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NBS SPECIAL PUBLICATION 621

U.S. DEPARTMENT OF COMMERCE/National Bureau of Standards

Failure Prevention in

Ground Transportation

Systems

MFPG

31st Meeting

c. 2
NATIONAL BUREAU OF STANDARDS

The National Bureau of Standards' was established by an act ot Congress on March 3, 1901. The Bureau's overall goal is to strengthen and advance the Nation's science and technology

and facilitate their effective application for public benefit. To this end, the Bureau conducts

research and provides: (1) a basis for the Nation's physical measurement system, (2) scientific and technological services for industry and government, (3) a technical basis for equity in trade, and (4) technical services to promote public safety. The Bureau's technical work is per- formed by the National Measurement Laboratory, the National Engineering Laboratory, and

the Institute for Computer Sciences and Technology.

THE NATIONAL MEASUREMENT LABORATORY provides the national system of

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Absolute Physical Quantities^ — Radiation Research — Chemical Physics —

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Engineering — Building Technology — Fire Research — Chemical Engineering^

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research and provides scientific and technical services to aid Federal agencies in the selection, acquisition, application, and use of computer technology to improve effectiveness and

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J-ibrary, £-01 Admin. BIdg.

OCT 2 5 1982

Natfonal Bureau of Standards

Failure Prevention in

Ground Transportation Systems

7

Proceedings of the 31st Meeting of the

IVIechanical Failures Prevention Group,

held at the National Bureau of Standards,

Gaithersburg, MD, April 22-24, 1980

Edited by:

T. Robert Shives

William A. Willard

Center for Materials Science

National Measurement Laboratory

National Bureau of Standards

Washington, DC 20234

The proceedings were sponsored jointly by NBS and the:

Office of Naval Research

U.S. Department of the Navy

Arlington, VA 22217

National Aeronautics and Space Administration

Goddard Space Flight Center

Greenbelt, MD 20771

Naval Air Systems Command

U.S. Department of the Navy

Washington, DC 20360

U.S. DEPARTMENT OF COMMERCE, Malcolm Baldrige, Secretary

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

Issued October 1982
Library of Congress Catalog Card Number: 82-600597

National Bureau of Standards Special Publication 621

Nat. Bur. Stand. (U.S.), Spec. Publ. 621, 223 pages (Oct. 1982)

CODEN: XNBSAV

U.S. GOVERNMENT PRINTING OFFICE

WASHINGTON: 1982

For Sale by the Superintendent of Documents, U.S. Government Printing Office, Washington, D.C. 20402.

Price $7.50

(Add 25 percent for other than U.S. mailing.)
FOREWORD

The 31st meeting of the Mechanical Failures Prevention Group was held

April 22-24, 1980, at the National Bureau of Standards in Gaithersburg

Maryland. The program was organized by a special ad-hoc committee

to bring together experts concerned with failures in ground transportation systems. H. R. Hegner of ManTech International Corporation

served as Program Chairman. He was assisted by E. Passaglia, A. W.

Ruff, C. G. Interrante and J. H. Smith of the National Bureau of

Standards. Special appreciation is extended to K. L. Pierson, D.

Dancer and J. Krugler of the Department of Transportation for providing

speakers and contacts within the DOT.

Appreciation is extended to the following members of the NBS Fracture

and Deformation Division: T. Robert Shives and William A. Willard

for their editing, organization and preparation of the proceedings,

and Joel C. Sauter for photographic work. Most of the papers in the

proceedings are presented as submitted by the authors as camera ready

copy. Some minor editorial changes were required.

Gratitude is expressed to the following members of the NBS Metallurgy

Division: Marian L. Slusser for handling financial matters and

Kimberly A. Smith for typing.

Special thanks are extended to Jo Ann Lorden of the NBS Public

Information Division for the meeting, hotel and banquet arrangements.

HARRY C. BURNETT

Executive Secretary, MFPG

Center for Materials Science

National Bureau of Standards

III
TABLE OF CONTENTS

FOREWORD

SESSION I: RAIL VEHICLES AND STRUCTURES

A. Component Reliability of Railroad Freight Car Trucks.

Milton R. Johnson

B. Mechanical Behavior and Potential Failure Mechanisms

of Railroad Locomotive Pads (oral presentation only,

no manuscript submitted). R. Landel

C. Fracture of Steel Plate Materials under Abusive

Service Conditions in Railroad Tank Cars.

C. G. Interrante

D. Fracture Analysis of Cast Steel Components in Rail

Vehicles. Y. Sharir, D. H. Stone and W. S. Pellini

SESSION II: RAIL VEHICLES AND STRUCTURES

A. Requirements for On-Board Failure Detection Systems

for Rail-Vehicles. J. V. Mirabella

B. Detection of Rail Defects and Prevention of Rail

Fracture. Oscar Orringer and Harry L. Ceccon

C. Automated NDE for Detection of Braking Abnormalities

of Trains (abstract only). J. D. Ferguson

SESSION III: HIGHWAY AND ROAD BRIDGES

A. Failure Analysis of Highway Bridges. John W.

Fisher and Hans Hausammann

B. Failure Analysis of Dan Ryan Rapid Transit Structure.

John M. Hanson

C. Bridge Welding and Fracture Control. Carl E. Hartbower

IV
D. Theory and Design of Instrumentation for Bridge 143

Investigation. D. A. Bracher, D. A. Garrett

and C. 0. Heller

SESSION IV: PIPELINE TRANSPORTATION SYSTEMS

A. Ductile Fracture Analysis of Pipelines. F. Erdogan 153

B. Radiographic Variables and Weld Flaw Analysis. 165

Robert C. Placious

C. Development of Welding Consumables for Arctic 174

Pipelines (abstract only). M. Nakabayashi

D. Analysis of Pipeline Failure Data (oral presentation

only, no manuscript submitted). Paul Hinchliff

SESSION V: MOTOR CARRIERS

A. 1978 Roadside Vehicle Inspections. Kenneth L. Pierson 177

B. Regulation, Legislation and the Teamsters. 186

R. V. Durham

C. Stress Systems Related to Fracture of Ductile and 196

Brittle Materials. Donald J. Wulpi

D. Perspectives on Diagnostic Systems for the Trucking 201

Industry. V. H. Miles

E. Investigation of Defects in Motor Carrier Activities. 212

Robert F. Hellmuth

APPENDIX

List of Registrants for the 31st MFPG Meeting 217

MFPG Publications 221

V
ABSTRACT

These proceedinas consist of 18 submitted entries (15 papers and 2

abstracts) from the 31st meeting of the Mechanical Failures Prevention

Group which was held at the National Bureau of Standards, Gai thersburg

,

Maryland, April 22-24, 1980. The theme of the symposium was failure

prevention in ground transportation systems. Areas of interest included

rail vehicles and structures, highway and road bridges, pipeline transportation systems, and motor carriers.

Key words: bridges; diagnostic systems; failure; failure detection

systems; fracture; fracture control; ground transportation; motor

carriers; pipelines; rail structures; rail vehicles; reliability;

transportation systems.

UNITS AND SYMBOLS

Customary United States units and symbols appear in some of the

papers in these Proceedings. The participants in the 31st meeting of

the Mechanical Failures Prevention Group have used the established

units and symbols commonly employed in their professional fields.

However, as an aid to the reader in increasing familiarity with the

usage of the metric system of units (SI), the following references are

given:

NBS Special Publication, SP330, 1981 Edition, "The International

System of Units.

ISO International Standard 1000 (1973 Edition), "SI Units and Recommendations for Use of Their Multiples."

IEEE Standard Metric Practice (Institute of Electrical and Electronics

Engineers, Inc.), Standard 268-1979.

Disclaimer:

Certain trade names and company products are identified in order

to adequately specify the experimental procedure. In no case does such

identification imply recommendation or endorsement by the National

Bureau of Standards, nor does it imply that the products are necessarily

the best available for the purpose. Views expressed by the various

authors are their own and do not necessarily represent those of the

National Bureau of Standards.

VI
SESSION I

RAIL VEHICLES AND STRUCTURES

Chairman: D. Dancer, Federal

Administration

Railroad

COMPONENT RELIABILITY OF RAILROAD

FREIGHT CAR TRUCKS

Milton R. Johnson

IIT Research Institute

Chicago, Illinois 60616

Abstract: The modern freight car truck is a relatively simple structure which provides for weight transfer between the car and track,

maintains the proper positioning of the wheels with respect to the

rail, contains the suspension elements for isolating the effects of

wheel/rail forces and acts as a support for the brake. The number of

freight car truck components which fail and cause accidents is low with

respect to the number of cars in the railroad fleet. Continued efforts

for their improvement are warranted because of the potential seriousness of accidents. Most freight car truck components function as safelife rather than fail-safe structures requiring a conservative approach

to their design. The reliability problem is complicated by the practical difficulties of making adequate field inspections. Current efforts

to improve reliability include gaining a better understanding of complex wheel failure phenomena, improving design and performance specifications, developing new systems for wayside inspection and developing

new truck designs for the reduction of wheel /rail forces.

Key words: Railroad freight car; freight car truck; fatigue; derailments; railroad accidents; reliability; railroad testing.

Introduction: A number of important trends are evident in current

railroad operations. First, there is the use of larger freight cars. Most new cars are of nominal 100 ton capacity, which allows a 32,875 lb

wheel load. Larger wheel loads are not permitted in normal interchange

service. As wheel/rail loads become larger there is more opportunity

for developing damaging loads on both the track structure and vehicle.

The second important trend is the increased movement of bulk commodities

in unit trains. Unit train cars will all have similar dynamic behavior

so that motions and forces excited by vehicle/track interaction will

tend to be repetitive from one car to the next. These repeated loads

will accelerate the development of damage at particular track locations.

Third, there is the tendency for some railroads to allow track quality

to deteriorate. This increases the dynamic vehicl e/ track interaction

effects, hastens the onset of track deterioration, and increases the

risk of derailments

The increase in the derailment rate raises safety concerns. Derailments not only have the potential of causing serious casualties and

3
inconveniences to the public, but they have an adverse affect on railroad operations as well. Of special concern is the probability that

damage to cars containing hazardous materials might lead to a major

catastrophe. In recent years there have been a number of accidents

where widespread damage and numerous casualties were caused by the re- lease of hazardous materials from tank cars. Although the loss of life

has been relatively low from these events, one can conceive of accident

scenarios which could lead to major catastrophes.

Insuring the reliability of the freight car truck is one aspect of

safer rail transportation. This paper reviews the reliability characteristics of freight car trucks. Specifications used for their design

are also reviewed. Actions being considered to improve reliability are mentioned. Finally, some indications are given of the kind of research

and development which is likely to lead to further improvements in re1 iabil i ty

Aspects of Reliability: One must recognize several aspects of the term

"reliability". First, there is reliability with respect to accident

free operation. Safety is equivalent to reliability in this context.

Second, there is reliability with respect to the early removal of com- ponents from service because they are defective. A defective part may

be functional, but it exhibits characteristics which are indicative of

a premature failure. Removal of defective components is of economic concern because the desired life of the part is not obtained. Third,

the relationship between reliability and design must be recognized.

Reliability can be enhanced by design procedures which are based on an adequate definition of the environment in which the part functions as well as a complete understanding of the response of the component to

that environment. Finally, the relationship of reliability to inspection practices must be considered. Since most failures are the result

of defects which progress over a period of time, like a fatigue crack,

reliability can be enhanced by the increased frequency and thoroughness

of inspections. Adequate inspections are also required during the

manufacturing process so that the quality of components introduced into

service are maintained.

Description of Conventional Freight Car Truck: The principal features

of the conventional four-wheel freight car truck are illustrated in

Fig. 1. The weight of the car is concentrated by a structural element

called the body bolster. This mates with the truck bolster at the

center plate in a pivot connection which provides for the swivel ing of

the truck to accommodate yawing angular displacements between the truck

and car. The truck bolster, a steel casting, distributes the load from

the center plate to the two spring groups, one at each side of the

truck. These are supported by the side frames. Two side bearings, one

on either side of the center plate provide an alternate means of load

transfer between the car and the truck bolster. These are engaged by a

4
slight rolling motion of the car.

FIGURE 1 CONVENTIONAL FREIGHT CAR TRUCK

Conventional freight car truck designs utilize only one type of suspension element, namely, coil springs between the bolster and both the

side frames. The side frames, which are also steel castings, distribute the load from the base of the springs to the two axles of the

truck. The load is applied to the axles through bearings contained in

housings which are independent of the side frame, but effectively con- nected to it through the jaws or pedestals at the ends of the side

frame. The use of roller bearings has been required on all new cars built since August 1968. Over 60 percent of the freight car fleet is now equipped with roller bearings. The wheels are rigidly mounted to

the axles so that a wheel/axle set turns as a unit.

Shock and vibration isolation is accomplished by relative motion between the bolster and the side frames. The primary freedom for relative

movement is in the vertical direction; the total possible motion depends on the characteristics of the spring group. Relative motion is

also permitted in the lateral direction, although to a lesser extent.

