

The average hardness of the first shell course plate was 70 HRB (Rockwell B

scale) and the average hardness of the second shell course plate was 71 HRB.

These values are consistent with the measured tensile strength of

approximately 67,000 psi. Typical microstructures of samples taken from the

first and second shell course plates are shown in Figure 7.14. The

microstructure consists of a uniform distribution of ferrite and pearlite,

which is typical of low strength carbon steel. The ferrite grain size was

determined to be ASTM grain size number 7.

7.3.2 Evaluation of Welds

The horizontal weld (test sample 4E) and the T-weld (test sample 9) were

radiographed and then were evaluated by metallographic, hardness and Charpy-Vnotch tests. The radiographic examination showed that the entire length of

the horizontal weld from test sample 4E contained many weld defects including

lack of fusion, nearly continuous lack of penetration, and extensive areas of

porosity. The horizontal weld section was radiographed a second time after

being ground from the original thickness of 7/8 in to a thickness of 1/2 in to

88
remove the weld crown and to attempt to remove the regions of lack of

penetration. The second radiograph of the horizontal weld, shown in Figure

7.15, still shows extensive areas of porosity and lack of penetration. These

results are confirmed by the metallographic examination results discussed

below. The radiograph of the T-weld section (test sample 9), which includes

the only section of vertical weld available for this investigation, showed

similar extensive areas of weld defects such as lack of fusion, lack of

penetration, and porosity. No attempt was made to determine from the

radiographs if the defects found were located in the "old" welds used to

fabricate the original tank or if the defects were in the "new" weld used to

reassemble the tank at the Floreffe site. Although a detailed assessment was

not made of the radiographs from this limited weld sample, the quality of the

welds examined here would not satisfy the requirements of the current (1984)

version of API 650.

Transverse sections of the welds were taken in the as received condition from

nine different locations along the horizontal weld in test sample 4E.

Photomicrographs were taken of each section to characterize the weld metal and

to determine the type and number of weld defects in each specimen. Hardness

measurements (Rockwell B hardness) were taken on each specimen at several

locations within the old weld, new weld, and heat affected zones. Similar

sections taken from four locations in the horizontal weld from test sample 4E

(after the sample had been ground to a thickness of 1/2 in to remove the weld

crowns) were evaluated in the same manner. Four transverse sections of the

vertical weld in the test sample 9 containing the T-weld were taken to

characterize the vertical weld.

Figure 7.16 shows a typical transverse section of the horizontal weld. The old

weld is identified as the weld metal connected to the first shell course (the

right side in Figure 7.16) and the new weld metal is connected to the second

shell course. In this weld section, a significant area of lack of penetration

is found at the junction between the old weld metal and new weld metal. This

defect extends to a depth of about 40 percent of the thickness of the plate. A

large area of porosity was found in the center of the old weld metal and

numerous areas of small porosity were found in the upper part of the new weld

metal

.

89
Numerous hardness measurements were taken on this weld sample at location in

the base metal, old and new weld metal, and heat affected zones of the old and

new weld metal. As shown in Figure 7.16, the base metal had a Rockwell B

(HRB) hardness of 81 in the first shell course plate and a hardness of 72 in

the second shell course plate. The hardness of heat affected zones and old

and new weld metals ranged from HRB 85 to HRB 96. A hard spot, believed to be

an arc strike was found on the surface of the second shell course plate and

had a Rockwell C (HRC) hardness of 33. Another hard spot with a hardness of

HRC 28 was found in the surface "cap" pass of the weld and a very hard spot of

HRC 43 was found in the vicinity of the lack of penetration defect. High

magnification micrographs of the weld metal, heat-affected zones, and hard

spots are shown in Figure 7.17. The microstructure were typical of weld metal

and heat-affected zones in welds of this type of steel. There was evidence of

martensite in the hardened regions. The microstructures, hardness values, and

weld defects found in this specimen are typical of the those found in all of

the transverse weld specimens that were evaluated.

Additional specimens from the horizontal weld are shown in Figures 7.18 and

7.19. These specimens are shown to indicate the type and extent of weld

defects that were found in the horizontal weld. Numerous areas of porosity

and lack of fusion were found in nearly all of the weld specimens. Hardness

values (HRB) for the weld metal and base metals are shown on each of the

specimens in Figures 7.18 and 7.19.

Specimens taken from the vertical weld (test sample 9) are shown in Figure

7.20. These specimens show a region identified as base metal separating the

old weld metal from the new weld metal. This indicates that the old weld was

not fully removed when the tank was dismantled and reassembled. The

microstructure and hardness values in this vertical weld were similar to those

found in the horizontal weld described above. Numerous weld defects,

particularly lack of fusion, are found in all specimens of the vertical weld

shown in Figure 7.20.

An attempt was made to evaluate the toughness of typical weld metal using

Charpy-V-notch tests of specimens taken from the horizontal weld (test sample

4E) . The samples were taken from the best portion of the weld identified by

90
radiography on the weld that had been ground to 1/2 in thick to remove the

weld crown. The Charpy-V-notch specimens were etched before being notched and

the notch was located in either the old weld metal, the new weld metal, or the

new weld metal heat affected zone. The weld defects, particularly porosity,

were so extensive that several of the specimens broke at the defect rather

than at the notch of the Charpy specimen and valid results of the test could

not be obtained. The results of satisfactory tests conducted at 40° F and at

room temperature (71° F) are shown in Table 7.12.

These results show that the toughness of the weld metal at both room

temperature and at 40° F is significantly better than the toughness of the

steel plates from which the tank was fabricated.

7 . 4 Characterization of the Fracture Surface Defect

All characterization of the fracture surface defect was done by Battelle and

the results reported here and all photographs of the deferct were supplied to

the NBS by Battelle [7.15]. The NBS reviewed all test procedures with

Battelle before the characterization of the defect was started and

subsequently examined all test specimens used in the characterization.

A defect that extended part way through the wall of the tank was found on the

surface of the major vertical fracture in the tank. This defect was on the

inside of the tank in the first shell course plate at a location just below

the junction of a T-weld joint as shown in Figure 7.21. This defect is

believed to be the origin of the major vertical fracture in the tank because

"chevron" markings on the surface of the fracture point to this defect from

above and below the defect, as shown in Figure 7.22. No chevron markings or

other defects were found elsewhere on the fracture surface that would indicate

another origin of the fracture so it was concluded that this defect was the

origin of the vertical fracture in the tank. When present, these chevron

markings have been identified as one method of determining the origin of

fracture in steel plates. Test samples, as shown in Figure 7.23, were cut out

of the tank in the vicinity of this defect for detailed investigation of the

exact location and features of this defect.

91
The location of the defect on the fracture surface is shown at a magnification

of 4X in Figure 7.24. This figure shows that the defect was located in the

first shell course plate on the inside of the tank, below the horizontal weld . The size of the defect was determined from Figure 7.24 and is shown on the

Figure. These are the dimensions of the defect that were used in the fracture

analysis shown in Chapter 8. Other features of the defect and the fracture

surface that were determined by examination with the scanning electron

microscope are shown on Figure 7.24.

To determine the exact location of the defect with respect to the new weld

metal, old weld metal, and first shell course plate, the sample shown in

Figure 7.24 was sectioned as shown in Figure 7.25. The sections were cut

using a fine wire electrodischarge machine. Eight specimens were cut at

intervals of 0.080 in starting from the inner surface of the tank. These

specimens were then examined metallographieally to permit a 3-dimensional

reconstruction of the defect to be made and to determine the exact location of

the defect. An example of one of these sections is shown in Figure 7.26.

This figure shows that the defect was located entirely in base metal from the

first shell course. The edge of the defect extended just to the edge of the

heat affected zone of the old weld metal.

Battelle concluded that the weld metal at the edge of the defect was "old"

weld metal rather than "new" weld metal and therefore that the defect was

present prior to dismantling the tank for reconstruction at the Floreffe site

[7.15]. Battelle based this conclusion on the following evidence: (1) In the

horizontal weld, the old horizontal weld metal was always found to lie below

(that is, on the first shell course side) the new horizontal weld metal, (2) a slice (sectioned similarly to the slices shown in Figure 7.25) that was taken

from a region of the horizontal weld about 1 inch from the location of the

defect had a chemical composition and microstructure representative of the old

weld metal rather the new weld metal (see Table 7.1) and (3) as shown in

Figure 7.26, the heat-affected zone surrounding the defect did not go through

or alter the heat-affected zone of the weld metal adjacent to the defect. The

NBS did not have access to material in the vicinity of the defect to

independently verify this conclusion. However, the NBS did examine the

me tallographic specimens prepared by Battelle as well as do an independent

92
evaluation of typical horizontal welds. From these observations, the NBS did

not find any evidence to conflict with the Battelle conclusion that the defect

was below old weld metal and therefore was present prior to the dismantling of

the tank

.

As seen in Figure 7.26, a dark etching region is present surrounding the

defect below the fracture surface. This dark etching region is believed to

represent a region of heat-affected zone caused by heating of the metal,

probably at the time the defect was formed. Hardness measurements made by

Battelle showed that this heat- affected zone had a maximum hardness of

slightly in excess of 400 Knoop (HKN) (equivalent to a Rockwell hardness of

HRC 40) at the edge of the defect and that the hardness decreased with

distance below the defect until a hardness of about HRB 85 was reached at a distance of about 0.050 in below the edge of the defect. The hardness of HRB

85 is representative of the hardness of the base metal plate. Chemical

analyses using the electron microprobe and secondary electron emission imaging

were conducted on the surface of the defect and in the heat- affected zone

below the defect. A heavily carburized layer was found near the surface of

the defect at the location showing the maximum hardness. Formation of such a carburized layer can only occur from a combination of high temperature and

excess carbon in the atmosphere. This is the primary evidence for

concluding that the defect was formed by a cutting torch.

7 . 5 Stress Analysis

A detailed analysis of the stresses in the tank at the time of failure was

conducted to obtain the stresses necessary for the fracture analysis of the

tank. The stress analysis, described in section 5.1 of this report, was

performed on the tank to determine the stresses in the tank wall at the time

of failure.

From the stress analysis the maximum stress at the top of the first shell

course due to the hydrostatic pressure of the oil was calculated to be a circumferential membrane (hoop stress) of 13.7 ksi . Additional stresses due

93
to settling of the tank or thermal effects were not found to be of a significant magnitude and are not included in the stresses used for the

fracture analysis.

Residual stresses were measured by Battelle using the chip removal method

[7.16]. A typical T-weld joint removed from a location remote from the

primary vertical fracture was used for the residual stress measurements. The

results are considered to be representative of the residual stresses in the

vicinity of the defect in the primary fracture but may not represent the exact

stresses in the specific T-weld region near where the defect was located. The

maximum residual stress measured was a tensile circumferential membrane stress

of 15.1 ksi . Residual stresses as high as the yield strength of the base

metal have often been found to be present in the vicinity of welds in steel

plates [7.17] .

For the fracture analysis shown in Chapter 8, the stresses at the location of

the defect were taken to be:

Circumferential membrane stress due to the oil = 13.7 ksi

Circumferential residual stress = 15.0 ksi

Stress at the yield strength of the metal = 34.0 ksi

These stresses are used alone and in combination in the fracture analysis

shown in Chapter 8.

94
Table

7.1

-

Chemical

Composition

of

Tank

Steels

and

Weld

Metal

95
Table 7.2 - Tensile Properties of Plate Steels

Shell

Course

Specimen

Orientation

Tensile ^

)

Strength

(psi)

Yield^ 1 ) »

(

2 ) Strength

(psi)

Elongation^

)

in 2 -inches

(%)

First Transverse 66,900 33,400 33.8

First Longitudinal 66,900 34,500 34.2

Second Transverse 67,300 33,000 34.6

Second Longitudinal 67,200 34,900 34.8

(1) All values are the average of 3 or more specimens

(2) 0.2% offset yield strength

96
Table 7.3 - Summary of Charpy-V-Notch Test Results

First Shell Course - Transverse Orientation ( 1 )

Average

Temperature

(F)

Absorbed

Energy

(ft. -lbs.

)

Lateral

Expansion

(mils

)

Shear

Fracture

(%)

160 37.5 34.4 80

121.5 27.5 26.9 70

79.7 17.3 22.0 40

57.7 9.0 15.8 25

40.1 5.5 12.7 13

20.8 4.0 11.5 0

(1) All values are the average of 2 or more specimens

97
Table 7.4 - Summary of Charpy-V-Notch Test Results

First Shell Course - Longitudinal Orientation ( 1 )

Average

Temperature

(F)

Absorbed

Energy

(ft. -lbs.)

Lateral

Expansion

(mils)

Shear « Fracture

(%)

160.1 53.5 40.0 85

121.5 33.5 30.5 75

80.4 13.0 18.9 38

57.5 8.8 15.4 30

39.8 6.2 13.6 20

Csi «—

1

CM 6.1 12.8 0

(1) All values are the average of 2 or more specimens

98
Table 7.5 - Summary of Charpy-V-Notch Test Results

Second Shell Course Transverse Orientation ( 1 )

Average

Temperature

(F)

Absorbed

Energy

(ft. -lbs.

)

Lateral

Expansion

(mils)

Shear

Fracture

(%)

160.2 32 30.2 70

121.4 21 23.3 50

80.0 9.8 17.0 30

55.4 6.0 13.2 20

40.3 4.8 13.5 10

20.0 4.2 11.5 0

(1) All values are the average of 2 or more specimens

99
Table 7.6 - Summary of Charpy-V- Notch Test Results

Second Shell Course - Longitudinal Orientation ( 1 )

Average

Temperature

(F)

Absorbed

Energy

(ft. - lbs .

)

Lateral

Expansion

(mils)

Shear

Fracture

(%)

160 38.5 33.0 60

121.5 20.3 21.9 50

80.4 11.5 17.0 30

57.8 6.4 13.5 15

40.2 4.4 11.7 10

20.2 3.7 10.8 0

(1) All values are the average of 2 or more specimens

100
Table 7.7 - Ductile- to-Brittle Transition Temperatures Determined from

Charpy-V-Notch Tests

Shell

Course

Specimen

Orientation

15 ft- lb

(F)

15 mils Lateral

Expansion

(F)

50% Shear

Fracture

Appearance

(F)

First Transverse 80 55 95

First Longitudinal 100 55 95

Second Transverse 90 68 120

Second Longitudinal 100 68 120

101
Table 7.8 » Drop-Weight (NDT) Test Results

Shell

Course

Specimen

Type

Temperature

(F)

Test

Results

First P-2 40.0 Break

60.2 Break

60.0 Break

70.0 No Break

70.0 No Break

NDT Temperature for first shell course = + 60F

First P-3 40.4 Break

49.0 Break

49.4 Break

58.8 No Break

59.0 No Break

NDT Temperature for first shell course = + 50F

Second P-2 40.6 Break

61.0 No Break

60.4 Break

71.0 No Break

69.0 Break

80.2 No Break

80.2 No Break

NDT Temperature for second course = + 70F

102
Table 7.9 - Summary of Fracture Toughness Test Results

Specimen

No.

Shell

Course

Temperature

(F)

J C

( in- lb/in^)

KC

(1)

(ksiVin)

5LT-9 First 20 72 46.5

5LT-17 First 20 72 46.5

5LT-13 First 40 76 47.7

5LT-16 First 40 83 49.9

5LT-18 First 40 68 45.2

5LT-14 First 69 92 52.5

5LT-15 First 69 85 50.5

5LT-20 First 71 83 49.9

5LT-21 First 105 1011 174.2

3LT-1 Second 40 72 46.5

3LT-8 Second 40 69 45.5

3LT-7- Second 69 84 50.2

3LT-6 Second 69 81 49.3

(1) KC

2 = J C E

103
Table 7.10 - Summary of Notched-Bend Test Results

Specimen

Condition

Maximum

Load

(lbs.

)

Yield

Load

(lbs.)

Displacement

at Maximum

Load

(in.

)

Energy Absorbed

to Maximum Load

(ft. -lbs.)

Notched only 4748 2783 0.072 21.7

Notched, then

Heated

4037 2904 0.034 9.0

Heated First,

then Notched

3826 2662 0.035 9.7

104
Table 7.11 - Summary of Hardness of Steel Plates

Hardness (HRB)^^

Sample First Shell Second Shell

Number Course Course

421 73 69

422 71 71

423 72 73

424 65 70

- 4141 69 73

4142 68 71

4143 69 71

4144 69 72

Average Hardness 69.5 HRB 71.3 HRB

(1) Rockwell B Hardness

105
Table 7.12 - Charpy- V-Notch Results from Weld Metal

Sample

Number

Location

^ ^ Temperature

(F)

Energy Absorbed

(ft. -lbs.)

2,3,4 New W.M. 40 All failed outside

notch

5 New W.M. 71 85

6 New W.M. 71 117

7 New W.M. 71 57

Average 86

14 Old W.M. 40 21

15 ' Old W.M. 40 37.5

16 Old W.M. 40 7

Average 21.8

17 Old W.M. 71 17

18 Old W.M. 71 23

19 Old W.M. 71 29.5

Average 23.2

First Course B.M. 40 Average 5 .

9

First Course B.M. 60 Average 9 .

0

First Course B.M. 80 Average 17.3

(1) W.M. - Weld Metal

B.M. - Base Metal

106
107
Figure 7.2 - Location of Test Specimens from First Shell Course

108
Figure 7 .

3

Location of Test Specimens from Second Shell Course

109
7.4a - Overall Location of Specimens

7.4b - As Received Weld Sample

Figure 7.4

7.4c - As Ground Weld Sample

Location of Test Specimens from Horizontal Weld

110
Figure 7.5 - Location of Test Specimens from Vertical Weld

111
Horizontal Weld Metal

Figure 7.6 - Location of Samples for Chemical Analysis

112
(qi-U) A0d3N3

113

Figure

7.7

-

CharpyV-Notch

Tests

for

First

Shell

Course

Transverse

Orientation
(Pl-H) ADH3N3

114

Figure

7.8

-

CharpyV-Notch

Tests

for

First

Shell

Course

Longitudinal

Orientation
LL

o

LU

CC

<:

CC

LU

CL

LU

h"

O

CD

O O O O

LO CO - C\J

(Ql-U) A9U3N3

o

115

Figure

7.9

-

CharpyV-Notch

Tests

for

Second

Shell

Course

Transverse

Orientation
LL

o

LJJ

H

(Ql-U) A0U3N3

116

Figure

7.10

-

CharpyV-Notch

Tests

for

Second

Shell

Course

Longitudinal

Orientation
*co

*—> 0D c

CDsO

<UD.

C/5 •OC *3

<D 3

CQ 4J

C/5

Ao w w C 0 a) z e 1 <u

a) ~-i bO -u T3 U W -H

<D 0 —t w

bOC J-i •h o

C/5 C4_j

a>3

bO

117
Figure 7.12a - Embrittlement Zone at Notch Tip in Single-Edge

Notch Bend Test

Depth (in)

, 0.020

0.040

0.060

0.080

0.100

0.120

0.140

0.160

Hardness

HRC 35

HRC 36

HRC 37

HRC 43

HRC 31

HRC 27

HRB 85

HRB 80

Magnification: 21X

Figure 7.12b Hardness of Embrittled Zone in Single-Edge

Notched Bend Specimens

118
5000 o o o o

<* CO CVJ T_

(•sq|) QV01

119

NOTCH-MOUTH-OPENING

DISPLACEMENT

(in.)
7.14a - Microstructure of First Shell

Course Plate

7.14b - Microstructure of Second Shell

Course Plate

Figure 7.14 - Microstructure of Steel Plates

Magnification: 100X Etchant: 2% Nital

120
Figure 7.15 - X-ray Radiograph of Horizontal Weld after Surface Grinding

121
Figure 7.16 - Microstructure and Hardness of Horizontal Weld Metal

Rockwell B and C Scale Hardness Values are

given for various Locations

.

Inside

122
N

<

X

123

Figure

7.17

-

Microstructure

of

Horizontal

Weld

Metal
1st Course 2nd Course

Figure 7.18 - Defects in Horizontal Weld Metal

124
Figure 7.19 Defects in Horizontal Weld Metal

after Grinding Welds

125
Figure 7.20 Defects in Vertical Weld Metal

126
Figure 7.21 - Location of Fracture Surface Defect

127
Figure 7.22 - Fracture Surface Defect

128
Figure 7.23 - Location of Fracture Surface Defect

From Outside of Tank

129
Outside

of

Tank

I

Second

First

Shelf

Course

1

Weld

Metal

—

Shell

Course

130
To top of tank

Horizontal

weld

/ / /

Weld '/ /

metal

/

/

/

/

/

/

(Slice as shown in Fig. 7.26)

Figure 7.25 - Details of Sectioning to Determine

Location of Defect

131
Defect

132

Figure

7.26

-

Transverse

Section

of

Fracture

Surface

Defect

Showing

Exact

Location

of

Defect
8.0 ANALYSIS OF TANK FAILURE

8 . 1 Scope of Tank Failure Analysis

The tank was analyzed to determine (1) if the defect found on the fracture

surface initiated the fracture and (2) why the fracture propagated and caused

the tank to rupture instead of arresting and only causing the tank to leak.

To predict if the defect found on the fracture surface could be expected to

initiate fracture with the stresses determined to be present in the tank and

at the operating temperature of the tank, several established failure

assesment methods were used. These included predictions based on ductile-tobrittle transition temperature methods using Charpy-V-notch tests or the nilductility (NDT) temperature, and analyses based on linear-elastic fracture

mechanics principles to predict quantitively if the defect found on the

fracture surface was responsible for fracture initiation. The linear-elastic

fracture mechanics analyses were performed to compare the applied stress

intensity factor calculated from the applied stress and defect size with the

fracture toughness of the material. The analyses were done in sequence by

first making the assumption that the defect found on the fracture was a partthrough- surface defect and next by assuming that the part- through- surface

flaw was converted (by failure of the uncracked ligament below the surface

defect) to a through- thickness defect. To predict if fracture propagation is

to be expected once the fracture had initiated, an assessment of fracture

propagation and arrest based on the measured nil-ductility- transition (NDT)

temperature was made. A determination of the most probable cause of the

failure was made and an assessment was made of possible means to reduce or

prevent similar failures from occurring in other storage tanks.