The system must be damped to prevent excessive vertical motions, which

5
would result in bottoming out the springs with accompanying hard solid

blows to the bolster. Most trucks are damped by a shoe which provides

frictional forces between the bolster and the side frame columns. The

spring-loaded friction shoes also offer a resistance to skewing forces

which tend to distort the rectangular array of the truck side frames

and axles.

The freight car truck also supports the brake. The air cylinders which

initiate the braking forces may be attached to the truck components or the car, but in either case the linkage which transfers the piston rod

forces to the brake shoes is connected to the truck components. In

this way the truck structure transmits the braking forces from the

wheels to the car.

The overall reliability of the freight car truck is a function of the

reliability of the components which go to make up the system. It is

also influenced by the nature of the structural design philosophy,

whether it is of fail-safe or safe-life design. Fail-safe design relies on the utilization of redundant structural members. The probability of a catastrophic system failure is small because failure of the

system requires simultaneous defects in each of the redundant members.

There are only a few examples of fail-safe design principles in the

freight car truck. One is the truck-bolster/body-bolster connection

where the center plate rim and the center pin provide redundant structural elements in the connection between the two bolsters.

Safe-life design requires maintaining the integrity of each structural

element. Structural components must be designed for an extremely low

probability of failure during intervals between inspections. Conventional freight car truck construction is based primarily on safe-life

principles. Failure of one of the major structural elements, such as the side frames, bolster, wheels, or axles inevitably leads to derailment.

Changes in freight car truck design have taken place at a relatively

slow pace. This is due in part to the large number of cars that are in

the railroad fleet and the need for maximizing the useful life of this

equipment. Changes in design must be compatible with existing operations, and a changeover of the entire fleet must be planned to occur only over an extended period of time. The economic factors are important too. The nature of freight car operations, wherein the average

car moves only 60 miles per day, requires that it have a low capital

value and this severely limits the alternate approaches that can be

considered in freight car truck design.

Truck Component Loads: The vertical load is the principal load acting

on the truck and is due to the weight of the car and lading. It is

modified by transient factors, such as rail irregularities, suspension

6
system oscillations, wheel flat spots, etc. Most of the fluctuations

in the vertical load occur at frequencies below 10 Hz. Car rolling

motions result in an alternating component of load on each side of the

truck at a characteristic frequency of approximately 1 Hz for a loaded

car. Car bouncing and pitching motions cause load fluctuations in the

3 to 5 Hz range (see Ref. 1). The 1 i ne-of-action of the vertical load

is normally at the center of the bolster, but this can be shifted be- cause of load transfer to one of the side bearings. Load transfer to

the side bearings is most severe on curves or during car "rock and roll"

motions. Car rock and roll motions are enhanced by a high center of

gravity, a truck center distance approaching rail length, and operation

at critical speed on jointed rail .

Lateral truck loads are the result of load transfer from the car body,

such as would be caused by the angularity of the draft force on the car,

the superelevation of a curved track rail not commensurate with speed,

guiding forces for the negotiation of curves, or flange contacts during

the hunting motion of the wheel-axle set. Lateral loads are also associated with the internal truck forces which tend to maintain the truck

in a rectangular configuration while traversing curved track. Lateral

creep and slip forces are also built up at the wheel /rail contact

points. These forces are normally directed toward the flange. Occasionally large transient forces can be directed in the opposite direction due to guardrail contacts. Severe lateral loads also can be

developed at the wheel/rail interface under the self-excited vibrational

condition commonly referred to as truck hunting. This condition is

usually associated with a light weight car traveling at high speeds.

Horizontal twist loads between the side frame and bolster result from

the tendency of the side frame to rotate into an out of square position

with respect to the bolster during the traversal of curved track or due

to truck swivel ing motions. Longitudinal loads result from braking

forces and the inertia! forces accompanying acceleration of the truck.

The most severe forces occur under car impact conditions when an un- loaded car is coupled at high speeds into a standing string of cars. The longitudinal load is applied to the bolster at the center plate rim

and is reacted at the side frame columns. The load transfer through

the center plate rim has been shown to be one of the most severe loading conditions for the truck bolster (Ref. 2).

Wheels are subjected to thermal loading when tread brake shoes are applied. The deposition of energy in the rim leads to temperature

gradients within the wheel. This causes an expansion and twisting of

the rim relative to the plate. There are two distinct types of severe

braking service which lead to significant thermal loads. The first is

drag braking which is associated with the descent of a long grade.

This might involve braking for over an hour at moderate rates of energy

deposition in the tread of the wheel (e.g., 20 bhp) . The relatively

7
long duration of the brake application allows the heat to penetrate

through the rim down into the plate (see Ref. 3). The second type of

service is emergency braking. The relatively short time of brake

application means that most of the heat is retained near the surface of

the tread resulting in sharp thermal gradients within the wheel.

Failure Modes: Catastrophic failures of truck castings (side frames

and bolsters) are usually due to the development of fatigue cracks. Be- cause of the complexity of truck casting designs and the inherent characteristics of foundry practice, it is difficult to eliminate all con- ditions which may lead to the development of fatigue cracks. Fatigue

cracks may originate at shrinkage cavities within the castings, the

edges of holes, cracking strips, sharp surface corners, or geometrical

discontinuities at core joints.

The development of a fatigue crack on a side frame is a relatively rare occurrence. The most common locations are through the tension member

(see Fig. 2) or through the column. The tension member crack is the

FIGURE 2 SIDE FRAME NOMENCLATURE

most critical because the propagation of the crack through the member

can lead to the collapse of the side frame. On relatively rare occasions fatigue cracks develop in the compression member near the point

where it joins the tension member above the pedestal opening.

Bolsters are also subject to the development of fatigue cracks. The

most serious location of the initiation of a fatigue crack is on the

bottom surface because these cracks eventually propagate up the side

and lead to its collapse. These cracks often originate at one of the

bottom lightner holes and propagate up to one of the side lightner

holes (see Fig. 3)

8
FRICTION SHOE POCKET CENTER PLATE RIM

BRAKE ROD OPEN INGS

turn of the spring

seat" region

lightner" hole.<^,

side and bottom

FIGURE 3 TRUCK BOLSTER NOMENCLATURE

Fatigue cracks also develop at the center of the bolster below the

center plate and propagate up through the sides and inside ribs. Fatigue cracks can also develop on the bottom surface in the turn-of-thespring-seat area, but this is a relatively rare occurrence.

The Association of American Railroads (AAR) Wheel and Axle Manual (Ref.

4) lists 22 types of wheel defects which require removal of wheels from

service including excessively worn tread or flange, evidence of overheating, and the presence of cracks. From the standpoint of ensuring

the structural adequacy of the wheel the most serious of these defects

are thermal (radial rim) cracks, plate cracks and shattered rims be- cause they are the type of defects which are most likely to lead to the

catastrophic failure of the wheel.

Cracks which originate on the periphery of the wheel and propagate inward toward the hub are commonly referred to as thermal cracks. Once

these cracks propagate through the rim they will often turn and propagate as a circumferential plate crack. Sometimes the crack will then

reenter and propagate through the rim in a radial direction causing the

separation of a large segment of the wheel. A thermal crack might be

initiated and propagate over an extended period of time or the crack

may develop rapidly causing a sudden failure of the wheel.

The thermal crack development is associated with modifications in the

residual stress field within the wheel which is caused by severe tread

braking. Under these conditions the periphery of the wheel is heated,

but it is restrained from expanding by the colder plate and hub of the

wheel. This causes the development of circumferential compressive

stresses in the rim and radial tensile stresses in the plate. If the

9
thermal gradient is large enough, plastic deformation will take place

first in the highly stressed plate fillet regions and, with increasing

severity of the thermal load, in the rim adjacent to the surface of the

tread. If the thermal load causes inelastic deformation in the rim

this region will develop residual tensile stress when the wheel cools.

The existence of tensile stress in the rim will promote the initiation

and growth of thermal cracks.

Cracks that initiate in the plate of the wheel and propagate circumferential ly around the plate are commonly referred to as plate cracks.

They normally develop and propagate as fatigue cracks from a point of

initiation in the outside plate hub fillet. Plate crack development is

believed to be caused by the effects of both thermal and mechanical

loads. The highest stresses in the regions of crack initiation result

from the thermal effect, but there are a relatively small number of

cycles of high stresses over the life of the wheel. Mechanical wheel

loads produce an alternating stress pattern within the critical fillet

regions of the wheel once per wheel revolution, but the stress levels are of relatively low magnitude. The relative importance of these

stress cycles has not yet been established.

The development of cracks in the rim slightly below the surface of the

tread, and their propagation in a generally circumferential direction

so that portions of the tread are broken off, is commonly referred to as spalling or shattered rim failure. It is caused by the fluctuations

in the stresses surrounding the wheel/rail contact point as the load is

repeated once per wheel revolution.

The most common cause of complete journal bearing failure is overheating. As a result of the loss of lubrication or some other defect

in the bearing, the frictional forces increase causing the bearing to run hot (Ref. 5). This condition, if not detected, progressively be- comes worse. Temperature in excess of IBOO^F can be reached and the

ability of the end of the axle to carry the bearing load is lost. As

a result the end of the axle breaks off which often leads to a derailment.

Other modes of truck component failure include side bearings becoming

loose and fal 1 i ng off the bolster, which leads to excessive sway of the

car, and excessive wear in the components of the friction snubber

system, which leads to uncontrolled oscillations of the car.

Accidents Caused by Failure of Freight Car Truck Components: Accidents

caused by failures of freight car truck components are relatively rare

events when one considers the large number of truck components in ser- vice. Table 1 presents a 5 year average of the number of derailments

included in the Federal Railroad Administration (FRA) accident statistics which were caused by the failure of one of the principal freight

10
car truck components. During this period there was an average of 29.4

billion freight car miles per year.

TABLE 1 DERAILMENTS CAUSED BY TRUCK

COMPONENT FAILURES, 5 YEAR AVERAGE (1973-77)

Cause Average Derailments Per Year

Broken or Loose Wheels 173

Worn Wheel

175

Broken Axles 15

Defective Journals 290

Defective Side Bearings 190

Broken Side Frames or Bolsters 63

The information in the FRA Railroad Accident/Incident Reporting System

(RAIRS) data base can be processed to examine certain characteristics

of derailments caused by the failure of freight car truck components.

Of particular interest is the conditions under which a typical accident

takes place and the resulting damage.

The results of such an analysis for 1977 data are summarized in Table 2.

The first column shows the groupings of cause codes used to summarize

categories of failure. The next column shows the total number of derailments that were analyzed for the year. The totals differ slightly

from FRA accident summary tables because derailment cases were eliminated where certain items were omitted in the report.

The third column shows the median speed of derailment. The median

value is used rather than the mean in order to get an indication of a typical derailment. Note that there is a large difference between

certain derailment categories. For example, derailments caused by

excessive wheel wear, improper truck swivel ing, and defective side

bearings occur most often at relatively low speeds. On the other hand

derailments which are due to fractured components such as bolsters,

side frames, wheels or axles tend to occur at moderately high speeds.

This may be due to the fact that the final separation of the component

is caused by a relatively high impact loading which is more apt to occur at higher speeds.

11
TABLE 2

ANALYSIS OF DERAILMENTS CAUSED BY TRUCK COMPONENT FAILURES

Defect Category

(Cause Codes)

Number

Of

Cases

Median

Derailment

Speed

(mph)

Median

Number

of Cars

Derail ed

Derailment

Damage (dollars)

90

Median Percentile

Defective Side

Bearings (440-2)

234 15 4 $ 7,900 $ 46,000

Broken Truck

Bolster (443)

25 35 3 5,200 500,000

Broken Truck

Side Frames (444)

9 30 1 10,000 270,000

Improper Truck

Swivel ing (445)

oo 6 7,400 49,000

Defective Truck

Springs or Snubbers (446-7)

11 17 3 12,500 80,000

Broken Wheels

(460-3)

145 38 2 13,000 200,000

Excessive Wheel

Wear (464-6)

171 5 4 5,000 20,000

Loose Wheels

(467)

41 23 4 14,700 220,000

Broken Axles

(450)

13 40 4 19,200 120,000

Defective

Journal

s

(451_454)

269 30 2 14,000 140,000

The fourth column in the table shows the number of cars derailed.

Again the median value is used. Note that there are no major differences. Two, three, or four car derailments are typical. This is

somewhat surprising in view of the differences indicated for the median

speeds of derailment. It would be expected that higher speed derailments would result in more derailed cars, but a further review of the

-12
data indicates no significant correlation between speed and the number

of derail ed cars

.

The fifth and sixth columns show the value of damage caused by the

derailment. This includes both equipment and track damage. The fifth

column shows the median value. There is no significant trend except a slight correlation with respect to speed. Typical derailments produce

$5,000 to $15,000 damage. The sixth column gives an indication of the

potential severity of a derailment by presenting the 90 percentile

damage value. The meaning of this value is that 90 percent of the

derailments would have damages below the level indicated and 10 percent

would have damages above this value. There are large differences between derailment causes. Derailments resulting from excessive wheel wear show a 90 percentile value of $19,000 where $500,000 is shown for

broken truck bolsters. Note that broken components like wheels and

truck castings have the largest values and thus offer the potential for severe derailments.