8 . 2 Assessment of Fracture Initiation and Arrest

The simplest assesssment of the tendency for fracture initiation in the steel

plates of the tank is based on the transition temperature of the steel.

Extensive service experience with fracture in mild steel structures,

particularly welded ships, and correlations with Charpy-V-notch tests has

shown that fracture initiation is likely for steel plates that have an energy

133
absorbed in the Charpy-V- notch test of less than 10 foot-pounds at the

temperature at which they are being used. Fracture arrest does not occur for

steel plates that have a Charpy-V- notch energy of less than 20 foot-pounds

[ 8 . 1 ].

It has become common practice to specify that mild steel plates must have a minimum Charpy-V- notch energy of 15 foot pounds at the lowest operating

temperature at which the steel will be used [8.2]. The temperature at which

the steel has a minimum Charpy-V-notch energy of 15 foot-pounds has been

accepted as the definition of the ductile to brittle transition temperature

for low to medium strength structural steels. This is considered a conservative definition of the highest temperature at which fracture

initiation can occur in steel. The current requirement for selecting steels

for oil tanks constructed as specified in the API Standard 650 requires that

the steel have a minimum Charpy-V-notch energy of 15 foot-pounds at the design

metal temperature which is +5° F for this tank. For these tanks, this minimum

Charpy-V-notch energy is considered to be sufficient to prevent fracture

initiation and to arrest any propagating fractures.

As can be seen from Table 7.7, the 15 foot-pound transition temperature of the

steel plates used in both the first and second shell course plates was well

above (80 F to 100° F^ the operating temperature (32 F) at the time of

failure of the tank. Therefore, fracture initiation and propagation is

predicted to be possible.

A second, empirically determined, definition of the transition temperature

below which fracture initiation is expected to occur is the temperature at

which the lateral expansion in the Charpy-V-notch specimen is 15 mils [8.3].

This definition is also considered a conservative estimate of the lowest

service temperature at which the steel can be used without risk of fracture

initiation. For higher strength steels, the 15 mils lateral expansion

criterion has been found to be a more suitable definition of the transition

temperature than the 15 foot pound definition [8.2]. The 15 mils lateral

expansion criteria is less conservative (defines a lower transition

temperature) for the low strength structural steel such as that used in this

tank. As can be seen from Table 7.7, the 15 mil lateral expansion transition

134
temperatures of the steel plates used in this tank (55° F to 68° were also

O substantially above the operating temperature (32 F) at failure of this tank

and therefore this method predicts that fracture initiation is possible.

A more exact, but still empirically based, procedure to assess the

probability of fracture initiation and arrest is the determination of the

drop-weight nil-ductility transition (NDT) temperature. This NDT temperature

is used in conjunction with the generic Fracture Analysis Diagram (FAD) , shown

in Figure 8.1, to make an assessment of the probability of fracture initiation

and arrest [8.4]. The Fracture Analysis Diagram provides a generalized

description of the relationships between the crack size, relative stress, and

operating temperature relative to the NDT. At temperatures below the NDT, the

Fracture Analysis Diagram predicts that fracture will initiate from small

defects (typically less than 1-inch in length) at stress levels below the

yield strength of the steel and that the fracture will be brittle in nature.

As seen in Table 7.8, the NDT for the first and second shell course steel

plates was 20° F to 40° F above the operating temperature of the tank at the

time of the failure. Based on the measured NDT and the Fracture Analysis

Diagram,- fracture initiation is predicted to be likely to occur in this tank

o at the operating temperature of 32 F.

The NDT temperature also can be used to make a qualitative assessment of the

probability of fracture propagation or arrest once fracture has been

initiated. A crack arrest temperature (CAT) curve is defined by the Fracture

Analysis Diagram and shows that fracture propagation is expected to occur at

and below the NDT temperature even when the stress levels present are very low

(<10 per cent of the yield strength) . Fracture propagation is predicted at

stress levels of approximately one half the yield strength level at

temperatures about 30° F above the NDT temperature. Fracture arrest is not

expected at stress levels equal to the yield strength of the steel until a temperature of approximately 50° F above the NDT temperature is reached.

135
8.3 Fracture Analysis

8.3.1 Basis for Fracture Analysis

A more quantitative and deterministic analysis to predict fracture initiation

can be made on the basis of linear elastic fracture mechanics principles. The

objective of this analysis is to calculate the applied stress intensity in the

structure due to the presence of a defect and the known applied stresses and

then to determine if this applied stress intensity exceeds the critical value

stress intensity (Kjc) for initiating fracture in the material. The data used

for this fracture analysis came from the measured materials properties, the

measured size of the identified surface defect, the stress and thermal

analyses done by the NBS , and the residual stress measurements conducted by

Battelle. The values of all parameters used in the analyses that follow are

summarized in Table 8.1. The fracture analysis was done using materials

o properties values measured at 40 F. The actual temperature of the metal in

the tank was calculated to be 32° F but the small difference in temperature

between the tank temperature and the temperature at which the materials

properties were measured was not found to significantly affect the values of

the materials properties. For the following fracture analyses, the three

specific stress levels that were analyzed are: (1) stress only from the

hydrostatic pressure of the oil (2) stress from the hydrostatic stress plus

the measured residual stress, and (3) stress equal to yield strength of the

steel plates.

The fracture analysis conducted here was based on linear elastic fracture

mechanics principles. The analyses were done only for the defect located in

base metal in the first shell course and the stresses used were those

calculated at the position of this defect and all materials properties were

those measured for the base metal in the first shell course. For a plane

strain stress state to exist, the following relationship must be satisfied

[Reference 8.5]:

T . -\

t > 2.5

KC

V " y y

136
where

K(] = Toughness

cjy = yield strength

t = thickness

Using the thickness of the steel plate as 0.875 -inch, the toughness of the

steel plate as 47.7 ksiJTn, and the yield strength as 34.0 ksi, it was

determined that the shell of the tank was not in plane strain but should be

analysed as a plane stress problem.

For purposes of the fracture analysis-, it was determined that the curvature of

the tank could be ignored and that the tank could be assumed to be a flat

plate. This conclusion was reached because the tank radius to tank wall

thickness ratio (R/t) was very large (>820) . It has been determined that

under some conditions, a part- through-wall or through-wall defect may cause a local increase in the stress in a thin walled, pressurized structure due to

small amount of bulging of the wall in the vicinity of the defect. This

increase in local stress is accounted for using a stress magnification factor

(Mp) based on an analysis by Folias [8.6]. The stress magnification factor is

calculated by the following relationship:

1.62C 2

R t

where

C = half defect length

R = radius of the tank

t = thickness of the tank wall

For the size of defect in the tank being analyzed here the stress

magnification factor was calculated to be 1.0002. This indicates that the

increase in local stress due to the defect causing bulging of the tank wall is

insignificant. Because the curvature of the tank could be ignored and no

significant stress magnification occurred due to bulging of the tank wall in

the vicinity of the defect, the stress intensity in the tank wall due to the

defect can be accurately calculated by modeling the problem as a defect in a flat plate.

137
8 . 3.2 Analysis of Part-Through- Surface Defect

The stress intensity for a part through wall defect was calculated based on a stress intensity formulation developed by Raju and Newman [8.7] for partthrough- surface defects in a flat plate. This calculation is based on the

following equation:

KI = aH

F

ttC

J Q

where

(?Yi — hoop stress

Q = defect shape factor = 1 + 1.464 (a/c)^°^,

C = defect half length

F = stress magnification factor

fW

where

t

Ml

M2

fW

w

a

plate thickness

1.13 - 0.1 (a/c)

J 7r/4

[ sec (It)]

1/2

plate width

defect depth

As shown in Table 8.2 the applied stress intensity calculated from stresses

created by the hydrostatic pressure was only 10 .

7

ksi^ in for the size of

defect found in the tank. When residual stresses were taken into account and

added to the hydrostatic pressure stresses the calculated applied stress

intensity was increased to 22 .

7

ksij in. Even if it is assumed that stresses

from all other undocumented sources could be as high as the yield strength of

the steel plate, the calculated applied stress intensity is increased only to

27 .

2

ksi4Tn. The critical stress intensity (Kq) for the first shell course

steel plate was 47.7 ksiJTn. For fracture initiation to occur the calculated

138
applied stress intensity must be larger than the critical stress • intensity

(Kc) for the steel plate. Therefore, it is concluded that even if stresses up

to the level of the yield strength of the steel plates were present at the

location of the defect, fracture would not be expected to have initiated from

a surface defect of this size. This analysis predicts that, at most, the

ligament of metal in the tank wall below the defect might be expected to break

and a through- the -wall defect would be formed that would result in leaking of

the tank without rupture of the tank occurring.

8.3.3 Analysis of Through- the -Wall Defect '

The second stage of the fracture analysis was to determine, that if the partthrough-wall defect became a through-the wall-defect as described above,

whether fracture would be initiated. The applied stress intensity for the

defect in the tank wall was calculated by using procedures for cracks in

hypothetical structures that are available in the technical literature.

The defect was first modeled as a center-cracked-plate stressed in tension.

This analysis gives results of acceptable accuracy when the curvature of the

tank can be ignored because the tank radius to thickness ratio (R/t) is large.

This condition was met for the problem analyzed here. Analytical solutions

for both the case of an infinite plate and for the case of a wide plate of

finite dimensions were used. The stress intensity factor for the center

cracked plate is calculated from the following relationship given in reference

[ 8 . 8 ]:

Ki = J ttC

where

C = defect half length

For this case, because the tank is very large relative to the size of the

defect, the stress intensity calculated using both the infinite plate and

finite plate analyses gave the same answers. The results for these

calculations of the stress intensity are shown in Table 8.2.

139
As shown in Table 8.2 the applied stress intensity calculated from stress

created by the hydrostatic pressure alone was 14 .

5

ksi4Tn for the size of

defect found in the tank. When residual stresses were taken into account and

added to -the hydrostatic pressure stresses the calculated applied stress

intensity was increased to 30 .

7

ksiJTn. Even if it is assumed that stresses

from all other undocumented sources could be as high as the yield strength of

the steel plate, the calculated applied stress intensity is increased only to

36 .

7

ksiJTn. The calculated applied stress intensity was the same for the

infinite plate analysis as for the finite width plate. The critical stress

intensity (K^) for the first shell course steel plate was 47.7 ksiJTn. For

fracture initiation to occur the calculated stress intensity must be larger

than the critical stress intensity (Kq) for the steel plate. Therefore, it is

concluded that even if stresses up to the level of the yield strength were

present at the location of the defect fracture would not be expected to have

initiated from a through-wall defect of this size.

The stress intensity factor for the through-wall defect was calculated by an

alternate method that models a defect in a thin walled, pressurized tank

[Reference 8.9]- This analysis models the actual defect in a more rigorous

manner and should give a more accurate value of the actual applied stress

intensity in the tank. The stress- intensity factor for a through-wall defect

in a thin walled pressurized structure is calculated from the following

equation:

Kj = Mp 4 7rC

where

Mp = stress magnification factor (defined previously)

As shown in Table 8.2 the applied stress intensity calculated by modeling the

defect as being in a thin walled pressurized structure was exactly the same as for modeling the defect as being in a flat plate. The applied stress

intensity from stress created by the hydrostatic pressure alone was 14 .