Removal of Defective Components : The study of accident statistics alone

does not present a complete picture of the number of components which

fail in service. Many components are found to be defective during

routine inspections and are removed before they fail . Thus the number

of components which become defective under service conditions is substantially larger than the number of component failures which lead to

accidents

.

The AAR Truck Casting Removel Reports, for example, indicate that about

1000 truck bolsters and approximately one- fourth this many side frames are removed from service yearly because they are defective. Excessive

wear is one of the major reasons for their removal . When severe wear patterns develop they occur at the interconnections between components

(e.g., the bolster friction shoe pockets, side frame columns, the bolster center plate area).

The most common location for the development of severe wear on a side

frame is the column. This type of wear results from interactions with

the truck bolster caused by persistent truck swivel ing motions or by

excessive vertical oscillations resulting from a poorly damped suspension system.

Many bolsters are removed from service because of cracked or broken

center plate rims. This is generally due to high speed car coupling

impacts of empty cars. Another reason for bolster removal is the

development of badly worn center plate rims. The major cause of this

condition is uncontrolled truck swivel ing motions which occur during

high speed operations with light weight cars.

Several hundred thousand wheels are removed from service each year be- cause of defects. An analysis of the reasons for their removal shows

13
that only a small number (on the order of 3 percent) are removed be- cause of evidence of cracking or overheating, which would make the

wheel more susceptible to fracture (Ref. 6).

Bearings are usually removed from service for seal defects, loose end

caps, and evidence of overheati ng( Ref . 5). It has been established that

in many cases overheating results from overl ubrication of the bearing.

When bearings are removed for inspection they are examined for defects

such as oversized ring or cone, fatigue pits or spalls on a cup race,

brinelling, and corrosion on roller or raceway.

Design Specifications: Side frames and bolsters are cast with low to

moderate carbon steels meeting the requirements of AAR Specification

M201 (Ref. 7). Most castings are made with either Grades B or C steel

having minimum yield strengths of 38 and 60 ksi respectively.

Bolster and side frame castings must be designed in accordance with AAR

specification M202 and M203 (Ref. 7). These specifications require

that nominal design stresses must be kept below 16,000 psi in Grade B castings for prescribed combinations of vertical and lateral loads.

The corresponding maximum design stress for Grade C castings is 25,000

psi. The specified vertical load used in side frame design calculations is 1.5 times the nominal maximum static load on the side frame.

The specified vertical design load used in the bolster calculation is

the nominal static load.

Side frame and bolster designs must be subjected to static load tests

in order to demonstrate certain load/deflection criteria and maximum

load carrying capacity.

Side frames are also subject to fatigue test requirements under M203.

This specification calls for new side frame designs to be tested in

machines which apply a complex pattern of vertical, lateral, twist, and

impact loads on side frame castings. The requirement calls for testing

a group of four sample castings for an average of 100,000 cycles of

load without the development and extension of a transverse crack to 1/2

i nc h

Railroad wheels are manufactured in accordance with AAR Specification

M107 for wrought steel wheels or M208 for cast steel wheels (Ref. 7).

Five classes of moderate to high carbon wheel steels are defined on the

basis of carbon content and hardness.

There are no design stresses specified for wheels. Instead various

geometric configurations are specified. Minimum dimensions are specified at critical locations (e.g., plate thicknesses at hub and rim

fillets)

Roller bearings must possess a minimum life expectancy of 500,000 miles

14
with the full rail load acting radially for one-half the mileage. Life

expectancy is defined as the life at which no more than 10 percent of

the bearings require replacement due to fatigue initiated spelling or flaking on the load carrying surfaces (Ref. 8).

Axle design is specified in terms of dimensional requirements which are

based on limitations in stress values (Ref. 4). The allowable stress

levels reflect extensive fatigue tests.

Inspections: Inspection requirements for freight car truck components

are based on the AAR Interchange Rules and the current FRA Railroad

Freight Car Safety Standards (49CFR Part 215). The FRA standards re- quire a predeparture inspection to insure that cars are not dispatched

with critical defects which have the potential to cause derailments.

Wheels must be checked for excessive wear, cracks, and signs of overheating. Any axle with a crack or with a gouge more than 1/8 inch

deep must be withdrawn from service. Roller bearings must be checked

for signs of overheating and damaged seals. Roller bearings must also

be checked for defects when disassembled to permit wheel replacement.

Side frames and bolsters must not be permitted to operate with cracks

1/4 inch or more in the transverse direction on a tension member.

Actions to Improve Reliability: A number of activities are underway

which are expected to improve the reliability of freight car trucks.

Wheels are expected to become more reliable by the development of an improved design specification. The AAR has a program to determine the

feasibility of specifying maximum design stresses under various load

conditions as a means of improving the integrity of wheel designs. The

feasibility of using a design stress condition has been made possible

by the development of modern finite element analysis techniques which

allow for the calculation of stresses in wheels under complex loading

conditions including mechanical loads at the wheel/rail interface and

thermal loads introduced by tread braking.

The reliability of truck bolsters is expected to be improved by the

development of a fatigue performance specification. The RPI/AAR Truck

Research Safety and Test Project has been working on this specification

for several years. The work began by making measurements of the loads

acting on freight car trucks over many thousands of miles of railroad

service in order to adequately define the environment in which freight

car truck components operate (Ref 9). Fatigue tests were also conducted on a number of bolsters to determine their fatigue strength. Based

on these results it has been possible to develop recommendations for a fatigue test specification which will insure that a given bolster design has an adequate margin against failure during its anticipated life.

A trend in the construction of conventional freight car trucks is utilization of longer spring travel. Most new cars now are equipped with

15
3-11/15 inch travel springs. Within the last 2 years, the use of the

D-7 spring, which gives 4-1/4 inch travel, was introduced. Wheel/rail

interaction forces are significantly reduced by the use of longer travel suspension springs.

The concern over truck hunting phenomena has led to the development of

devices which prevent this unstable type of motion. One way of reducing

the hunting tendency is to provide some damping restraint to truck

swivel motions. The use of the center plate extension pad or constantcontact side bearings are ways of providing this restraint.

Over the last several years, much activity has been directed toward the

development of new types of freight car trucks. The goal has been the

development of trucks with features which will reduce wheel wear and

wheel /rail interaction as a means of increasing safety and reliability.

These trucks would be more expensive than conventional trucks, but the

reduction in wheel wear and other beneficial effects would be expected

to more than compensate for the added cost.

Different design philosophies are evident in these advanced truck designs. Truck hunting phenomena, for example, are minimized if the

truck is rigidized to reduce its tendency for out-of-square deformations, and this feature is evident in most advanced truck designs. The

principal design objective of other trucks is the reduction of wheel wear and wheel /rail forces on curves through the use of self-steering

mechanisms (Ref. 10).

Long Range Activities: Work of a long range nature which is now in

progress can be expected to lead to more reliable freight car truck

operation. In recent years instrumentation systems have been developed

for accurately measuring the wheel/rail interaction forces. Measurement programs have been conducted, and more are planned, to quantify

this aspect of the service load environment. The lack of such data in

the past has hampered efforts to get more reliable designs.

Other experimental work is being conducted at the Facility for Accelerated Service Testing. This facility is at the Transportation Test

Center (TTC) and is used to investigate the long-term effects of railroad operations on track and equipment. This work is being conducted

on a special track which is made up of a number of different types of

track construction. A mixed train of 100-ton capacity cars operates

over the track to generate a high rate of usage.

Another facility, the Stability Assessment Facility for Equipment, is

being developed at TTC to evaluate the dynamic response characteristics

of rail vehicles. The track is instrumented to measure vertical,

lateral, and longitudinal forces, and displacements. The track will

contain known geometric variations to induce vehicle dynamic modes.

16
The utilization of advanced wayside inspection equipment offers the

possibility of identifying and removing defective cars before they lead

to accidents. Some wayside inspection systems, like hot box detectors,

have been utilized for many years. Other equipment, such as a system

to detect improper dynamic operation of the suspension system, is in

the early stages of development. A joint government/industry group has

been organized to oversee the installation and operation of a wide

variety of wayside inspection systems at TTC. These devices are located on a special track section where cars containing known defects

can be used to examine the responsiveness of the detection equipment.

References

1. Johnson, M.R., "Summarization and Comparison of Freight Car Truck

Load Data," Transactions of the ASME, Journal of Engineering for

Industry, Vol. 100, No. 1, February 1978

2. Johnson, M.R., "Effects of Longitudinal Impact Forces on Freight

Car Truck Bolsters," Federal Railroad Administration, Report FRA0R&D-75-10, September 1974

3. Johnson, M.R., "Measurement of Railroad Car Wheel Strains Under

Operational Conditions Including Prolonged Drag Braking," Final

Report, Volume I, Contract DOT-TSC-855, December 1976

4. "Wheel and Axle Manual," Association of American Railroads,

Mechanical Division

5. McGrew, J.M., Krauter, A.I. and Moyar, G.J., "Reliability of

Railroad Roller Bearings," Transactions of the ASME, Journal of

Engineering for Industry, Vol. 99, No. 1, January 1977

6. Hengel , M.F., "Freight Car Wheel Performance," Sixth International

Wheel set Congress, Colorado Springs, Colorado, October 1978

7. "Manual of Standards and Recommended Practices," Section A,

Association of American Railroads, Mechanical Division

8. "Manual of Standards and Recommended Practices," Section H,

Association of American Railroads, Mechanical Division

9. Evans, R.A. and Johnson, M.R., "Analysis of Environmental Truck

Component Load and Bolster Fatigue Test Data," AAR Research and

Test Report R-246, September 1976

10. Tsai, N.T., Rama Chandran, P.V. and Cappel , K.L., "Truck Design

Optimization Project-Assessment of Truck Design Characteristics,"

ASME Technical Paper 79-WA/RT-15, December 1979

17
FRACTURE OF STEEL PLATE MATERIALS UNDER ABUSIVE

SERVICE CONDITIONS IN RAILROAD TANK CARS

C. G. Interrante

National Bureau of Standards

Washington, DC 20234

Abstract: In recognition of the potential for catastrophic accidents

involving railroad tank-car operations, the Federal Railroad Administration has actively participated in a continuing industry/government

research program to enhance the safety of these operations. Selected

parts of this program will be discussed with emphasis on the challenges

presented by abusive service conditions that are associated with accidents, the ameliorative measures that have been taken to date, and the

direction of the present work aimed at furnishing improved guidelines

for fracture-safe design.

Key words: Fracture control; hazardous materials; impact transition;

pressurized tank car; stress-rupture.

Background- -Studies related to the behavior of railroad tank cars in

abusive service are conducted in recognition of the potential for hazardous derailment and costly other consequences of failures of railroad

components. Facts about railroad service in the United States indicate

that freight service is big business [1]. About 500 billion net ton

miles of cargo are transported annually in about two million freight

cars. Out of this total railroad freight, about 40 billion net ton

miles of cargo are transported in almost two hundred thousand tank cars. As tank cars are commonly used to transport hazardous materials, their

safe performance under normal operating conditions as well as in accidents is of great importance to the railroad industry and to the public.

Recent data [2] indicate that the accident rate for all hazardous materials shipped as railroad freight (primarily in tank cars) is about

twenty accidents per billion ton miles. Thus, hazardous materials, are involved in about 800 railroad accidents annually. In these accidents,

the number of tank cars carrying hazardous materials is of the order of

4000, and about 200 of these cars release hazardous materials as a result of these accidents.

Accidents involving tank cars include both self-pressurized and unpressurized cars. Unpressurized cars are used to transport commodities that

have very low vapor pressures, such as flammable liquids (like naptha)

as well as various dry chemicals (like sulfur). Self-pressurized tank

cars are used to transport liquified gases that have significant vapor

pressures under normal operating conditions at atmospheric temperatures.

These pressurized liquids include chlorine, anhydrous ammonia, and propane.

18 I
While accidents involving the unpressurized cars can sometimes be serious

and costly, these cars can often be abused badly and even ruptured

without consequences of significance. Pressurized tank cars have a much

greater potential for catastrophe in accidents. The lading is frequently

a liquid, which may be flammable or toxic, with a significant vapor

pressure that increases with temperature and gives rise to increases in

the applied tensile stresses in the shell- and head-plate materials of

the tank car. The increased applied stresses can lead to a decrease in

the critical flaw size required for unstable fracture, as the critical

flaw size is related to the inverse square of the applied stress. For

these and other reasons, fracture-safe design is a continuing concern

for the pressurized tank cars. Pressurized steel tank cars are produced to the Department of Transportation (DOT) specifications for four basic classes: 105A, 109A, 112A,

and 114A, with Classes 114 and 112 having special requirements [3].

Although pressurized cars are sometimes fabricated from aluminum plate

materials (some cars of DOT Classes 105 and 109), the studies at the

National Bureau of Standards (NBS) have been concerned only with steel

plate materials. Most commonly, these steels are produced to specification AAR M-128 (Grades A and B). Other grades of steel permitted [3]

for use in this application are produced to various ASTM specifications

and they include ASTM A285-70a (Grades A, B, and C), A515-70, and

A516-70a (Grades 55, 60, 65, and 70), A537 (Grade A), and A302-70a

(Grade B).