5

ksi>] in. When residual stresses were taken into account and added to the

hydrostatic pressure stresses the calculated applied stress intensity was

increased to 30 .

7

ksiJTn. Even if it is assumed that stresses from all other

undocumented sources could be as high as the yield strength of the steel

140
plate, the calculated applied stress intensity is increased only to 36,7

ksiJTn. The calculated stress intensity was the same for the infinite plate

analysis as for the finite width plate. The critical stress intensity (Kq) for

the first shell course steel plate was 47.7 ksi4Tn. For fracture initiation

to occur the calculated stress intensity must be larger than the critical

stress intensity (K^) for the steel plate. Therefore, it is concluded that

even if stresses up to the level of the yield strength were present at the

location of the defect fracture would not be expected to have initiated from a through-wall defect of this size.

Because the steel plates in this tank had a relatively low yield strength and

some plasticity is expected to be present in the vicinity of the defect, an

additional calculation of the applied stress intensity was made to account for

the size of the plastic zone at the edge of the defect. The analytical

solution for the stress intensity factor that takes into account the

plasticity is given by the following equation from reference [8.10]:

K

i -

8 —2 „ „ * mf"h —a | C in sec

7r 2 <*

-i 1/2

where — yield strength + tensile strength

a — flow stress = “ —a—

As can be seen from table 8.2, the applied stress intensity is increased

slightly when the plasticity correction is added to the analysis. However,

even at stress levels equal to the yield strength of the steel plate, the

applied stress intensity is still only 41 .

4

ksiJTn which is still well below

the critical stress intensity (47.7 ksiJTn) required to initiate fracture.

8 . 4 Discussion of the Fracture Analysis Results

For all cases considered in the fracture analysis, the calculated applied

stress intensity for the defect identified as the fracture origin was

substantially below the stress intensity required to initiate fracture and

cause rupture of the tank. This can readily be shown by the stress-flaw size

curve shown in Figure 8.2. For the measured toughness (47.7 ksiJTn) the

141
relationship between the nominal applied stress and the length of a throughthickness defect is calculated. Combinations of nominal stress and defect

size that lie above the line shown in Figure 8.1 are predicted to cause

fracture initiation and any combination of defect size and nominal stress that

lies below the curve is not expected to result in fracture initiation. As can

be seen from the points plotted in Figure 8.1, fracture is not expected to be

initiated from a through- thickness defect of the length of the defect found on

the fracture surface (0.742 inch.) even if the stresses in the vicinity of the

defect are equal to the yield strength of the steel plates. Therefore,

additional causes must be identified to ‘ determine why the tank ruptured. For

the fracture analysis to predict rupture of the tank, either the defect must

be significantly larger than the size identified on the fracture surface or

the fracture toughness of the steel plate must be lower than that measured in

this investigation.

The defect from which the fracture originated was clearly identified by

corrosion products on the fracture surface of the defect. Detailed

fractographic examination of the fracture surface in the vicinity of the

defect by scanning electron microscopy (SEM) that was done by Battelle did not

reveal any features that would indicate that the defect was significantly

larger that the size used in this analysis.

Therefore, it was concluded that the measured fracture toughness determined on

representative samples of the steel plates from the first shell course was not

representative of the fracture toughness of the metal in the vicinity of the

defect. This conclusion was supported by the microstructural examination of

the defect area, shown in Figure 7.26, and by the microhardness measurements

made near the defect. These results indicate that the metal in the vicinity

of the defect was altered by heating, probably during the initial fabrication

or during at the time of disassembly of the tank, and that a hard spot was

created around the defect. It is likely that the fracture toughness of this

hard spot was lower than the fracture toughness of the unaffected steel. This

lower fracture toughness in the vicinity of the defect could cause fracture to

initiate in this hardened area and to propagate without arresting through the

tougher steel plates.

142
Previous studies [8.11]. have shown that the stress required to initiate a fracture at a defect in steels of the type used in this tank that has been

altered by heating, such as by welding, is considerably reduced from the

stress required to initiate fracture in unaltered steel. Heating the metal

surrounding an existing defect, such as by a torch cut or welding over the

defect may reduce the stress to initiate fracture very significantly.

It was not possible to determine directly the fracture toughness of the

altered metal in the vicinity of the defect and therefore, it was not possible

to do a quantitative fracture analysis to predict the fracture stress and

critical defect size. To demonstrate that it was feasible for the stress

required to initiate a fracture from a defect in a region of heat altered

metal to be substantially reduced, the series of notched bend tests described

in section 7.2 was conducted. This series of tests was conducted to determine

the notch sensitivity of the steel at room temperature and 40° F when 1) a weld was deposited, then a notch machined into the weld, 2) a weld deposited

over a prexisting notch, and 3) no weld was present near the notch.

In this Investigation, the effect of the heat-affected zone on the steel is

determined from the load/displacement curve. It is seen in Figure 7.13 that

the notch mouth-opening displacement at the maximum load (point of fracture

initiation) for the heat affected (embrittled) specimen was 0.032- inch

compared with a value of 0.072 -inch for the notch mouth-opening displacement

for the normal steel plate. The area under the load-displacement curve

,

(Figure 7.13) which can be expressed in terms of energy, can also be taken as

a measure of the tendency to fracture. For example, two specimens that were

tested at 40° F with just a notch in the base metal had energy values at

maximum load equal to 21.4 and 21.7 foot-pounds. A similar specimen tested at

40° F, in which the metal had been embrittled by heating before cutting the

notch had an energy at maximum load of 9.0 foot-pounds and a specimen that was

heated after cutting the notch had an energy at maximum load of 9.7 footpounds . It was found that the energy to initiate fracture in the notched and

heat embrittled specimens at 40° F was less than one half the energy to

initiate fracture in the steel base metal that had not been embrittled by

heating

.

143
The notched -bend specimens did not meet all the requirements to permit

determination of the fracture toughness using the crack- tip-openingdiplacement (CTOD) method described in reference [8.11]. However, the results

of the notched bend tests indicate that fracture will initiate from a defect

in a heat affected and embrittled region at a lower stress than from a defect

in the unaltered base metal. If the notch-opening displacement at maximum

load or the energy derived from the area under the load- displacement curve for

the notch-bend tests is used to compare the fracture initiation of the heat

affected and unaltered specimens , it is estimated that the toughness of the

material in the heat affected region is less than 1/2 of the toughness of the

unaltered base metal. That is, the toughness of the heat affected region

surrounding the defect found on the fracture surface could be expected to be

less than 25 ksiJTn, that is, less than the stress intensity factor of 30.7

ksiJTn due to the combination of the hydrostatic and residual stresses in the

tank.

This supports the conclusion that although the defect found on the fracture

surface was not large enough to cause fracture initiation in the unaltered

base metal, the defect was large enough to initiate the fracture when it was

embedded in a hardened region of reduced toughness.

8 . 5 Evaluation of Fracture Propagation and Arrest

The objective of the fracture analysis described in section 8.4 was only to

predict if fracture initiation could occur with the stress, defect size, and

fracture toughness present in the tank. Once it has been predicted that

fracture can be initiated, it is necessary to determine if the fracture will

propagate to an unlimited extent and a complete rupture occur or if fracture

arrest will occur, resulting in only a leak. The limited scope of this

investigation did not permit a complete and quantitative fracture analysis to

be made to predict the conditions under which fracture propagation and arrest

could occur. A complete investigation of this type requires the determination

of dynamic fracture toughness (Kq) and arrest fracture toughness (Ka ) before a quantitative fracture analysis could be made to predict fracture propagation

and arrest.

144
A qualitative assessment of the tendency for fracture propagation and arrest

can be made from the results of the drop-weight nil-ductility- transition (NDT)

temperature tests. The test results on the steel plates from the first and

second shell course plates showed that the NDT temperature of these steel

plates was 60° F to 70° F or about 30° F to 40° F above the operating

temperature of the tank. The predictions made from the NDT temperature in

conjunction with the Crack-Arrest Curve on the Fracture Analysis Diagram [8.4]

show that fracture arrest would not be expected to occur in structures

operated at temperature below the NDT temperature of the steel plates, even

for stress levels well below the level of the yield strength. A crack arrest

temperature (CAT) is defined on the Fracture Analysis Diagram, Figure 8.1,

which shows that fracture propagation is expected to occur at and below the

NDT temperature even when the stress levels present are very low (<10 per cent

of the yield strength) . Fracture propagation is predicted at stress levels of

approximately one half the yield strength level at temperatures about 30° F

above the NDT temperature. Fracture arrest is not predicted for stress levels

equal to the yield strength of the steel until a temperature of approximately

50° F above the NDT temperature.

On the basis of the NDT temperature tests alone, unlimited fracture

propagation without fracture arrest is predicted to occur once fracture had

initiated in the tank. Fracture arrest would not be expected even at stress

levels due only to the hydrostatic pressure of the oil. This assessment

agrees with the observed complete rupture of this tank.

145
Table 8.1 - Summary of Values Used in the Fracture Analysis

Property Value

Tank Temperature 32° F

Defect Size

Length 0.742 inch

Depth 0.601 inch

Plate Thickness 0.875 inch

Hydrostatic Stress + 13.7 ksi

Residual Stress + 15.0 ksi

Yield Strength^ ) (ay) 34.0 ksi

Tensile Strength^ ) (a-p) 67.1 ksi

Flow Stress (crp) 50.6 ksi

Elastic Modulus 30,000 ksi

Fracture Toughness^)

Kc at 40° F 47.7 ks i4 in

J c at 40° F 75.7 in-lb/in^

(1) Average value from measured properties at first shell course

(2) *

F = + a

T

2

146
Table

8.2-

Summary

of

Calculated

Stress -

Intensity

Factors

032u0

UaC

01e5

0) 2

T3C

03

03 06Co

03

<U

CO

03 cQ

oo

oo

<U u

03

CL.

iJ

03 U*

<D TD3

5^

<D

U-iCCo

•o

co

03 CQ

O

00

ON JZu

oo •«-* — 5

0)

co

CO

0)>

-oVcco

•o

CD

CO

03 CQ

0)

COCO

0)>

T3

a;c

JZco

T3

0)

CO

03 CQ

147
LL

oO

SS3H1S 3ALLV13H

148

Figure

8.1

-

Fracture

Analysis

Diagram

(FAD)
in

c\j

o

c\i

LO

d

o

c

LU

N

(f)

o

LU

LL

LUQ

(!S>» SS3ldlS 1VNIIAI0N

149

Figure

8.2

-

Calculated

Stress

-

Defect

Size

Relationship

for

Through-Wall

Defect

9.0 CONFORMANCE WITH STANDARDS

9 . 1 General

The nationally recognized standard for above-ground, atmospheric, welded steel

tanks is the American Petroleum Institute Standard 650, Welded Steel Tanks for

Oil Storage, 1984 [9.1]. In this chapter a comparison will be made between

the practices employed in the reconstruction of Tank No. 1338 and the

provisions of API 650. Only those provisions which are relevant to the

structural performance of the tank will be addressed.

API 650 does not explicitly address the reconstruction of existing tanks. It

does, however, include provisions for the use of "materials on hand which are

not identified as complying with any listed specification" (see API 650,

Appendix N) . Thus API 650 can be used for reconstruction of a tank when

documentation is not available for the steel.

9 . 2 Materials

Section N.1.2 of API 650 states that each plate shall be subject to the

chemical check analysis and physical tests as required in the designated

specification, and the carbon and manganese content shall be determined in all

check analyses. It also requires tensile tests. In this instance no record

of a "designated specification" was found for review. Appendix N is,

therefore, interpreted to require that the material comply with one of the

specifications listed in Sections 2.2.2 through 2.2.5. The plate material

used in this tank, because of its age, was not fabricated in accordance with

any of the ASTM, CSA, or ISO standards listed. However, it complies with ASTM

Specification A 283-84, listed in Section 2.-2. 2. Mechanical properties

obtained for the plates are listed in Table 7.2. These can be compared with

Table 1 in ASTM A 283. The comparison indicates conformance with Grade D.

Carbon contents for the plates, listed in Table 7.1, ranged from 0.207% to

0.299% and averaged 0.238%. ASTM A 283 does not specify carbon content.

However, the steel should be considered weldable without pre-heating in

151
accordance with presently accepted criteria [9.2].

Section 2.2.9 of API 650 specifies toughness requirements. Section 2. 2. 9.

2

states that when used at design metal temperatures below the design metal

temperature specified for the material in Figure 2-2 of the standard "...the

material shall demonstrate adequate notch toughness at the design metal

temperature ..."( in use). In accordance with Table 2-3 of the standard, the

material of the plates used does not fit any of the groups listed, because it

was not killed or semi-killed. However, for the pupose of assessing

conformance with the toughness requirements the material is placed in Group I (A 283 Grade C is listed in Group I, but A 283 Grade D is not listed.

However, as noted in Section 9.3 of this report, the design stress does not

exceed the working stress allowed for A 283 Grade G) . For a plate thickness

of 7/8 inches, the specified design metal temperature for Group I materials is

42° F. However, in accordance with Figure 2-1 and Section 2. 2. 9. 3 of API 650,

the design metal temperature for Pittsburgh , Pennsylvania is +5° F. Thus

adequate notch toughness must be demonstrated by testing. Table 2-2 of API

650 specifies a minimum absorbed energy requirement for the Charpy V-notch

Test (ASTH A 370). For Group I material an absorbed energy of 15 ft- lb is

specified for longitudinal orientation (parallel to the direction of rolling)

.

The absorbed energy measured for specimens taken from the plates (Table 7.4)

was 6 ft- lb or less for temperatures below 40° F. Thus the toughness of the

material is seriously deficient in respect to the standard.

In addition to the base metal, the plates also included old welds which were

not removed prior to the reconstruction of Tank No. 1338. Thus these old

welds were part of the material and therefore must comply with the -materials

provisions of the standard. In accordance with a radiography report prepared

by Progress Services, Inc. of Monroeville, PA and submitted to Ashland Oil

Company on November 25, 1986, the old welds were not of acceptable quality per

Paragraph UW-51(b) of the ASME Boiler and Pressure Vessel Code, Section VIII,

Division 1 [9.3]. Figure 9.1 shows defects in old and new welds.

In summary, a comparison with of the test results of the plate material with

the provisions in Section 2 and Appendix N of API 650 indicates that the plate

material does not comply with the toughness requirements of API 650 and is

152
seriously deficient in this respect. In accordance with the standard, the

material used was inappropriate for the climatic conditions of Pittsburgh,

Pennsylvania.

9 . 3 Shell Design

The original plans and design calculations for Tank #1338, if they exist, were

not available for review. Also, no design calculations were found for the

reconstructed tank.- The inspection report for the Whiskey Island Tank #16

indicates that the wall thickness for all shell courses was satisfactory.

Presumably this determination was based on calculations made for the measured

wall thicknesses, specific gravity of the product, corrosion allowance,

allowable stress of the steel, and design metal temperature. These data,

while required by API 650 to be supplied by the purchaser, were not available.

One can compute the required shell thickness using the "one- foot method" (API

650, Section 3. 6. 3.1) in which the design thickness, t, is given by,

t =

2 - 6 D

f1} G

+ C

.

A.

d

where

,

H = height from the bottom of the shell course to the top of the

shell

D = tank diameter = 120 ft

G = specific gravity =0.87

= design allowable stress = 20,000 psi

C„A. = Corrosion allowance = 0

These values are reported in Table 9.1 for a specific gravity of 0.87 and a design stress of 20,000 psi. The wall thicknesses measured by NBS are also

shown in Table 9.1. By subtracting the required thickness from the measured

thickness, one obtains the available corrosion allowance. It is seen that,

in all cases, the shell plate thickness provided either meets or exceeds the

required thickness, but the available corrosion allowance is only

approximately 1/16 in.

153
9.4 Inspection

Section 6 of API 650 requires that butt-welded joints of the shell plates,

both vertical and horizontal, be spot radiographed. The spot radiographs are

to be taken as the welding progresses (Section 6. 1.2. 7) and are intended to

determine if a welder is making good quality welds. The acceptance criteria

for welds examined by radiography are contained in Paragraph UW-51(b) of the

ASHE Boiler and Pressure Vessel Code, Section VIII, Division 1 [9.3] and are

adopted by reference (API 650, Section 6.1.5).

The number and location of spot radiographs are described in Section 6. 1.2.1

for vertical joints and Section 5. 1.2.

2

for horizontal joints. In the lowest

shell course, for example, API 650 requires two spot radiographs in each

vertical joint and a radiograph at all junctions of vertical and horizontal

joints. Tank #1338 was radiographed on one occasion, after completion of

construction. A total of 39 radiographs were taken; 26 in a vertical joint, 9

at the intersection of a horizontal and vertical joint, and 4 in a horizontal

joint. The radiographer reported that, out of the 39 radiographs, 22 did not

meet the - acceptance criteria. There is no record that any repairs were made.

Several provisions of API 650 regarding radiographic inspection were not met:

fewer radiographs were taken than were required; the radiographs were not

taken during construction, but rather after the tank was completed; and,

although 22 radiographs failed to meet the acceptance criteria, no repairs

were made. In the letter of transmittal which accompanied the radiography

report, the following comment was made:

"It is my opinion that the areas which are rejectable are in the old

weld and not in the weld recently done."

Had API 650 been followed, there would have been a radiograph at the location

of the flaw that initiated the fracture of the shell that might have revealed

the presence of the flaw.

154
9.5 Testing

Requirements for testing are included in Section 5 of API 650 which deals with

erection. Section 5.3.6 deals with testing of the tank shell. Part 1 of this

section requires that, if water is available, the tank shall be filled with

water to the maximum design level or to a level 2 inches above the weld

connection of the roof plate. Part 2 of the section permits an alternate

method of testing for joint leakage if sufficient water is not available.

The availability of sufficient water at the Floreffe site is not questioned.

Thus, in accordance with API 650, the tank should have been filled with 46 ft

of water and carefully inspected during the filling, before external oil

piping was connected.

It was reported that the tank was filled with only 5 ft of water rather than

46 ft. Thus the erection procedures specified by API were not followed. Had

the test been carried out as specified, the tank might have failed during the

water test preventing the severe environmental damage that occured. However,

abrupt failure of a water-filled tank would itself be hazardous. It is also

possible that a water test at a temperature higher than that at the time of

the failure would have resulted in no leak at all or a slow leak, rather than

a rapid crack propagation.

9 . 6 Foundation

Section 3.3.1 of API 650 specifies -that site selection and foundation design

and construction "shall be given careful consideration as outlined in

Appendix B" . This statement implies that Appendix B is a recommendation

rather than a provision.

Section B.2.1 deals with subsurface exploration and settlement control. No

data were available to NBS to indicate the nature and extent of subsurface

exploration prior to construction of the tank. The statement in the standard

on settlement is non-specific. However, an analysis of the foundation

indicates that the maximum anticipated settlements would be within tolerable

155
limits

.

Sections B.2.2 and B.2.3 do not apply because the underlying soil is

considered competent to support the tank without special remedial measures.

Section B.2.4 deals with fill materials, but the degree of required compaction

is not specified. The foundation analysis indicates that the compaction

specified in the foundation construction contract was adequate and that the

compacted fill was constructed in accordance with the specifications.

Section B.3 deals with the grading of the foundation surface supporting the

tank bottom and the subgrade supporting the bottom plate. The tank

foundation complies with these provisions.

Section B.4 deals with foundation design. The foundation analysis indicates

that the design criteria in Section B.4.1 are satisfied. The recommended

design detail for a tank which is not supported by a concrete ring foundation

is shown in Figure B-2 of the Standard. The design of the tank foundation,

shown in' Figure 4.1 deviates from this detail. However, the gravel pad as a whole was 1 to 1 1/2 ft thick thereby eliminating the need for the crushed

stone ringwall shown in Figure B-2. In addition, a geotextile sheet was

inserted between the crushed-stone pad and the compacted fill, which could

inhibit radial spreading of the crushed-stone pad and erosion caused by

potential leaks.

In summary, on the basis of construction drawings and records and subsurface

data supplied by the Pennsylvania Task Force and the foundation analysis

presented in Section 4, it is concluded that the tank foundation complies with

the provisions and recommendations in Section 3.3.1 and in Appendix B of API

Standard 650.

156
Table 9.1 - Design Shell Plate Thicknesses

Plate Thickness (in) Computed

Corrosion

Measured by NBS Allowance

(nearest 1/16")

Design Minimum

(next 1/16")

Shell Course 1 7/8 11/16 3/16

Shell Course 2 11/16 9/16 1/8

Shell Course 3 1/2 7/16 1/16

Shell Course 4 3/8 5/16 1/16

Shell Course 5 1/4 1/4 -

Shell Course 6 3/16 1/8 1/16

157
Figure 9.1 - Defects in a Horizontal

Shell Courses

Weld between the First and Second

158
10 . CONCLUSIONS

10.1 Cause of Tank Collapse

The failure of the tank was caused by a brittle fracture that initiated at a flaw in the tank shell about 8 ft above the base and just below the horizontal

weld between the first and second courses of steel. The fracture propagated

to complete rupture of the tank shell because the steel lacked sufficient

toughness to arrest a propagating fracture at the temperature and stresses

existing in the tank shell.

Failure was reported to have occurred as the tank was being filled to capacity

for the first time at the Floreffe site. The temperature of the shell was

about 32° F, and the calculated stress at the initiating flaw about 13.7 ksi.

These conditions are fully consistent with initiation and complete propagation

of brittle fracture considering the flaw size, embrittlement of adjacent

steel, presence of residual stresses, and the low fracture toughness of the

shell steel at 32° F. The shell steel did not meet the fracture toughness

requirements of API 650.

The flaw at which the fracture initiated was determined by Battelle to be

present prior to the reconstruction of the tank at the Floreffe site.

Deposits on the surface of the flaw indicate it was formed during a cutting

operation.

Although the weld quality did not conform to API 650, the weld quality was

not a contributing factor in the initiation or propagation of this fracture.