Pressurized tank cars are fabricated by welding and they consist of a central section of a number of shell courses. This central section is

closed at each end by a 2:1 semi-el iptical head plate. The shell plates

are cold formed in the as-rolled condition. The head plates are hot

formed and the temperatures of forming should be controlled in a manner

that produces properties equivalent to those of normalized plates.

After forming, the plates are welded and the welded vessel, with its

integral attachments, is stress relieved. While some anomalies in the

welding practices employed on particular vessels have been observed [4],

results of impact tests of weldments have been favorable [5], and in

general weldments have experienced a rather favorable service record;

welding practices will not be discussed in this report.

A principal focus of the NBS studies of the plate materials used in the

fabrication of pressurized cars has been on a steel that is most commonly

used for these cars, AAR TC-128 steel (Grades A and B). This report

will deal with pressurized tank cars in general and with the relationship

between the properties of TC-128 steel and the requirements of the

abusive service that pressurized cars are apt to experience in service.

This steel is produced to the AAR M-128 specification [6], which is

revised periodically. This steel is commonly used as the plate material

for tank cars that are used in the transportation of hazardous materials

19
such as cars of the DOT Classes 112 and 114, which are used in transportation of propane and anhydrous ammonia. In this report, some aspects

of the M-128 specification will be discussed in relation to the requirements for service. In addition, general observations on the microstructure, the mechanical properties (especially impact properties), and

the quality of the plate material will be summarized. This is followed

by a brief discussion of the FRA/AAR cooperative programs that led to a rulemaking known as HM144 for hazardous materials, which involves significant changes in the regulations for transportation of hazardous materials.

Finally, work in progress on the development of analytical models for

computation of critical flaw sizes for unstable fracture in the pressure

vessels used in this application will be mentioned. In keeping with the

theme of this symposium, significant problems that have been solved

recently and challenging problems for the coming decade will be highlighted.

Comments on Specification AAR M-128--The specification M-128. 00 for high

strength carbon-manganese steel plates for tank cars requires a tensile

strength in the range of from 81,000 to 101,000 psi and a minimum yield

point of 50,000 psi, with elongation in two inches to exceed 19.0 percent [6]. Under this specification, AAR TC-128 Grades A and B steel

plate materials are produced to fine-grain practice and the plates are required to be normalized only when used for low temperature service.

The requirements for chemical composition, given as percent by weight,

include maximum levels of carbon and manganese of 0.25 and 1.35, respectively. Thus, the ratio of manganese to carbon (Mn/C) for steels produced

to this specification would be about five. This is favorable when

compared to that of older tank-car steels that are no longer permitted

for this application, ASTM A212-A and -B steels. These older steels had

still higher maximum levels of carbon (0.28 and 0.31 for Grades A and B,

respectively) and a lower maximum manganese level of 0.90 percent. The more favorable Mn/C ratio of the TC-128 steels would tend to promote

improved impact transition temperatures.

For cars built before 1978, the maximum level of phosphorus was 0.04 percent

and that of sulfur was 0.05 percent. These levels can be achieved

without special steelmaking practices and they might be considered to be

generously high. In July of 1978, these maximum levels were adjusted

downward to the current maximum phosphorous and sulfur levels of 0.035

and 0.040, respectively. Steel that had been aluminum killed and produced

to meet the requirements of Grade B of the M-128 specification (pre-1978)

has been shown [7] to fail the bend test. The inclusion content of this

particular steel is high in comparison with other plates tested (4, 5,

8), all of which passed the bend requirement of the M-128 specification.

The bend test requires that the specimens shall stand being bent cold

through 180° without cracking on the outside of the bend portion, through

an inside diameter equal to twice the thickness of the plate.

20
While the bend test can be used as a qualitative test of ductility and

steel quality, it is not a very definitive test related to the requirements for steel plates in abusive service. In this type of service,

temperatures may be significantly below room temperature, at which the

bend test is conducted, and loading rates may be substantially greater

than those used to conduct the bend test. The bend specimen is neither

notched nor pre-cracked. However, the fracture behavior of steel plates

of the types in question are well known to be sensitive to strain rate,

temperature, and the presence of notches and cracks. For these and

other reasons, the bend test is not considered a definitive test of

resistance to fracture. A definitive test must be one that is based on some measure of fracture toughness.

Fracture Toughness --For abusive service, it becomes apparent that a balance of properties is required. The requirements for strength and

ductility are based on traditional design requirements. The results of

tensile tests provide adequate measures of the yield strength, the

tensile strength, and the ductility. For abusive service, other properties may become important. It will be shown later that the short-term

creep (or stress-rupture) properties can be important when the plate

materials are exposed to a fire. Further, the fracture toughness must

be considered. This material property governs the size of a flaw that

will propagate unstably in a plate that has been ruptured or partially

ruptured in an accident.

Fracture toughness is generically defined as the resistance to propagation of a crack. Various measures of fracture toughness are available. For purposes of this report, these measures are categorized into

two classes, direct measures and indirect (or alternative) measures. The direct measures are those that have a highly analytical basis in

fracture mechanics. The results of these measures are normally expressed

in terms of the stress-intensity factor (K), or values of J derived

using the J-integral concept. The results of these tests are used

directly in fracture control. Normally, these tests are expensive when

compared with the alternative tests.

Results of alternative tests are often useful for fracture control,

based upon either service experience or empirical correlations with

fracture mechanics tests. While there are about seven different shapes

of specimens used as alternative tests for steels of the types in question here, only one is used widely. It is the single-edge notched

specimen, tested in three-point bending, such as the Charpy V-notch

(CVN) specimen and the dynamic-tear (DT) specimen; another common test

is the nil-ductility transition (NDT) test, which uses a face-notched

crack-starter weldment in a specimen tested in three-point bending.

These types of tests are used to characterize the fracture toughness of

tank-car steels because steels of the types used in this application

usually exhibit a strong transition from ductile to brittle fracture and

21
under impact loading this transition can occur at temperatures in the

service range of temperatures. Further, these tests are generally

inexpensive and effective in characterization of the transition.

With the above mentioned balance of properties in mind, the fracture

resistance of TC-128 and other carbon-manganese steels used in tank cars

is now further described. The microstructure is a mixture of ferrite

and pearl ite that is characteristic of carbon-manganese steels. The

fracture resistance of these steels is highly dependent on orientation

of the test specimen [5,9]. This is illustrated in Figure 1, which

gives the CVN energy absorption data for six orientations of specimens.

The specimens were tested over a range of temperatures selected to

establish the transition from ductile to brittle behavior. Specimen

orientations are designated by a two-letter code, with the first letter

representing the normal to the crack plane and the second letter representing the expected direction of crack propagation. The letters L, T,

and S are used to represent the longitudinal, transverse, and shorttransverse directions of the plate.

The effects of orientation of specimen and notch were found to be similar

for each of two steels tested, a head plate and a shell plate [9]. In

this work, it was shown that the orientations of the specimen and the

notch: (1) greatly affect upper-shelf values of lateral expansion

(L.E.) and of energy absorption (E.A.); (2) affects the transition

temperatures at which a given level of L.E. or E.A. is obtained; and

(3) has little effect on the lower-shelf values of L.E. and E.A. or on the temperature limits of the transition zone or on the 50 percent

shear-fracture-appearance val ues.

As an example of the above specimen orientation code, in TL specimens,

the normal to the crack plane is the transverse direction. In TL specimens, the direction of crack propagation is the longitudinal (principal

rolling) direction. Specimens of this orientation are commonly called

transverse specimens. This is the orientation of lowest resistance to

extension of a through-thickness flaw, and it is commonly called for in

specifications.

Resistance to extension of flaws present on the plate surfaces, represented by the TS and LS orientations, is greater than that for throughthickness flaws. Internal flaws, those contained within the surfaces of

the plate, are represented by orientations SL and ST. Tests of the SL

and ST orientations reflect the resistance to extension of these internal

flaws under stress that is normal to the rolled surfaces of the plate,

such as the stress that results from the welding of attachments to the

surfaces of the plate materials.

The results shown in Figure 1 indicate that resistance to extension of

surface flaws (LS and TS) is greater than that for the corresponding

through-thickness flaws (LT and TL) and resistance to extension of

22
internal flaws (SL and ST) is least. This anisotropy of impact properties typifies that of many steels commonly used for pressure-vessel

applications, and it demonstrates that resistance to propagation is

highly dependent on orientation. This anisotropy, and the fracture

resistance in general, are related to the quantity, types, and shapes of

inclusions in steels. Therefore, mill practices have been developed to

minimize these tendencies. For example, transverse properties in plate

materials are enhanced by (1) more extensive cross-rolling of the plates

and (2) inclusion shape control by chemical means. Fracture toughness

is generally improved by decreasing the amount of sulfur, as this

decreases the number and size of manganese-sul f ide inclusions.

Thus, transition temperatures are affected by the orientation of the

test specimen. This is one of many factors. Others include specimen

geometry, loading rate, and the presence of a pre-crack.

The dynamic-tear (DT) specimen is much larger than the CVN specimen. It

is 5/8-inch thick by 7 1/8-inches long by 1.6 inches wide. The CVN

specimen is 10 mm (0.394 inches) square by 55 mm (2.165 inches) long.

The thickness of plate materials used for tank cars is often 5/8- to

11/16-inches thick. As a result, the DT specimen can be prepared as a full-plate thickness specimen, but the CVN specimen represents only part

of the plate thickness. Therefore, variations in properties throughout

the thickness, which sometimes are observed particularly in head plates

[8,9], could be misrepresented in tests of Charpy specimens. Further,

it is noted that, due principally to its greater thickness, the DT

specimen yields higher transition temperatures when compared with those

of the CVN specimen. However, when the CVN specimen is tested in the

pre-cracked condition, the transition temperatures are often similar to

those of the DT test specimens. Thus, in using these alternative tests

for estimation of fracture toughness of pressure-vessel steels, it is

important to recognize that the CVN specimen is not a full plate thickness specimen. Further, if the CVN specimen is not pre-cracked, it is

likely to give transition temperatures that are significantly lower than

those obtained with either a pre-cracked CVN specimen or with the DT

specimen.

The DT specimen is prepared with a sharp pressed notch. During a DT

test this sharp notch facilitates initiation of a crack. Then the crack

propagates to completely fracture the specimen; the bulk of the energy

absorbed in this test represents the energy required to propagate the

crack. With its 1.6 inch width, the full crack extension is more than

three times that obtained with a CVN specimen. This permits the

DT specimen to build up large shear lips during propagation of the crack

along the full width of the run-out distance of the specimen. Thus,

with this long propagation path, this specimen is suitable for comparing

materials for relative resistance to propagation of a crack. The comparison can be made using either measures of energy absorption or lateral

expansion. The latter is measured at the striking line of the specimen.

23
Using the DT test, energy absorption and lateral expansion of longitudinal (LT) and transverse (TL) specimens were measured for each of two

plates of TC-128 steel [9]. Tests were conducted over a range of temperatures that extends well beyond the onset of the ductile upper shelf.

The DT test lateral-expansion data are shown in Figure 2 for one of

these two steels. These results, and those for the other plate that is

not shown, indicate an embrittlement at about 550 °C (1050 °F). The

nature of this embrittlement has not been fully established. In an attempt to explain this embrittlement [10], CVN impact specimens were exposed for 30 minutes at 550 °C and then tested over the range of

near-ambient temperatures at which the transition from brittle to ductile fracture is normally observed. The results of the untempered (or

unexposed) specimens were not significantly different from those for the

exposed or tempered specimens, and therefore it was concluded that this

embrittlement is not due to temper embrittlement of the steel.

What remains are the questions of (1) the mechanism of this embrittlement, (2) the effect of time of exposure on the magnitude of the embrittlement, and (3) the significance of this embrittlement in relation to

service performance.

Further, it is noted that the resistance to crack propagation in the

transverse direction (the lower line in Figure 2) is substantially lower

than that of longitudinal specimens. The direction of crack propagation

represented by transverse specimens is the circumferential direction in

the tank-car shell plates. These observations give rise to another

question: "Can improved transverse properties substantially reduce the

tendency for an undesirable type of circumferential failure in tank cars exposed to elevated temperatures?" This type of failure, shown in

Figure 3, is not uncommon and it leads to rocketing of a section of the

tank car. The rocketing behavior is serious, with apparently higher

potential for personal injuries or property damage. The magnitude of

the decrease in fracture resistance of tank-car steels at elevated

temperatures (the 550 °C embrittlement) appears to be sufficiently large

so that a significant reduction in the resistance to crack propagation

is to be expected at temperatures near 550 °C. Further, because trans- verse properties are much poorer than longitudinal, the potential for

the development of an easier path of propagation in the circumferential

direction of a tank car exists. For these reasons, this work will be

continued.

Cooperative FRA/AAR Research--The NBS studies on tank-car steels are part of the broader research programs conducted cooperatively by the

Federal Railroad Administration (FRA) and the Association of American

Railroads (AAR). As part of this FRA/AAR cooperative effort, tank cars have been subjected to collisions similar to those encountered in rail

yards. These tests were used to understand the events that lead to

failures of tank cars. In addition, statistics of actual impact events

from accidents have been compiled by the FRA in an attempt to determine

the most vulnerable area (or the area most worthy of protection) on a tank car.

24
When tank cars fail in accidents, fires sometime develop. Due to the

serious consequences of these fires, the cooperative research program

has included fire tests of tank cars. As a result of these studies, it was found [11] that the time to failure for a tank car exposed to an all-enveloping fire could be significantly increased by the use of an ablative coating on the plate materials of the tank car. An uncoated

car failed in only 24.6 minutes, just about the time that it takes for a fire company to arrive at the scene of an accident and set up its equipment. A thermally coated car remained for 93.7 minutes before it failed;

this car had vented most of its hazardous cargo (propane) before the

failure occurred. As a result, the explosion that accompanied the

failure of the thermally shielded car was much less violent.

As part of this work, a failure analysis of the insulated tank car, was conducted [12]. This analysis indicated that the failure mechanism

begins with the formation of a stress-rupture (short-time creep) crack

which extends by void coalescence until its length is sufficient for

further extension by tensile overload or shear failure. The stressrupture properties of this plate of TC-128-B steel were characterized [13]. This work points up the importance of the elevatedtemperature properties of tank-car steels. The strength and the notch

tensile strength at temperatures in the range of from 430 °C (800 °F) to

680 °C (1250 °F), as well as the stress-rupture properties for times

within two hours of exposure, are important material properties that

must be considered in the overall balance of properties that are needed

for optimum performance of plate materials used for tank cars in abusive

service.

Improved Safety--As a result of these NBS studies and many other studies

in the cooperative FRA/AAR program, a regulation known as Docket HM144

for Transportation of Hazardous Materials was written into law. This

rulemaking is designed to improve the safety of transportation of flammable compressed gases and anhydrous ammonia. It provides specifications

for thermal shielding, improved couplers, and puncture protection of

head plates in pressurized tank cars of DOT Classes 112 and 114. The

thermal shields may be either a coating of insulating materials or insulating materials encased in an outer steel jacket. The couplers are shelf couplers of the types E or F, which are designed to prevent cars from uncoupling in accidents and thereby to prevent punctures. The

head-plate puncture protection involves options, one of which is the use of head shields that are installed in front of the head plates at both

ends of the tank cars. Their presence increases the relative speed

required to cause fracture from two colliding cars. Thus, the regulation

involves three steps for improved safety: thermal shielding, improved

couplers, and head shields. In addition, relief valve capacity has been

altered for improved performance in accidents. Existing cars will be

retrofitted by the end of 1980 and new cars are built to these new requirements.

25
Future Work--The optimum balance of properties for plate materials used

in tank-car service must be based upon the requirements of normal service

and abusive service. This fact has focused attention on various

approaches to f racture'safe design. One analysis [14] has indicated

that the fracture toughness of plate materials used for pressurized tank

cars is adequate for the normal service requirements. This means that

the flaws inherent in the plate materials and those that are introduced

during fabrication are small in relation to the service stresses and the

fracture toughness of the materials. Under abusive service, large flaws

are punctured into the vessel, stresses can become relatively much

higher, and the propagation of these large flaws can be expected. The

questions are, "What levels of flaw size can be tolerated with the

present grades of steels that are used in this application and are there

alternative steels that would furnish a better balance of properties for

improved fracture resistance at ambient and elevated temperatures?"

The present specifications [3] for steel plate materials used for pressurized tank cars has no fracture toughness requirement, except when the

steel is used for low temperature service—such as for the transportation

of carbon dioxide, which requires steel with 15 ft-lb energy absorption

at -50 °F (-46 °C). Thus, the fracture toughness of the plate materials

used in most pressurized tank cars is assumed to be adequate for fracturesafe design, provided that the steel meets the chemical and other

requirements. The specifications do include a thickness requirement

that is based upon the burst pressure. This is a design requirement

that assumes the plastic collapse mode of failure. Under this failure

mode, an increase in the strength of the material leads to increases in

the burst pressure and in the critical size of the flaw required for

unstable fracture. The flaw sizes for the plastic collapse mode of

failure are greater than those for the elastic or elastic-plastic failure

modes, and this makes the transition temperature of the material an important consideration.

To assure that failure will be by the plastic collapse mode, the ductileto-brittle transition must occur at temperatures below the service

temperatures at which failures are anticipated. When it is not that

low, the analytical problem is to relate the fracture toughness to the

critical flaw size for the elastic or elastic-plastic failure mode under

which failure is expected. When this is done, adequate resistance to

crack extension can be assured by judicious selection of materials on the basis of their fracture toughness properties.

The transition temperatures of materials used in the current pressurized

tank cars are not sufficiently low to always assure upper-shelf plastic

collapse behavior. Thus, failure is to be expected under the other failure

modes, and analytical models are being developed for understanding the

levels of protection that can be expected from the materials now in use and from candidate materials for the future.

26
Hence, the paradox is that both improved strength and improved toughness

may be sought for improved service performance in pressurized tank cars. These properties are normally inversely related, unless cost is of no concern. Therefore, the challenge of the coming years is to furnish

this improved balance of properties, including improved stress-rupture

properties, at a marginal increase in cost of the steel plate materials.

It has been stated [14] that no significant transition- temperature

improvement can be achieved by changing to other ferrite-pearl i te types

of steels. In view of the balance of properties required for this

application, this author would agree with that statement. Perhaps other

types of steels hold promise for the future. Candidate materials would

have to include micro-alloyed steels, commonly referred to as highstrength low-alloy (HSLA) steels.

For the ambient-temperature aspect of the problem, there are two

approaches to fracture-safe design. One is the historic transition

temperature approach which provides safety against unstable fracture by

specifying that the lowest anticipated service temperature always be

above some ductile-brittle transition temperature (DBTT). The DBTT can

be defined by tests such as the drop-weight NDT, the Charpy V-notch, and

the dynamic tear. Since the transition temperature approach does not

directly specify acceptable stress levels or flaw sizes, it is highly

dependent on correlations with service performance.

The second approach to fracture-safe design is based on fracture mechanics and it was initially developed to handle materials that do not

exhibit a ductile-to-brittle transition behavior (high strength steels,

aluminum, titanium). This approach provides the relationship between

the applied stress on a flawed structure and the stress intensity (K) at

the leading edge of the flaw. This approach was initially used in

applications in which plane-strain conditions are developed near the

front of a crack. In tougher materials, this condition does not commonly

prevail and methods have been developed to deal with plane-stress conditions near the crack front. In addition, there are many versions of the

fracture mechanics approach, and a number of fracture toughness parameters are now used, such as the dynamic initiation fracture toughness,

crack arrest fracture toughness, and elastic-plastic fracture toughness.

Attempts have been made to correlate data from the two fracture-safe

design approaches (DBTT and fracture mechanics), and to mix the two

approaches into a hybrid approach such as the ASME reference fracture

toughness curve (Kj^) approach.

Future work on pressurized railroad tank cars will examine past and

current attempts to apply these two approaches to fracture-safe design.

A preliminary effort has already been published [15] by the AAR. It

incorporates the K-rp curve concept and several parameters from the DBTT

approach. This is a beginning. Other studies are in progress under

sponsorship of the AAR and the FRA. An approach being developed at the

27
NBS incorporates direct and indirect measures of fracture toughness.

The direct measures are used in the establishment of the toughness

requirements for the application and the indirect measures are used

through correlation with results of the direct measures, in the establishment of materials qualification standards.

REFERENCES

1. Yearbook of Railroad Facts, Association of American Railroads,

Washington, DC, June 1978.

2. Accident/Incident Bulletin No. 147, Calendar Year 1978, U. S. Department of Transportation, Federal Railroad Administration,

Office of Safety, October 1979.

3. Code of Federal Regulations (CFR), Title 49, Transportation, Part

179—Specifications for Tank Cars, Subpart C--Specif ications for

Pressure Tank Car Tanks (Classes DOT- 105, 109, 112, and 114), U. S. Government Printing Office, October 1979.

4. "Analysis of Findings of Four Tank-Car Accident Reports," C. G. Interrante, J. G. Early, and G. E. Hicho, NBS, FRA OR&D 75-50,

January 1975, NTIS PB 251097/AS.

5. "Impact Properties of Steels Taken From Four Failed Tank Cars,"

C. G. Interrante, NBS, FRA OR&D 75-51, January 1975, NTIS PB 255854/AS.

6. Specifications for Tank Cars , Association of American Railroads,

59 E. Van Buren Street, Chicago, IL 60605, January 1971.

7. "Metallurgical Analysis of a Steel Shell Plate Taken From A Tank

Car Accident Near South Byron, NY," C. G. Interrante and G. E. Hicho, NBS, FRA OR&D 75-47, October 22, 1971, NTIS PB 250063/AS.

8. "A Metallurgical Evaluation of Two AAR M-128 Steel Tank Car Head

Plates Used in Switchyard Impact Tests," J. G. Early and C. G. Interrante, NBS, May 1980, NBSIR 80-2039/AS.

9. "A Metallurgical Analysis of Eleven Steel Plates Taken From A

Tank Car Accident Near Callao, MO," C. G. Interrante, G. E. Hicho,

and D. E. Harne, NBS, FRA OR&D 75-49, September 1972, NTIS PB

250544/AS.

10. "Elevated-Temperature Embrittlement of a Carbon-Manganese Steel

Under Impact Loading Conditions," C. G. Interrante, Dynamic Fracture

Toughness, Proceedings of an International Conference, London,

England, July 1976, the Welding Institute, Abington Hall, Abington,

Cambridge, UK, 1977.

28
11. Phase II Report on Fullscale Fire Tests, J. J. Manda, RA-11-6-31,

Railway Progress Institute, Association of American Railroads

Technical Center, Chicago, IL, December 1975.

12. "A Metallurgical Investigation of a Full-scale Insulated Rail

Tank Car Filled with LPG Subjected to a Fire Environment," J. G. Early and C. G. Interrante, NBS, FRA OR&D 75-52, January 1975, NTIS

PB 250587/AS.

13. "Ambient- and Elevated-Temperature Mechanical Properties of AAR

M-128-69-B Steel Plate Samples Taken From Fire Tested Insulated

Tank Car RAX 202," J. G. Early, NBS, FRA OR&D 76-74, May 1975, NTIS

PB 255907/AS.

14. Final Phase 03 Report: "Material Study on Steels Used in Current

and Former Tank Car Construction and From Cars Involved in Accidents," R. J. Fiber, and L. L. Olson, Report RA-03-5-33 (AAR R-193),

Association of American Railroads Technical Center, Chicago, IL,

August 1975.

15. "Introduction to AAR Guidelines for Fracture-safe Design Involving Temperature-Transition Sensitive Structural Steels," W. S. Pellini, AAR Report R-383, Association of American Railroads Technical Center, Chicago, IL, August 1979,

29
30
TEST TEMPERATURE {*C)

Figure 2. Lateral expansion in dynamic tear (DT) tests

conducted over a broad range of temperatures.

31
Figure 3. A tank car failed at elavated temperatures

showing a circumferential crack propagation.

32
FRACTURE ANALYSIS OF CAST STEEL COMPONENTS IN RAIL VEHICLES

Y. Sharir, D. H. Stone, and W. S. Pellini

Association of American Railroads

3140 S. Federal

Chicago, Illinois 50616

Abstract : Couplers in rail vehicles made of AAR C-grade cast steel have

been tested in the laboratory in simulated, full-scale fatigue tests.

Fracture analyses were carried out on broken components to characterize

cause and mode of failure for comparison with field failures. The ma- jority of the failures initiated at flaws on the surface, mainly casting

defects, characterized to fall in the range of K^=1.5. In the laboratory, cracks initiating at these flaws propagated in fatigue to the

point where catastrophic failure occurred in a brittle mode. This fracture sequence is similar to that found in field failures. A metallurgical analysis which included microscopic examination as well as mechanical testing was carried out. Most of the failed components were composed of normalized steel with pearl itic structures having poor

dynamic toughness and transition impact properties which are believed

to be the major factors leading to failure.

Key words : Cast steels; fatigue crack growth rates; fracture analysis;

mechanical testing; microstructure; rail vehicles; SEM fractography

.

Introduction : The railroad industry is the largest consumer of cast

steels for rail vehicle parts such as couplers, yokes, knuckles, bolsters and side-frames. These parts are made from four grades of cast

steel, differing from one another by heat treatment and slightly by

chemical composition, resulting in a variety of mi crostructures as well

as strength and toughness levels [1].

The couplers, yokes and knuckles are made primarily of C-grade quenched

and tempered steel and C-grade normalized and tempered steel. In 1971-

1972, a joint project of AAR and a few railroads was carried out in

which couplers, yokes and knuckles that had failed in service were examined [2], [3]. Within a period of one year, and with only a few

railroads participating, there were over 4,000 failed parts reported.