However, welding adjacent to the flaw contributed to embrittlement of the

metal

.

Foundation conditions were investigated to determine if instability or

irregular settlement of the foundation might have increased stresses in the

shell. No evidence was found of foundation instability or excessive

settlement

.

159
10.2 Lessons Learned

This study focused on the tank that collapsed at the Floreffe Terminal on

January 2, 1988, and on the standards applicable to its dismantling and

reconstruction. In spite of this limited scope, lessons of general

significance have been learned from this failure, and relearned from earlier,

similar failures.

The catastrophic failure of the Ashland tank resulted from a flaw that might

have been, but was not, detected in tests for welding quality, and the use of

steel that did not possess sufficient fracture toughness to arrest a fracture

that initiated at a localized flaw. While there are many reasons for control

of weld quality, it is not feasible to guarantee the absence of localized

defects by control of fabrication and detailed inspection for defects during

fabrication. Neither is it feasible to guarantee the absence of localized

defects during service by recurrent inspections. Therefore, it is advisable

to use steel of adequate toughness to provide "leak before break" fail-safe

behavior. If a crack is initiated, due to a flaw or local damage in service,

in a tank with steel of adequate fracture toughness, the crack will arrest,

the tank will leak, normal spill-control measures will confine the spread of

contents, and the tank can be repaired all without catastrophic consequences.

Available standards should be reviewed to identify those calling for

sufficient fracture toughness to prevent catastrophic brittle fracture at

design stresses and temperatures. This study found the provisions of API 650

(1984) to be sufficient for the structural steels in its Materials Group I.

Compliance with sufficient fracture toughness provisions should be required

for steel to be used in construction or reconstruction of tanks, or for

changes in service conditions

.

There is concern for the safety of existing tanks whose catastrophic failure

would cause unacceptable human, environmental or economic losses. If

documentation exists to show that the steel of an existing tank meets

sufficient fracture toughness provisions, the concern will be alleviated. If

documentation shows inadequate fracture toughness, remedial actions should be

taken. Possible remedial actions include conversion to adequately high

160
temperature service, retirement, or installation of crack arresters.

Test and assessment protocols should be developed for assessing the fracture

safety of potentially hazardous existing tanks that lack adequate

documentation for definition of the fracture toughness of their steel.

161
REFERENCES

2.1 ASME Boiler and Pressure Vessel Code , Section VIII - Rules for

Construction of Pressure Vessels, Division 1, 1986 Edition, The

American Society of Mechanical Engineers, New York, 1986.

4.1 Duncan, M.J. and D'Orazio, T.B., "Stability of Steel Oil Storage

Tanks", Journal of Geotechnical Engineering, ASCE, Vol. 110, No. 9,

September 1984.

4.2 Bjerrum, L. and Overland, A., "Foundation Failure of an Oil Tank in

Fredrikstad, Norway", Proceedings of the 4-th Intemalional Conference

on Soil Mechanics and Foundation Engineering, Vol. 1, pp . 287-290,

1957.

4.3 D'Orazio, T.B. and Duncan, M.J., "Differential Settlements in Steel

Tanks", Journal of Geotechnical Engineering, ASCE, Vol. 113, No. 9,

September 1987.

4.4 Langeveld, J.M.,"The Design of Large Storage Tanks for Crude Oil and

Natural Gas", Proceedings of the Annual Meeting of International

Institute of Welding, 1974.

4.5 Malik, Z. et al . , "Qvalization of Cylindrical Tanks as a Result of

Foundation Settlements", Journal of Strain Analysis, Vol. 12, No. 4,

April 1977.

4.6 Penman, A. D. M. , "Soil-Structure Interaction and Deformation Problems

with Large Oil Tanks", Proceedings of the International Symposium on

Soil-Structure Interaction, Vol. 1, Roorkee , India, January 1977.

4.7 Marr, W.A. , Ramos, A.J., and Lambe , T.W.

,

"Criteria for Settlement of

Tanks", Journal of the Geotechnical Engineering Division, ASCE, Vol.

108, No. GT8 , August 1982.

5.1 PATRAN Plus User Manual, PDA Engineering, Costa Mesa, California,

1987.

5.2 Stress Analysis Module Primer, PDA Engineering, Costa Mesa,

California, 1985.

6.1 Jakob, M. , Heat Transfer, vol. 1, John Wiley & Sons, Inc., New York,

1949.

6.2 Spiers, H.M.

,

Technical Data on Fuel, Fifth Edition, The British

National Committee, World Power Conference, London, 1955.

6.3 McAdams, W.H., Heat Transmission, 3rd ed. , McGraw-Hill, New York,

1954.

6.4 Holman, J.P., Heat Transfer , 5th ed.

,

McGraw-Hill, New York, 1981.

162
6.5 Wind Loading and Wind- Induced Structural Response

,

ASCE, New York,

1987.

7.1 ASTM E-415-85, Standard Method for Optical Emission Vacuum

Spectrographic Analysis of Carbon and Low Alloy Steel, published by

the American Society for Testing and Materials, 1916 Race St.

Philadelphia, PA. U.S.A.

7.2 ASTM A10-39, Standard Specification for Mild Steel Plates,

published by the American Society for Testing and Materials, 1916 Race

St. Philadelphia, PA. U.S.A.

7.3 ASTM A283-84, Standard Specification for Low and Intermediate Tensile

Strength Carbon Steel Plates, Shapes and Bars, published by the

American Society for Testing and Materials, 1916 Race St.

Philadelphia, PA. U.S.A.

7.4 Welded Steel Tanks for Oil Storage, API Standard 650, Seventh Edition,

November 1980, Revision 1, February 1984, American Petroleum

Institute, Washington, D.C.

7.5 Barr, W. and Honeyman, A.J.K., "Some Factors Affecting Notched-Bar

Impact Properties of Mild Steel," J. Iron & Steel Inst., Vol.157,

pp. 242-246 (1947).

7.6 ASTM E-8-86, Standard Methods of Tension Testing of Metallic

Materials, published by the American Society for Testing and

Materials, 1916 Race St. Philadelphia, PA. U.S.A.

7.7 ASTM E-23-86, Standard Methods for Notched Bar Impact Testing of

Metallic Materials, published by the American Society for Testing and

Materials, 1916 Race St. Philadelphia, PA. U.S.A.

7.8 Johnson, H.H. and Stoudt, R.D., "Comparison and Analysis of Notch

Toughness Tests for Steels in Welded Structures," ppl-28. WRC Bulletin

62, July 1960.

7.9 ASTM E-208-85, Standard Method for Conducting Drop-Weight Test to

Determine Nil-Ductility Transition Temperature for Ferritic Steels,

published by the American Society for Testing and Materials, 1916 Race

St. Philadelphia, PA. U.S.A.

7.10 ASTM E-813-87, Standard Test Method for Jjc> A Measure of Fracture

Toughness, published by the American Society for Testing and

Materials, 1916 Race St. Philadelphia, PA. U.S.A.

7.11 ASTM E-399-83, Standard Test Method for Plane-Stain Fracture Toughness

of Metallic Materials, published by the American Society for Testing

and Materials, 1916 Race St. Philadelphia, PA. U.S.A.

163
7.12 Rosenfield, A.R. , and Shetty, D.K., "Cleavage Fracture of Steel in the

Ductile-Brittle Transition Region," Elastic-Plastic Fracture Test

Methods: The User's Experience, ASTM STP 856, E.T. Wessel and F.J.

Loss eds . , American Society for Testing and Materials, 1985, pp.196-

209.

7.13 Barsom, J.M. and Rolfe, S.T., Fracture and Fatigue Control in

Structures, Second Edition, Prentice-Hall, Englewood Cliffs, New

Jersey, 1987, pp.569.

7.14 Private Communication- -R.Mesloh, Battelle, April, 1988.

7.15 Micrographs supplied by R.Mesloh, Battelle.

7.16 Letter from Kirkpatrick & Lockhart Attorneys at Law, March 14, 1988,

reporting results of tests by Battelle.

7.17 Masubuchi, K. Analysis of Welded Structures, Pergamon Press, Oxford,

England, pp. 209.

8.1 Pellini, W.S., Guidelines for Fracture-Safe and Fatigue-Reliable

Design of Steel Structures

,

The Welding Institute, Cambridge, England,

1983, pp. 12-18.

8.2 Williams, M.L. and Ellinger, G.A., "Investigation of Structural

Failures of Welded Ships," Welding Journal, Vol.32 (10), Research

Supplement, pp . 498s- - 528s , 1953.

8.3 Gross, J.H. and Stout, R.D., "Ductility and Energy Relations in Charpy

Tests of Structural Steels", Welding Journal, Vol.37 (4), Research

Supplement, pp,151s--159s, (1958)

8.4 Pellini, W.S. and Puzak, P.P., "Fracture Analysis Diagram Procedures

for the Fracture-Safe Engineering Design of Steel Structures," Welding

Research Council Bulletin No. 88, May 1963.

8.5 Barsom, J.M. and Rolfe, S.T., Fracture and Fatigue Control in

Structures, Second Edition, Prentice-Hall, Englewood Cliffs, New

Jersey, 1987, pp.109.

8.6 Folias, E.S., "A Finite Line Crack in a Pressured Cylindrical Shell,"

International J. of Fracture Mechanics, Vol. 1, (1965), pp. 104-113.

8.7 Raju, I.S. and Newman J.C., "Analyses of Surface Cracks in Finite

Plates Under Tension or Bending Loads," NASA TP-1578, National

Aeronautics and Space Administration, Dec. 1979.

8.8 Broek, D. Elementary Engineering Fracture Mechanics

,

Martinus Nordhoff

Publ., The Hague, Holland, (1982), pp.12.

8.9 ibid, pp.379

164
8.10 Hahn, G.T., Sarrente.M. and Rosenfield, A.R., "Criteria for Crack

Extension in Cylindrical Pressure Vessels," International J. Fracture

Mechanics, Vol.5 (1969),- pp. 187-210.

8.11 Hall, W.J., Nordell, W.J. and Munse, W.H. , "Studies of Welding

Procedures", Welding Journal, Research Supplement Vol. 41 (11),

(1962).

8.12 ASTM Draft Standard "Test Method for Crack-Tip Opening Displacement

(CTOD) Fracture Toughness Measurements, August (1987).

9.1 Welded Steel Tanks for Oil Storage , API Standard 650, Seventh Edition,

November 1980, Revision 1, Febryary 1984, American Petroleum

Institute, Washington, D.C., 1984.

9.2 Ricker, D.T.

,

"Field Welding of Existing Structures," Engineering

Journal , AISC, Vol. 25, No . 1 , 1988.

9.3 ASME Boiler and Pressure Vessel Code, Section VIII - Rules for

Construction of Pressure Vessels, Division 1, 1986 Edition, The

American Society of Mechanical Engineers, New York, 1986.

165
APPENDIX

SUBSURFACE EXPLORATION DATA

Boring Logs

Oedometer Tests from Boring B-4 and B-5

Unconsolidated Undrained Compression Tests

166
ENGINEERS FIELD BORING LOG

PROJECT NO.

TEST PIT NO.