Since that time, a shift was recommended towards the C(Q&T) steel which

is believed to show a better fatigue life endurance. A full-scale

laboratory fatigue test, simulating real-life loading history, was conducted to study cause and mode of failure for comparison with field

failures [4]. The subject of this paper is the metallurgical fracture

analysis of these failed components.

33
Experimental Procedure : Within the framework of the laboratory tests,

four couplers, two yokes and twenty knuckles were tested to failure.

Load, cycles to failure, and a summary of test conditions were noted

and a preliminary visual inspection of the failed parts and fracture

surfaces was made [5]. All failed parts were then carefully examined

to reveal fracture origins and defects to be considered as possible

crack initiation sites, and were photographed for documentation. The

failed parts were then sectioned to provide samples for various tests.

Fractography was performed using an ISI-60 scanning electron microscope. Chemical analyses were conducted at random representative locations to check for specification compliance. Microstructures were examined to reveal micro-defects and heat treatment history. Transition

impact strength for both grades was measured using both the Charpy

V-notch and the dynamic tear techniques for 5/8" thick specimens [6],

and also using an SRI double pendulum machine. Fatigue crack growth

rates were measured in accordance with ASTM standards [7] using a closed

loop hydraulic testing machine system (MTS-810). A IT compact tension

specimen failed to provide a valid Kj^ test for either grade at room temperature

.

Resul ts : A summary of test parameters for couplers and yokes is presented in Table 1. The same testing parameters also applied to the

testing of the twenty knuckles that failed.

COITLRRS ]

" YOKES

No. A-OI A-02 A-03 A-04 R-05 B-06 AAR Cast Steel Crade C(N&T) f:(Q&T) C(NiT) E C(N(iT) C(<)&T)

I^ad (kip) +2on +2 on +200/+ TOO + 200 + 300

Cycles (10^) 226.5 1 \ 1 66.5 288. 5

Est Invited 1.1-1.5 1 . 1- 1 . 5 1.5 1.3-1.5 1.5 1.5

Fatigue 2 origins at 3 origins at 2 origins at 2 origins at Multlplp tnltlati'in Multiple Characteristics corner and corner surfaces, corner surfaces. comer surfaces. Joined to form In It la t Ion surface, 80% 1x0.5 Inches 1 . 5 Incites and 0.5x0.5 Inches a huge rrark- Joined to form

of section. slT^e. 0.5 Inch, long. size. 5 Inches long shallow cr.nrk. (80% of section)

.

1 . 5 Inches wide.

Fracture Brittle. Rough texture Brittle, Rough texture Brittle. Rough texture. Appearance with large chevrons with evident shear lips. point Ing shear lips.

to origin.

*Englne coupler, loaded at +200 kip for '.0«,000 cycles, followed by 30,000 cycles at +300 kip to failure.

Table 1: Couplers and yokes, test parameters and fracture

visual inspection.

An attempt was made to identify the locations of fracture origins, and

it was observed that most of the fractures started at high stress con- centration areas, as is schematically shown in figure 1 and demonstrated

in figure 2.

34
YOKE COUPLER KNUCKLE

Figure 1: Location of fracture in couplers (A-03) , yokes

(B-05) and knuckles (D-18)

35
As indicated, most of the failed components showed some amount of

fatigue crack growth. Huge fatigue cracks, as large as 80% of the cross section, can be seen in figure 3, whereas shallow, multiple-origin

cracks are shown in figure 4.

Figure 3: Large fatigue cracks in failed coupler (A-01)

yoke (B-05).

In 90% of the failed components, the cast surface roughness, estimated

to have values of about 1.5, was responsible for crack initiation

sites. Two knuckle failures were associated with huge casting voids, as shown in figure 5. One of these failures (C-18) exhibited a large

fatigue crack with chevron marks pointing very clearly to the fracture

origin.

Chemical analyses at random locations show that all the components were

within the limits of the AAR specifications, as shown in Table 2.

36
COUPLERS YOKES KNUCKLES AAR-SPEC.

Iron balance balance balance balance

Carbon 0 .24 -0.27 0 .27 -0.31 0 .18 -0.31 0.30 max. Manganese 1 .09 -1.48 1 .08 -1.41 1 .26 -1.60 1.85 max. Phosphorus 0 .016-0.017 0 .021-0.023 0 .012-0.027 0.04 max. Sulfur 0 .030-0.031 0 .033-0.038 0 . 020-0.036 0.04 max. Silicon 0 .37 -0.40 0 .40 -0.49 0 .35 -0.46 0.55 max.

Table 2: Chemical composition of failed components.

Metallographic examination to reveal the mi crostructure was carried out

on all components. Figure 6 shows representative microstructures for

both C grades.

The C(N&T) grade, having primarily a pearl itic structure, shows slightly

different patterns of mi crostructure from part to part, depending on cooling rates (TTT diagrams), the manufacturer, and more so, the

position in the heat treatment load. This variability repeats itself

in the CVN data (discussed later). The C(Q&T) grade, having a martensi tic structure, did not show any anomalies. Typical porosity and

inclusions, as seen in figure 6b, were found in most parts and are

37
Figure 6: Typical microstructure of: (a) C(N&T) grade, B-05

yoke, and (b) C (Q&T) grade, B-06 yoke, (X500)

.

regarded as normal. A few parts, as mentioned before, had large, ran- domly spread cavities, but these cavities did not seem to influence test

results. Scanning electron fractography supports the results of the

visual inspection (Table 1), and, as seen in figure 7, shows a brittlecleavage mechanism for the C(N&T) grade and a ductile-shear, microvoidcoalescence mechanism of fracture for the C(Q&T) grade. The particles

and inclusions seen in figure 7a were generally rich in phosphorus,

nickel and manganese, as determined by the SEM energy dispersive X-ray

technique.

A small shrinkage cavity within the fracture path is shown in figure 8.

These cavities ranged in size from 10 mils to over 0.5 inch and show a three dimensional dendritic solidifying pattern (figure 9).

Four knuckles, all C(N&T) grade, and one C(Q&T) yoke provided the

samples for the Charpy V-notch impact tests to determine the transition

behavior (figure 10). It was mentioned earlier that variations in

C(N&T) grade microstructures were observed, probably due to differing

cooling rates. The wide band of energies for lower and upper shelves

for this grade is probably attributable to these variations. However,

the range of transition temperatures seems to stay fairly constant, in

the range of 50-75°F.

The C(Q&T) grade is very much superior in terms of upper shelf, as well

as transition temperature which is in the range of -100 to -50 F. The

percentage shear as a function of temperature is shown in figure 11 for

the C(Q&T) grade.

One coupler, C(N&T) grade, and one yoke, C(Q&T) grade, provided samples

for the dynamic tear tests. The dynamic tear test data, shown in figure

38
Figure 7: Typical fractography of: (a) C(N&T) grade, A-03

coupler, and (b) C(Q&T) grade, B-06 yoke, (X500)

Figure 8: Shrinkage

cavity on a fracture

path, C (Q&T) grade,

B-06 yoke (XlOO)

.

12, support the CVN data and result in values of N.D.T. = +50 F for the

C(N&T) grade and N.D.T. = -100 F for the C(Q&T) grade. Figure 13 shows

the percentage shear for the DT C(Q&T) grade fractures.

39
Figure 9 : A three

dimensional dendrites pattern,

in a shrinkage

cavity, C(N&T)

grade, B-05 yoke

(X150)

.

-200 -ISO -100 '50 0 TEMPERATURE CF)

60 RT 100 ISO

Figure 10: CVN transition impact energy for both grades

40
ov

5B

75V

58

Figure 11: Percentage shear of C(Q&T), B-06 yoke in the CVN

test. Temperature and energy (ft-lbs.) are listed.

Figure 12: DT transition impact

energy for both

grades

.

-100 0 RT 100 ZOO

TEMPERATURE CF)

41
Figure 13: Percentage shear of C(Q&T), B-06 yoke in the DT

test. Temperature and energy (ft-lbs.) are listed.

Compact tension specimens were tested to determine fatigue crack growth

rates. Data from these tests are shown in figure 14. As expected, no significant differences arise between the two grades. In figure 15,

which shows CT fatigue specimens, cavities and porosity can be seen in

dafm \ dn[cfcj

—I 1

\ I I I 1

YOKE B-06, (0aT;

FATIGUE CRACK GROWTH-RATE

to-'* 9

8

7

- 6

- 5

LO6^-a.7fZ.4(L00AK) J I I I I I I I

I I I I I I I COUPLER A-03, (N B T)

lo-s 98

76

54

3

10

/cr6

L0G^'9.0te.6(L0SAX)

I I I I I I I I

do

[CYOJ

20 30 40 60 80 10 20 30 40 60 80

AKiKSl/lU) AK(KSl/TFi)

Figure 14: Fatigue crack growth rates for both grades resulting in almost identical rate equations.

42
Figure 15: CT specimens fracture surfaces of: (a) C(N&T)

grade (A-03) coupler, and (b) C(Q&T) grade (B-06)

yoke

the fatigue path, and ductile tear associated with large shear lips at

the K-zone explains why we could not get valid linear elastic fracture

toughness data, Kj^.

Summary and Conclusions : The extensive work done in the metallurgical

evaluation did not appear to reveal any unusual findings, but rather the

very explicit and well known behavior of the AAR C-grades of cast steel.

It should be emphasized that the statistics of the C(Q&T) grade within

this program are poor and it is really hard to establish a conclusive

summary based on two components. However, as mentioned earlier, the

results seem to agree very well with what is already known and can be

summarized as follows:

1. All test materials appear to be of conventional and accepted quality

for production castings.

2. Cast surface roughness with estimated values of 1.5 seem to be

unavoidable and are regarded as normal. This roughness is probably

responsible for most crack initiation sites.

3. Voids and cavities in the material's bulk do not seem to influence

component performance.

4. Most of the failed parts exhibited a certain amount of fatigue crack

growth, although extremely scattered in terms of number of cycles to

failure. The ASTM fatigue crack growth test reveals identical rates

for both grades (see figure 14). It is therefore really hard to

understand the wide variety of cycles to failure. A few factors

that may have some effect are suggested: (a) it is impossible to

estimate the initial K values and the shape of the initiation site

defects and therefore differences will result in the number of

cycles for the defect to start growing as a fatigue crack, (b) the

differences in microstructure of the different components, as well

as the related scattering of the CVN transition curves for the

C(N&T) grade present probably the most solid explanation for these

variations, (c) loading rates calculated from test frequencies and

peak loads show a range of 8 to 40 KIP/min. Being in the intermediate range of loading rates [8], a factor of five might turn out

to be significant, and (d) there are uncertainties in regard to load

43
history, primarily with possible overshooting on starting or stopping as well as retardation effects due to long breaks within a test

5. Applying Pellini's approach for fracture transition of steels [9],

and using the DT data presented earlier, the following values are developed: for r,(N&T) grade, L=50°F and YC=90°F, and for the C(Q&T)

grade, L=-100 F and YC=-20 F. These values are in good agreement

with what is known and applicable for these grades of cast steels

[9,10].

6. Laboratory experience and service experience appear very similar.

Fatigue crack sizes at failure appear normal for yokes, scattered

for knuckles, and somewhat larger for couplers, probably due to the

absence of occasional impact loading during the tests.

7. Fracture appearance of the C(N&T) grade is typical of brittle fracture as expected and as found in service failures [2,3]. The appearance of the C(Q&T) grade is also typical for ductile fracture,

as expected at room temperature. There is not enough service analysis documentation to support the above observation.

8. It is recommended that extensive fracture mechanics characterization

of these two grades of steel, as well as the other two grades that

are in use - the B normalized and the E quenched and tempered, be

performed for a better understanding of their performance.

Acknowl edgement : Gratitude is expressed to Dr. S. Mostovoy of IIT,

Chicago, Illinois for time and support and very beneficial discussions,

to B. E. Largent for the photographic work, to M. C. Fee for graphics,

and to B. Vaughan for typing the manuscript.

References :

[1] "AAR Manual of Standards and Recommended Practices of Steel Castings," Spec. M-201-79 , Oct. 1979.

[2] "Coupler Failure Investigation (Summer Program)", AAR R-LT-350 , Feb. 1974.

[3] "Coupler Failure Investigation (Winter Program)", AAR R-LT-364 , Feb. 1 974.

[4] "Final Report on Freight Car Couplers Safety Project," Morella, N.,

AAR R-394 , Expected June 1980.

[5] "Preliminary Review of Coupler Data," Pellini, W. S., Railroad

Coupler Safety Research and Test Project, March 1979.

[6] "Standard Test Method for Dynamic Tear Energy of Metallic Materials," ANSI/ASTM E604-77.

[7] Tentative Test Method for Constant-Load-Amplitude Fatigue Crack

Growth Rates Above 10-^ m/cycle," ASTM E647-78T.

[8] "Fracture and Fatigue Control in Structures, Applications of Fracture Mechanics," Rolfe, S. T. and Barsom, J. M. , Prentice-Hall,

Inc. , New Jersey, 1977.

[9] "Fracture Properties of AAR Cast Steels," Pellini, W. S.,

AAR R-272, July 1977.

44
[10] "Evaluation of the Toughness Properties of Various Cast Steels,"

Sigala, P. C. and Wallace, J. F., M.S. Thesis, Case Western Reserve

University, June 1977.

45

SESSION II

RAIL VEHICLES AND STRUCTURES

Chairman: C. G. Interrante, National Bureau

of Standards

REQUIREMENTS FOR ON-BOARD FAILURE

DETECTION SYSTEMS FOR RAIL-VEHICLES

J. V. Mirabella

Federal Railroad Administration

Washington, D. C. 20590

Abstract: On-board failure detection systems for rail

vehicles are proposed as a means of reducing train accidents

particularly in the high speed range. In this paper, accidents are broken down in broad categories of cause and speed

range. Special emphasis is given to equipment caused accidents as a percentage of total accidents for all speeds and

then for 3 speed ranges--10 mph and below, 11 mph-30 mph

,

31 mph and above. When the accidents occurring at speeds of

31 mph and above are isolated, equipment caused accidents

represent a larger portion of the accidents than when the

analysis includes the lower speeds . Isolating component

failure accidents, a similar trend is noted with certain

components causing a very small percentage of accidents at

lower speeds, but causing an increased percentage of the

accidents at higher speeds . These components are discussed

in terms of past research with on-board sensors as well as

possible candidates for future research and development.

Various approaches are discussed to examine methods to use

the sensors in a system for possible safety benefits. The

advantages and disadvantages of each type of system are

discussed

Key words: Contact derailment sensor; g-sensing derailment

detector; local derailment; nitinol sensor; on-board failure

detection system; overheated bearings; thermal switch

sensor; train line.

INTRODUCTION

Before evaluating on-board failure detection systems for

rail vehicles, the problem must be initially defined in

terms of the accidents which could possibly be prevented.

This paper, therefore, will first discuss railroad accidents

as reported to the FRA for the last 10 years, and then examine various on-board failure detection systems for rail

vehicles for their advantages and weaknesses.

49
ACCIDENT DATA ANALYSIS

The total number of accidents reported to the FRA during the

period of 1968 to 1978 is shown in Figure 1. These accidents represent all those involving a rail vehicle which

incurred damages at or above a monetary reporting threshold!^

Although yearly fluctuations are present, an overall rise

in accident numbers can be seen between 1968 and 1974.

While in 1975, it appears that the number of accidents

dropped, it must be noted that a change in reporting requirements was made in that year and affected the statistics accordingly. The increasing trend resumes between

1975 and 1978 and merits investigation as to the cause.

In 1978, of the total number of reported accidents, 11,277

(77.77o) were derailments; 1,477 (13.1%) were collisions;

and 1,037 (9.2%) were other types of accidents. VThile the

following analysis of speed and damage factors will include

all of these accident types, it should be kept in mind that

the predominant type of accident is the derailment category.

Classification of 1978 accidents according to speed (See

Table 1) reveals that approximately two-thirds (67.87o) of

all accidents occur at speeds of 10 mi/hr (16 km/hr) or

less. This speed range, however, represents only 30.5% of

the dollar damage. Conversely, the speed range 30 mph and

above accounts for only 10.17o of the accidents, but 32 . 37o

of the dollar damage. The relationship between damage

incurred and speed is the result one would expect and is

further shown in Figure 2 which shows damage per accident

as a function of speed.

Train accidents can also be classified according to their

causes. For the year 1978, accidents broke down into the

following categories and percentages:

Table 2 further delineates this information into speed

ranges versus cause. In analyzing this data, it is noted

''^For the years 1968-1974, the reporting threshold was

$750; in 1975-1976, it was increased to $1,750; and for the

years 1977-1978, it rose again to $2,300.

Track caused

Equipment caused

Operating practices

Other

42 . 5%

19 . 3%

25.2%

13.0%

50
51
ACCIDENTS-DAMAGE

BY SPEED RANGE

SPEED y« OF % OF

(MPH) ACCIDENTS DAMAGE

0-10 67.8% 30.5%

11-30 20.3% 36.1%

31 & ABOVE 10.1% 32.3%

Tnble " 1

52
ACCIDENT SPEED - $ DAMAGE

10 20 30 40 50 60

SPEED (MPH)

J

Fir, # 2

53
54
that while track causes represent 48.2% of the low speed

accidents, they represent only 20.57o of the high speed

accidents. Inversely, while equipment causes result in

only 12.47o of the low speed accidents, in the high speed

accident range, equipment causes result in 48.0% of the

accidents. As shown earlier, the higher speed accidents

result in the most amount of damage. It can be asserted,

therefore, that equipment caused accidents represented

nearly one-half of the most serious train accidents in 1978.

Causes relating to component failures are further delineated for more specific examination using standard FRA

cause codes such as

• Frogs, switches & track appliances

• Axle bearings

• Rails and joint bars

• Wheels

o Couplers and draft system

• Brakes

• Side bearings

Based on train accident data for the years 1977 and 1978,

the leading causes of accidents due to equipment failures

were identified. Tables 3, 4 and 5 rank the equipment

failures which caused the most accidents for each of the

speed ranges. It is interesting to note that the primary

cause of accidents in one speed range may not show up at all

as a major cause at another speed range. For example, axle

bearing failure is the leading cause of train accidents in

the high speed range of 30 mi/hr and above. However, this

classification of cause does not appear as a major cause

in the low speed range of 10 mi/hr and below.

It is apparent that the leading causes of accidents at the

most serious , higher speed ranges are vehicle related and

these are the prime candidates for failure detection systems .

DETECTION SYSTEMS

One possible method of providing for prevention of these

equipment failures is through the use of on-board failure

detection systems, On-board detection systems could, by

early detection of impending equipment malfunction, prevent

a large number of the serious accidents which occur on

railroads

.

Most past research concerning use of on-board sensor systems

has involved detection of overheated bearings (hot boxes)

55
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56


and the occurrence of local derailments On-board sensors

can be developed much further, however, for other train

applications, such as monitoring of:

• overheated or cracked wheels

• excessive rock and roll

• hazardous material leakage

• potential damage to special cargo

• unreleased hand brakes

.

In general, the complete on-board detection system is made

up of three segments: (1) the on-board sensor itself;

(2) a method for transmitting the signal emitted by the

sensor and (3) a correcting response for avoiding an impending accident.

Qn-Board Detection Sensors

To date, research has been conducted on on-board sensors

which are designed to detect hot boxes on freight cars.

Figure 3 illustrates one particular type of a hot box sensor

which utilizes a wire made of nitinol, a unique metal with

the property of changing shape when a certain temperature

is reached. In this case, the nitinol ring will spread

when the metal reaches 250 F. As the ring spreads, a

firing pin is released which in turn activates a thermal

battery and a signal is emitted.

A second type of overheated bearing sensor involves the use

of a thermal switch (See Figure 4) . When the switch

reaches a given temperature, the system is activated and

sends a radio signal to the engineer. The unique aspect of

this sensor is that it is in the form of a bolt and replaces one of the actual bolts which secures the bearing

end cap.

In terms of detecting local derailments and taking corrective action to prevent a major derailment, the following

two sensors have been developed. One such device is a

g-sensing derailment detector (See Figure 5) . This truck

mounted device senses "g" levels in a vertical direction

and when the seismic mass is sufficiently displaced upon

"A local derailment is defined as a derailment of a single wheelset.

59
60

62
derailing, a release mechanism allows the signal power

supply to be initiated. During the testing of this device,

it was found that normal "g" levels encountered over rough

track in the winter could be higher than that caused by an

actual low speed derailment in the summer. This would mean

that the sensors would be required to be adjusted frequently to accomodate environmental changes or it would result

in false alarms.

In an effort to avoid the disadvantage associated with the

seismic sensor, a second type of local derailment sensor

has been investigated. This sensor, shown in Figure 6, is

based on contact with the rail upon derailment. With this

sensor, a disc on the bottom of the sensor attached to the

wheel opposite the derailed wheel will make contact with

the rail. The disc is attached to a plunger type of mechanism which severs a shear pin and activates a pulse battery

which emits a signal.

Signal Transmission

There are basically two ways

may be transmitted to the cr

line. There exist drawbacks

however, as discussed below.

in which the sensor's signal

ew; by radio or by a train

to both transmission methods.

First of all, there is the problem that the locomotive

heading the train consist must have the proper receiver to

match the emitter on the particular freight car. In view

of interchange requirements, it is not practical to key a

freight car to a specific locomotive. Moreover, all freight

cars and locomotives would be required to operate on the

same radio frequency. This may cause problems if the signal

is not sufficiently directional. When trains are in close

proximity to one another, for example, as in a yard operation, the wrong locomotive may receive the signal. . Contemplation of a train line transmission system must

start with the understanding that train lines, at present,

do not exist on freight cars. An extensive modification

program to install the on-board system or train line alone,

therefore, would be required to implement such a plan.

Moreover, should the line on one freight car malfunction,

the entire system would go down. This alone may dictate

that should the transmission of the signal to the crew be

desired, radio transmission would be the more practical

approach

.

63

System Corrective Response

Once a problem arises and the sensor is activated, there

are two options as to what corrective response action can

take place. One type of system response has already been

discussed in that the signal is transmitted to the train

crew in the locomotive and the crew will then take appropriate measures to correct the problem. A second method

of response entails automatic activation of the brakes , to

be accomplished with a device such as that shown in

Figure 7. This device, when activated by an electric signal, vents the brakeline at a predetermined rate. The

overall automatic system is shown in Figure 8.

At present there is no clear preferred approach as to

corrective response action, as both present their own

advantages and disadvantages. The automatic system requires no decision and brake action is immediate. There

are arguments, however, that there are situations in which

immediate brake action is inappropriate and may actually

result in a more severe derailment. There also exists

opposition to any modification to the brake pipe as required in the automatic system. On the other hand, with

an automatic system, a specially equipped locomotive is not

required, thereby facilitating portions of one fleet to

function safely when used on other railroads. As in the

case of radio transmission, the automatic system's benefits

are proportional to the number of cars equipped.

In support of transmission of the signal to the crew, there

exists the capability to transmit noncritical information

which does not require immediate action. For example,

information such as detection of rail flaw could be transmitted to the engineer who could then relay that data to

the dispatcher. An automatic system, being limited to a

predetermined brake application, could not serve this

purpose

.

CONCLUSION

As indicated above, each type o tages . Since each system, if p

accomplish the safety function,

necessarily consider economics,

information is not available on

reliability of on-board failure

research and evaluation will be

information.

system has its own advanoperly developed, could

further investigations must

At this time, adequate

cost, potential benefits or

detection systems . Further

required to provide this

65
66
67
REFERENCES

Federal Railroad Administration, Accident/Incident Bulletin

No. 147 . Calendar Year 1978.

Nayak, P. R. and Palmer, D. W. "Issues and Dimensions of

Freight Car Size: A Compendium" FRA/ORD-79/56

Smith, Richard L. and Frarey, John L. On -Board Failure

Protection Requirements for Railroad-Vehicle Equipment . FRA/ORD-78/72

68
DETECTION OF RAIL DEFECTS AND PREVENTION

OF RAIL FRACTURE

Oscar Orringer and Harry L. Ceccon

DOT Transportation Systems Center

Cambridge, Massachusetts 02142

Abstract: Rail transportation safety is maintained in part by

periodically conducting a continuous search of track to find and

correct rail defects. The state of the art of this detection

technology is reviewed to illustrate, in general terms, the capabilities available to the railroads. A joint government/industry effort

is now in progress to determine whether the periodic inspection

strategy can be improved by means of an adaptive control algorithm

that permits variation of the inspection interval based on actual

histories of detected defects. The results of some preliminary

studies are presented to illustrate how simulation techniques can be

used to assess the effectiveness of proposed changes of inspection

strategy.

Key words: Crack detection; inspection interval; rail flaw

detection.

BACKGROUND

About five hundred derailments occur annually in the United States

as a result of rail fracture. Although many of these derailments

have only economic consequences, public and/or train crew safety can

be affected by occasional accidents involving high-speed operations

and/or hazardous material transport. The government and the railroad

industry thus have parallel interests in the goal of reducing derailments caused by rail fracture.

Since the complex behavior of crack makes it impossible to predict

the times and locations of fractures, control is established primarily

by means of periodic inspection with detector cars specifically

designed to find rail defects while they are still growing slowly in

a subcritical state. The railroads and rail inspection services, as well as the government, have actively engaged in research to improve

subcritical flaw detection capability. The flaw detection capabilities available to the railroads are outlined in the following section.

It will be seen that that technology is imperfect, viz: detection

cannot be guaranteed for large defects approaching fracture, even though there is some chance of detecting small ones which are still

in the early stages of growth; also, the ability to accurately

classify defects by size is limited.

69
In the railroad environment, these imperfections in flaw detection

can lead to an overcommitment of scarce maintenance resources in order

to repair or remove small defects. Since rail defects generally appear

to grow very slowly, it would seem advisable to concentrate maintenance on the larger defects approaching fracture, if an effective

strategy could be found to achieve this goal

The current Federal regulations require a minimum of one detector-car

inspection annually for certain classes of track. However, many

railroads inspect certain sections of track more often, and all railroads back up the detector-car search with much more frequent visual

inspections. This existing approach to control rail defects will be

compared to an adaptive control scheme which is currently the subject

of a joint government/industry study. The study objective is to

determine whether it is possible to rebalance resource commitments

by varying the inspection interval based on actual counts of detected

defects.