**-*><<.

SHEET QF 3

SORING NO. 'i- )

PROJECT LOCATION: LR. SEC. ;CTY.,

STATION OFFSET ELEVATION ZSP. z. p.'TT ' li£ -Yi/ti EQUIPMENT used J.

ENGINEER

C'-^ w •*<- ENGINEERS REP.

,‘ T,r>. a ? , L. -A L^» Y7 -- •

ORIUJNG METHOOS AH.A u-./

7

V3.Q

WA'ER LEVEL DEPTH: FEET 9 HRS.; FEET 9 HRS.;

UwIs

O

Uiti

2 2 z 3

a®

i s

v)

BLOWS

/Q. 5

FT.

ON

SOIL

SAMPLER

RECOVERY (IN) H /

> /

§ /*?

8/ a

Lu/ a

ad ai

POCKET PENETROMETER (TSF) DESCRIPTION REMARKS

11

1

11 \2. r Lh»mt ^ Cum CCc^ j

'3o<Xn^<i O —

^T- OF

S" *x<.«

„

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0-2.

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l_. _ x-x

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OK -* i A \.Z

Lxiwr i n > t Oiw-ry Ciu^ j S-L \% G.2.

i MSvl,-

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i*t T«<*X<VT

o-o "7 N % —

•H

:J s NOTES:

STRATIFICATION unes represfnt the approximate bounory between

SOIL OR ROCK TYPES AT THE BOREHOLE LOCATION ONLY.
ENGINEERS FIELD BORING LOG SHEET 7^ 0F 3

BORING NO. I

PROJECT NO PROJECT LOCATION: LR. „ „,;SEC< ; CTY

TEST PIT NO STATION OFFSET— ELEVATION

PATT \k£- EQUIPMENT USED -—

ENGINEER &*+-»*-**-*r ENGINEERS REP.

DRILLING METHODS 1 vz' ^ ^ ' t.^.

WATER LEVEL DEPTH: FEET 0 ___ HRS.; FEET 0 HRS.;

NOTES:

, STRATIFICATION UNES REPRESENT THE APPROXIMATE 8OUN0RY BETWEEN

SOIL OR ROCK TYPES AT THE BOREHOLE LOCATION ONLY.
ENGINEERS FIELD BORING LOG SHEET_1_ QF 3

BORING NO. 3 M .

' 5^)0 < C5U.^p<»€.

PROJECT NO PROJECT LOCATION: LR. ;SEC ;OTY _

test PIT NO STATION OFFSET ELEVATION iJLl

DATE \l2£i EQUIPMENT USED -2-1

6*^<=.rr ENGINEERS REP.

ORILUNG METHOOS

WAVER LEVEL DEPTH: 2. FEET O O HRS.; FEET 0 HRS.;

NOTES:

STRATIFICATION UNES REPRESENT THE APPROXIMATE BOUNDRY BETWEEN

SOIL OR ROCK TYPES AT THE BOREHOLE LOCATION ONLY.
ENGINEERS FIELD 30RING LOG . SHEET OF A

80RING NO. G-z -

PROJECT NO PROJECT LOCATION: LR. .;SEC — ;CTY

TEST PIT NO STATION OFFSET- ELEVATION -ZLLJ_

D att \/zZ-uAa EQUIPMENT USED Z Co^c.

r\ir;;NF?R GHo>jiuT^-e^ pnginffps REP.

DRILLING METHOOS

^ - WAVER LEVEL OEPTH: FEET 3 HRS.; FEET 3 HRS.;

NOTES:

STRATIFICATION UNES REPRESENT THE APPROXIMATE BOUNORY BETV/EEN

SOIL OR ROCK TYPES AT THE BOREHOLE LOCATION ONLY.
ENGINEERS FIELD SORING LOG SHEET CF 3

. SORING NO. *.z ,

PROJECT NO PROJECT LOCATION: LR. ; SEC ;CTY._

TEST PIT NO STATON OFFSET ELEVATION L2H£_

Patt \ h*-z* A% EQUIPMENT USED ^ >?. * .s^

^

c-\ir::MF-R FL.^ ENGINEERS REP. z,

_ ORIUJNG METHOOS - NASA

WA:ER LEVEL OEPTH: FEET « HRS.; FEET « HRS.;
ENGINEERS FicLQ BORING LOG SHEET _2_ OF 3 — SORING NO. ?.-*. _

PROJECT NO PROJECT LOCATION: LR. — ,;SECc ;CTY __

TEST PIT NO STATION OFFSET—— ELEVATION -Z^SLC

C ATg Ai equipment used

Tr'NF” PKIf^iNFr RR REP. —

DRILLING METHODS -W^_ _ -

— • -

NOTES:

STRATIFICATION UNES REPRESENT THE APPROXIMATE 9OUN0RY BETWEEN

SOIL OR ROCK TYPES AT THE BOREHOLE LOCATION ONLY.
ENGINEERS FiEL.0 SORING LOG

PROJECT NO. _____ PROJECT LOCATION: LR. ;SEC.

TEST PIT NO. STATON OFFSET

SHEET _i CF I SORING NO. 3> - T>

;CTY.,

P.atp t/l« I** EQUIPMENT USED

ELEVATON

^

ENGINEER

ORILLJNG METHODS V^»V

WA:ER LEVEL OEPTH:

_=2±_ ENGINEERS REP.

' -- c--V~r—

Vvrk, >jr:- .Vi» FEET « HRS.; FEET O HRS.;

Uiw ££

a

SAMPLE

NO.

/TYPE

OR

RUN

NO. d

*3

it

RECOVERY

!

(IN) H /

> / az L*

/**

a / w

a/ a

lu/ a

od cc

POCKET PENETROMETER (TSF) DESCRIPTION

I r Z

a

V \

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5 L ^ ..Nt. ^

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-7 7

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7a Z

«-w

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RQUARKS

w«u S’.Z.' T^r» 4-K4

'V

i \0 *"dSC tp’f \).jnr^

Uj<VCVt. o _ -2.-3

H

_J

NOTES:

STRATRCATON UNES REPRESENT THE APPROXIMATE BOUNORY BETWEEN

SOIL OR ROCK TYPES AT THE BOREHOLE LOCATON ONLY.
ENGINEERS HELD BORING LOG SHEETS. ar_Z

BORING NO.

PROJECT NO PROJECT LOCATION: LR. ;SEC ;CTY..

TEST PIT NO STATION OFFSET ELEVATION

DATE EQUIPMENT USED -

ENGINEER ENGINEERS REP. —

DRILLING METHODS - - WA“;ER LEVEL OEPTH: FEET « HRS.; FEET 9 HRS.;

NOTES:

STRATIFICATION UNES REPRESENT THE APPROXIMATE BOUNORY BETWEEN

SOIL DR ROCK TYPES AT THE BOREHOLE LOCATION ONLY.
PROJECT NO.

test pit NO.

ENGINEERS FIELD BORING LOG

V^M

PROJECT LOCATION: LR.

STATION OFFSET.

SEC.

SHEET L_ OF

BORING NO. ~

. ;CTY..

ELEVATION T2 7 2. S~

D.^rr ^/x_ EQUIPMENT USED

ENGINEER ENGINEERS REP. <xwiw^

ORILUNG METHOOS S>gw , 2

WA:ER LEVEL DEPTH: FEET 0 HRS.; FEET 0 HRS.;

NOTES:

STRATIFICATION LINES REPRESENT THE APPROXIMATE BOUNORY BETWEEN

SOIL OR RQCX TYPES AT THE BOREHOLE LOCATION ONLY.
ENGINEERS FIELD BORING LOG

'F'vv-.rt. 'FtS"—«"i-

- PROJECT NO. PROJECT LOCATION: L.R. : SEC.

TEST PIT NO STATION OFFSET

Patt . EQUIPMENT USED

ENGINEER ENGINEERS PEP. - DRILLING METHODS

WAilR LEVEL DEPTH: FEET 0 HRS.; _

SHEETS. OF ?

SORING NO.

, ;CTY..

ELEVATION 77Z. iT*

FEET 0 HRS.;

QTES:

RATIFICATION UNES REPRESENT THE APPROXIMATE BOUNORY BETWEEN

"IL DR ROCK TYPES AT THE BOREHOLE LOCATION ONLY.
Void

Ratio,

Type of Specimen Shelby Before Test After Test

Dia 2.50 m Ht 0.75 in Water Content w8 1 18.9 w* 20.1

Compression Index C. 0.16 Void Ratio ®o ! 0.592 e, 0.553

Classification CL Saturation S

I 87.7 S, 1Q0.Q

w, 48.6 l 0 25.5 Proiect ASHLAND TANK FAILURE

w0 23.1 ; u -o.i6 Bonng No B-5 Sample No U-l

Remarks Wet Unit Wt. » 128.2 Depth 6 .

0

'-7 .6’ Date 4/88

PCF . Specimen required a seating load of 0.20 TSF

5 . cr

4 .

3 .

2 .

1 .

to prevent swelling

.

GANNETT FLEMING GEOTECHNICAL LABORATORY

CONSOLIDATION TEST REPORT

Coef f

’of

C^dnsol°

C

u (in 2

/

min )

x

10~
Void

Ratio

Pressure, p(tsf)

Type of Specimen Shelby Before Test After Test

Dia 2. SO in HT 0.7S in Water Content 23.6 w, 17.9

Compression Index Ce 0 © 17 Void Ratio 0 . 668 ef 0.487

Classification Cl Saturation So 96.0 s, 100

-25.2 >e 7.5 Project ASHLAND TANK FAILURE

Wg 17.7 U 0.79 Bering No B~5 Sample No 0-2

Remarks Wet density^ 125.8pcf Depth 26=27.9' Date 4/88

GANNETT FLEMING GEOTECHNICAL LABORATORY

CONSOLIDATION TEST REPORT

Caeftf.

of

Consofi.
.600

Pressure, p(tsf)

Type of Specimen Shelby Before Test After Test

Dia 2.50 m Hr 0.75 in Water Content w. 17.8 w, 18.8

Compression index Ce ! 0.14 j Void Ratio e„ | 0.566 e. 0.516

Classification CL | Saturation i S0 f 86.7 ‘ S. 100.0

w’ 48.5 1 l 0 23.2 Project • ASHLAND TANK FAILURE

25.3 1 u -0.32 Bonng No 3.4 Sample No

Remarks Depth 4.0' -6.0' Date 4/88

Wet Unit Wt. = 129.2 PCF

GANNETT FLEMING GEOTECHNICAL LABORATORY

CONSOLIDATION TEST REPORT

Coeff

of

Consol

,

C

u (in 1

/

mm )
Void

Ratio

Type of Specimen Shelby Before Test After Test

Oia 2.50 in Hr 0.75 in Water Content we 23.1 Wf 14.8

Compression Index Ce 0.13 Void Ratio 00 0.646 0f 0.423

Classification ml Saturation So 97.2 s, 95.3

w, .22.9 ! e 5.3 Project ASHLAND TANK FAILURE

wb 17.6 U 1.2 Boring No B-4 Sample No U-2

Remarks Wet density^ 127 . Opcf Depth 2 6 -2 8

• Date 4/88

GANNETT FLEMING GEOTECHNICAL LABORATORY

CONSOLIDATION TEST REPORT

Coeff.

off

Coosa!.,

C

u (in. 2

/

min

.

)

X

10
•

Deviator

Stress

(ksf)

Shear

Stress

(kst)

Normal Stress (ksf)

Axial Strain (%)

Remarks:

Specimen No 1 2 3 Oiameter (in) Do 2.82 2.82 2.82

Height (in) H0

in in cr> 5.60 5.59

Initial

Water Content (%) 16.8 16.2 16.5

Dry Density (pcf) ^do 111.9 112.3 ! 111.8

Void Ratio ®o 0.534 0.528 0.535

Saturation (%) So 86.6 84.4 84.8

Before

Test

Water Content (%) wc — -— —

Dry Density (pcf) Toe - —

Void Ratio — —

Saturation (%) So i

Back Pressure (ksf) MS 0 0 0

Total Minor Princ Stress (ksf) <T3 1.27 2.71 4.15

Maximum Deviator Stress (ksf) P/A 4.16 5 . 34 i 6.02

Time Max Deviator Stress tf 21 21 21

Total Major Princ Stress (ksf) <S 1 5.43 8.05 10.17 :

Pore Pressure at Max

Deviator Stress (ksf) M — —

• —

Eff Minor Princ Stress (ksf) _ _

Eff Major Princ Stress (ksf)

1 <n — — —

Project ASHLAND TANK FAILURE Boring No D-l ,2 , 3 Sample No

‘H

Type of Test: Unconsolidated-Undrained Sample Type: Remolded j c = 1.3 ksf * = 14.5 c' = ksf (b

' =

c GANNETT FLEMING GEOT

TRIAX1AL COMPRES

ECHNICAL LABORATORY

SION TEST REPORT
Devialor

Stress

(ksl)

and

Pore

Pressure

(ksl)

Shear

Stress

fksf)

Axiai Strain (%)

Remarks:

Specimen No 1 2 3 Diameter (in) D0 2.83 2.83 2.83

Height (in) Ho 5.60 5.60 5.60

Water Content (%) Wo 14.3 14.3 13.7

.2 Dry Density (pcf) 116.4 116.2 117.4

Void Ratio 0.474 0.477 0.463

Saturation (%) So 82.9 82.4 81.3

Water Content (%) wc - -

u?

iS H— Dry Density (pcf) Toe - - S

o

Void Ratio ec - -

"s

CO Saturation (%) Se - -

Back Pressure (ksf) M0 - -

Total Minor Princ Stress (ksf) <73 1.27 2.71 4.15 - Maximum Deviator Stress (ksf) P/A 8.53 10.85 12.89

Time Max Deviator Stress tf 10 13 18

Total Major Princ Stress (ksf) <T1 9.80 13.56 17.04

Pore Pressure at Max

Deviator Stress (ksf) M - - -

Eft Minor Princ Stress (ksf) <T 3 - - -

Eft Major Princ Stress (ksf)

! <ri - - -

Project ASHLAND TANK FAILURE Boring No d-1 ,2,3 Sample No

Type of Test: Unconsolidated-Undrained Sample Type: Remolded

c =

2 ° 00

ksf W O e

CM tl c' * “ ksf <y = GANNIETT FLEMING GEO*

TRIAXIAL COMPRES

TECHNICAL LABOR*

SSION TEST REPORT

occ>
Deviator

Stress

(ksf)

Shear

Stress

(ksf)

0 2 4 6 8 10 12

Normal Stress (ksf)

0 5 10 15 20

Axial Strain (%)

Remarks: Max. Deviator Stress

taken at 15% strain ; Specimens

1 & 2 taken from S-4, and 3 & 4

Specimen No 1 2 3 Diameter (in) Do 2.83 2.83 2 . 85 i Height (in) Ho 5.60 5.59 5.60i

Initial

Water Content (%) *0 19.3 19.4 20.01

' ' Dry Density (pcf) T(3o 106.2 111.4 104.31

Void Ratio ®0 0.617 0.540 0.6451

Saturation (%) So 86.1 98.7 85.1

Before

Test

Water Content (%) —

Dry Density (pcf) — —

Void Ratio —

Saturation (%) Sc _

Back Pressure (ksf) M 8 11.52 11.61 11.52!

Total Minor Princ Stress (ksf) <73 3.43 1.90 1.991

Maximum Deviator Stress (ksf) P/A 2.45 4.24 3.45

Time Max Deviator Stress tf 120 120 120

Total Major Pnnc Stress (ksf) ay 5.88 6.14 5.44

Pore Pressure at Max

Deviator Stress (ksf) M 2.36 0.38 0 . 45

Eff Minor Princ Stress (ksf)

/ <73 —*

_

Eff Major Princ Stress (ksf)

r <71 _ —

from S-5

Project ASHLAND TANK FAILURE Boring No S-4,s-5 Sample No U”1

Type of Test: Saturated-Unconsolidated-fandrained Sample Type: Shelby

C= 1.9 ksf <*>= 0*0 0 c' = " ksf <t>

= . 0 GANNETT FLEMING GEO'

Page 1 of 2 TRIAXIAL COMPRES

TECHNICAL LABORA

JSION TEST REPORT

TORY
Pore

Pressure

(ksf)

Shear

Stress

<Ksf>

Normal Stress (ksf)

0 5 10 15 20

Axial Strain (%)

Remarks: Max. Deviator Stress

taken at 15% strain

Project ASHLAND TANK FAILURE Boring No S^4 , s-5 Sample No u-l

Type of Test: saturated-Unconsolidated-Undrained Sample Type: shelby

e = ksf <6 3 8 c' = ksf <*>' =

0 GANNETT FLEMING GEO'

°age 2 of 2 TRIAXIAL COMPRES

rECHNICAL LABORA

>SION TEST REPORT

TORY

m

Specimen No 4

Diameter (in) De 2.83

Height (in) H0 5.60

Water Content (%) w0 17.4

.2 Dry Density (pcf) Too 110.8

e

Void Ratio 9o 0.549 |

Saturation (%) So 87.2

Water Content (%) —

"35 <s Dry Density (pcf) Toe —

"©

CO

Void Ratio ®e —

Saturation (%) Se —

Back Pressure (ksf) 11.52

Total Minor Princ Stress (ksf) <Ti 4.87

Maximum Deviator Stress (ksf) P/A 5.05

Time Max Deviator Stress tf 120

Total Major Princ Stress (ksf) 9.92

Pore Pressure at Max

Deviator Stress (ksf) M 2,75

Eff Minor Princ Stress (ksf) <n' -

Eff Major Princ Stress (ksf)

t <71 •
Deviator

Stress

(ksf)

and

Pore

Pressure

(ksf)

Shear

Stress

(ksf)

Normal Stress (ksf)

8

7

6

5

4

3

2

1

0 5 10 15 20

Axial Strain (%)

Remarks:

Specimen No 1 2

j 3 Diameter (in) Do 2.83 2.85 2.83

Height (in) Ho 5.60 5.60 ! 5 .60

Initial

Water Content (%) Wo 21.1 19.9 20.6

Dry Density (pcf) Too L 04.7 105.

1

. 04.2

Void Ratio ®0 .640 . 633- .647

Saturation (%) So 90.7 86.4 87.6

Before

Test

Water Content (%) wc — __

Dry Density (pcf) Toe __ —

Void Ratio ®c --

Saturation (%) Sc — —

Back Pressure (ksf) MS 1.44 1.44 1.44

Total Minor Princ Stress (ksf) <T3 1.31 2 ’. 71 4.19

Maximum Deviator Stress (ksf) P/A 4.87 7.00 7.88

Time Max Deviator Stress 45 19 33

Total Major Princ Stress (ksf) <7 1 6.19 9.71 12 . 07!

Pore Pressure at Max

Deviator Stress (ksf) M — —

Eft Minor Princ Stress (ksf) oY __

Eft Major Princ Stress (ksf)

1 <r i ——

Project ASHLAND TANK FAILURE Boring No S-5a Sample No U-l

Type of Test: Unconsolidated-Undrained Sample Type: Shelby

e = 1.20 ksf — 21.0 9 c' = _ ksf <y = GANNETT FLEMING GEOTECHNICAL LABORA

TRIAXIAL COMPRESSION TEST REPORT

TORY
NBS»114A irev. 2 ««c) 1. PUBLICATION OR 2. Performing Organ. Report NoJ 3. Publication Date

REPORT NO. • NBSIR 88-3792 JUNE 1988

U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET (See instructions)

4. TITLE AND SUBTITLE

Investigation into the Ashland Oil Storage Tank Collapse on January 2, 1988

5. author(S) John L. Gross, Felix Y. Yokel, Richard N. Wright, A. Hunter Fanney,

John H. Smith, George E. Hicho, T. Robert Shives

6. PERFORMING ORGANIZATION (If joint or other than N BS, see instructions)

NATIONAL BUREAU OF STANDARDS

U,S. DEPARTMENT OF COMMERCE

GAITHERSBURG, MD 20899

7. Contract/Grant No.

f . Type of Report & Period Covered

9. SPONSORING ORGANIZATION NAME AND COMPLETE AODRESS (Street. City. State. ZIP)

10. SUPPLEMENTARY NOTES

L_"1 Document describes a computer program; SF-185, FIPS Software Summary, is attached.

oil storage tank at the Ashland Petroleum Company Floreffe Terminal near West Elizabeth;

Pennsylvania, collapsed as it was being filled to capacity for the first time since it

was reconstructed at the site. The tank had been dismantled in Cleveland, Ohio, after

more than 40 years of service and reconstructed at the Floreffe Site in 1986. The

National Bureau of Standards (NBS) conducted an independent investigation of the

physical causes of the Ashland Tank collapse. Data were obtained from NBS field observations, laboratory and analytical studies, from the investigation of the Pennsylvania

Tank Collapse Task Force appointed by the Governor of Pennsylvania, and from the

Battalia Columbus Division investigation sponsored by the Ashland Petroleum Company.

The cause of the failure was determined to be brittle fracture, initiating from a flaw

existing prior to the reconstruction of the tank. Complete rupture of the tank shell

occurred because the steel of the shell was of inadequate toughness at the operating

temperature to prevent brittle fracture propagation. The steel did not meet the

fracture toughness requirements of the American Petroleum Institute Standard 650, 1984,

"Welded Steel Tanks for Oil Storage," which was effective at the time of reconstruction

of the tank. The collapse shows the importance of using steel with sufficient fracture

toughness to prevent propagation of a brittle fracture in tanks whose sudden failure

would mean unacceptable human, environmental or economic losses.

12. KEY WORDS (Six to twelve entries ; alphabetical order; capitalize only proper names; and separate key words by semicolons) Brittle fracture; collapse; failure investigation; fracture analysis; steel; tanks;

welded steel tanks

13. AVAILABILITY

| •KfiUnlimited

n For Official Distribution. D© Not Release to NTIS

20^02

Fr°m Suoerintendem of Documents, U.S. Government Printing Office, Washington, D.C.

5E1 Order From National Technical Information Service (NTIS), Springfield, VA. 22161

14. NO. OF

PRINTED PAGES

204

15. Price

$24.95

USCOMM-OC 9043-P80