The effectiveness of adjusting inspection frequency based on defect

count can be assessed by means of a model which simulates the dynamics

of a population of defects that have been generated by track usage,

grow under further usage, and are subjected to periodic detection.

Results from some preliminary studies with a simplified population

dynamics model will be presented to show that the adaptive strategy

is a promising approach to control of rail defects. This paper con- cludes with some remarks about the possible effects of defect spatial

distribution on the effectiveness of the adaptive strategy and a summary of plans for future research.

DETECTION VEHICLE TECHNOLOGY

The fleet currently inspecting the nations 's railroads consists of 72

rail flaw detection vehicles. Forty-six of these are owned and operated by individual railroads and the remaining 26 by inspection

services. The inspection services provide service for railroads who

do not own their own equipment, and supplement the inspection capabilities for railroads who have inspection vehicles. Only one railroad currently has the capability to satisfy its inspection needs

without help from an outside service.

During 1979, the fleet logged over 300,000 miles of track inspection.

This inspection mileage was almost equally divided between the railroad fleet and the inspection service fleet. The total mileage inspected consists of both single (once per year) and multiple inspections. The single, annual inspection complies with Federal Track

Safety regulations. Individual railroads use multiple inspections

to insure high safety and maintenance standards on lines carrying

high tonnage.

70
The inspection methods used on the vehicles include ultrasonic, magnetic or a combination. Twenty- five of the combination systems are owned and operated by one of the inspection services and nine are owned by five of the railroads.

The overall performance of the systems (basically speed and flaw

detectability ) varies quite significantly. This variation can be

attributed primarily to the I'arge number of different designs and con- figurations used and the range of signal interpretation skill possessed

by different equipment operators. The mode of operation used in the

United States is to hand verify all flaw indications sensed by the

detection vehicle (Figure 1) . There is heavy dependence on human

judgement in both operations . The operator must interpret the test

data to determine if the system flaw indication warrants further investigation and whether the results of the hand test, using ultrasonic techniques, indicate a rail flaw requiring remedial action.

FIGURE 1. HAND MAPPING FOR DEFECT VERIFICATION

Approximately 12 basic designs appear in the national fleet. Various

combinations of sensor arrays, data processing and display schemes are

used. Sensor array designs vary from three sensors per rail on either

ultrasonic or magnetic systems to as many as 15 on combination systems

71
Ultrasonic transducers are coupled to the railhead through either

sled or wheel carriages, using a variety of fluid couplants.

Figure 2a represents the array with the fewest number of transducers

used in the industry. Figure 2b shows a tandem wheel configuration

with additional transducers oriented to enhance the detection of

certain types of defects and to provide coincidental (multiple channel)

detection of other defects to improve the signal-to-noise ratio and

thus increase the probability of detection. Even with the tandem

multiple transducer sensors, defects can be located and/or oriented

so as to not reflect or block probing signals, and can thus escape

detection. Another major contributing factor which limits the detectability of ultrasonic systems is the problem of coupling acoustic

signals into and from the rail. Grease from oilers on curved track,

or rail-head surface anomalies can prevent acoustic signals from

entering the rail. The combination of the technology limitations and

typical operating conditions make today's inspection operations imperfect.

The processed inspection data is presented to the operator in one of

two ways: a pictorial B-scan presentation which is found in 22 vehicles

of the fleet; or variations of gate/threshold processing coupled to a proper chart. Figure 3 compares the two presentation methods.

The systems using the B-scan presentation have an advantage over the

gate/threshold systems in that more descriptive data is presented to

the operator, allowing quicker and better characterization of the flaw.

However, the B-scan systems currently in the field have no data screening capability. Consequently, the large volume of data presented to

the operator can limit the maximum inspection speed. Maximum inspection speeds of up to 15 mph have been attained, but the fleet average

inspection speed ranges from 2 to 6 mph. Inspection speed is heavily

dependent on the condition of the rail under inspection and the number

of stops required to hand verify flaw indications

Both hy-rail and rail-bound vehicles (Figure 4) are used in the rail

inspection industry to house the inspection equipment. The trend in

the railroad fleet has been toward the use of more hy-rail vehicles.

This type of vehicle offers more mobility and, as a result, less interference with traffic when operating in an inspection mode or when

in transit between inspection sites.

There are 43 hy-rail vehicles and 3 rail-bound cars currently in the

railroad fleet. Rail-bound cars are used almost exclusively in the

inspection service fleet (25 out of 25) . The larger vehicles provide

space for the crew to live on-board and to house the heavy magnetic

induction equipment used by the service.

72
(A) SINGLE ULTRASONIC WHEEL PROBE

(B) TANDEM-WHEEL ULTRASONIC PROBES

FIGURE 2. TYPICAL FLAW DETECTION WHEEL SENSOR ARRAYS

73
CHARr lAAL-i-i

OIFICTI¥f Wit

i SCAN

HECOflWNG

FIGURE 3. COMPARISON OF CHART RECORDING AND B-SCAN

PRESENTATIONS

The trend within the railroad fleet over the past 10 years has been to

improve system capabilities by increasing the number of sensors. This

is being accomplished by adding transducers to the ultrasonic system

or using ultrasonic and magnetic systems in combination on the same vehicle. The ultrasonic and magnetic inspection methods nicely com- plement each other to provide a more thorough inspection than either

system operating independently. Magnetic systems have good detectability for defects located in the upper portion of the rail head,

particularly near the surface where ultrasonic systems are deficient.

Ultrasonic systems are effective in the lower portions of the head,

and in the web region where magnetic systems have almost zero detection

probability. The major service fleet has used a combination system

for the past twenty years; nine combination systems are currently

operational in the railroad fleet.

It is apparent that a wide diversity of detail design and configuration exists in the national fleet, although the major inspection

service maintains a uniform design throughout their 25-car fleet, and

the owner-railroads are working toward the same objective within their

own fleets. Since the introduction of ultrasonic inspection systems

74
(a) rail bound detector car owned and operated by sperry

RAIL service

(B) HY-RAIL detector car owned and operated by BURLINGTON

NORTHERN RAILROAD

FIGURE k. TYPICAL DETECTION VEHICLES

75
in the U. S. in the early 1960's5 there has been no significant change

in the "basic design used by the industry. This particular ultrasonic

application resulted partly as a spin-off from German technology. The

magnetic rail inspection systems are basically the same design and

configuration used in the 1930' s.

Some work is currently in progress in the industry to adapt high-speed

data processing to rail inspection. This work has been primarily

directed toward aiding equipment operators by screening data and

presenting the inspection results in a form that can be quickly and

reliably interpreted. The results of this work are planned to be

applied in the field within the next year.

Current government developmental work on automating inspection equipment has aimed at increasing inspection speed and reliability by

relieving the equipment operator of the need to manually adjust con- trols or to make operational decisions. Results of this work include

the development of automatic controls to adjust amplifier gain, adjust

signal gates and maintain sensor alignment. The automatic amplifier

gain control compensates for variation in acoustic signals caused by

changes in the metallurgical properties of rail steel and sensor-torail coupling conditions. The gate control device automatically re- adjusts signal gates, when the heights of rails under inspection change,

without having to stop the vehicle. The automatic sensor alignment

device controls the lateral position of the sensor relative to the rail

head to insure that the sensor is aligned directly over the acoustic

center of the rail. This alignment device compensates for sensor misalignment caused by worn rails or changes in rail head size. These

automatic controls have been tested and are currently in use in the

field.

The detectability and repeatability of rail inspection operations depend

primarily on two basic factors: equipment performance and operational

conditions. The performance variability within the fleet is probably

proportional to the number of different designs and configurations.

Also, changing environmental conditions cause performance variation,

even in a single equipment set. Finally, inspection results are directly dependent on a human operator. Studies of the performance of

operators have revealed a definite variation over a given time period.

Common practice in the rail inspection industry is to rotate operators

within the crew to prevent fatigue and boredom.

Although basic equipment performance can be improved through R&D and

operator skills can be enhanced by training, the environment on revenue

track is virtually uncontrollable. Rail surface conditions are probably the most significant. Shelling rail and engine burns, for example, can mask underlying defects from acoustic or electromagnetic inspection

systems. Dirty rail surfaces can prevent acoustic signals from entering the rail, rendering ultrasonic systems completely ineffective.

Weather conditions, particularly extreme cold, have a definite effect

76
on the performance of all types of inspection operations. Track

geometry deviations have a significant effect on sensor positioning

and thus influence the effective inspection speed. Thus, it is likely

that mobile rail inspection will remain an imperfect art in the future.

CONTROL OF RAIL DEFECTS

It is apparent from the foregoing review that net populations of undetected defects must be assumed to remain in track after periodic mobile

inspections. Railroads control these populations by means of frequent

visual inspections designed to spot large-sized defects and broken

rails as they occur. This second line of defense remains effective as long as it is not overwhelmed by large quantities of these critical

defects

.

Minimum inspection requirements embodied in the current Track Safety

Standards take an open-loop approach to control of the net population,

as shown schematically in Figure 5. The simplicity of this approach

is administratively advantageous. The scheme is technically disadvantageous because it requires that inspection resources be committed

uniformly, regardless of variations in the numbers of defects occurring

at different track locations, unless a railroad possesses the resources

to conduct more than the minimum required inspections

.

DEFECTS

GENERATED

BY TRACK

USAGE

INSPECT

AT FIXED

INTERVAL

DETECT NET

DEFECT

POPULATION

DETECT

BY

VISUAL

INSPECTION

t

REMOVE

FIGURE 5. CURRENT APPROACH TO CONTROL OF RAIL DEFECTS

77
The Federal Railroad Administration and the American Railroad

Engineering Association are investigating a modified scheme for defect

control. The results of detector-car inspections are to be fed back

as an adjustment to the inspection interval as shown schematically in

Figure 6. Here it is assumed that a certain generic defect type is

"critical", in the sense that derailment incidence is a function of the

population level for the given type . In addition to removal or other

remedial action, the detection process is now followed by classification to identify the count of the critical-type defects that have been

detected. The count for critical-type defects is then used to adjust

the interval to the next inspection.

The proposed approach is administratively complex, but it offers a potential for flexible allocation of resources away from "well behaved

track" toward track that is beginning to generate increasing populations of defects. It is hoped that this closed-loop approach will lead

to reduction of the net defect population without any significant increase in total inspection resources. However, it remains to be seen whether the inspection interval adjustment scheme is practical and

whether it will allow a railroad to react to changing track usage.

DEFECTS INSPECT DETECT

GENERATED AT

BY TRACK VARIABLE

USAGE INTERVAL

APPLY

RULE FOR

INTERVAL

ADJUSTMENT

CLASSIFY

DEFECT

TYPE

NET

DEFECT

POPULATION

1

1

f REMOVE

(ALL TYPES)

DETECT

BY

VISUAL

INSPECTION

I

REMOVE

FIGURE 6. PROPOSED APPROACH TO CONTROL OF RAIL DEFECTS

78
The latter question is the focal point of a comprehensive study, just

getting underway, in which the defect occurrence statistics of two

Class 1 railroads will be used to formulate a simulation model for

testing the technical effectiveness of the proposed control scheme.

The simulation studies will investigate flexibility available to the

railroads. In the following sections, results will be presented from

some preliminary studies in which the close-loop concept was investigated to first order by neglecting effects associated with the distribution of defects along the track.

DEFECT POPULATION DYNAMICS

The reaction ability of an inspection interval adjustment rule can be

assessed by means of a population dynamics model which simulates a nominally uniform section of track. Thus, the only independent variable in the simulation is track usage, represented by traffic tonnage

in million gross tons (MGT)

.

Figure 7 summarizes the conceptual elements of a generic simulation

model applicable to a single defect type. The continuous range of

defect size is divided into a finite number of subranges ("bins") to

permit detailed tracking of the population. These bins extend from an "initiation size", Sq, based on metallurgical laboratory data, to a "critical size", Sqr, based on fracture mechanics and/or analyses of

service defects. The range from S© to S^j^ spans the macrocrack growth

regime. In particular, S^^j^ is set such that rail breakage can be

expected within one or a few train passages after a defect reaches this

size.

An empirical initiation model feeds defects into the first bin as a function of accumulating tonnage. A macrocrack growth model then

distributes these defects across the other bins as a function of tonnage accumulated since initiation. As the defects are moved to larger

sizes, a "minimum detectable size", Sp, is encountered. For sizes

exceeding Sq, there exists a finite probability for detecting the

defects by means of the detector car. The model accounts for periodic

inspection by removing appropriate fractions of the defects in each

detectable bin. The sum of these counts Dq, is then considered to be

available for use in adjusting subsequent inspection intervals. Those

defects which escape detection to reach size ScR are accumulated into

a subpopulation representing the burden on the backup visual inspection

effort.

APPLICATION TO DETAIL FRACTURE

Use of the population dynamics model will be illustrated with data

which approximate very roughly the behavior of detail fractures.

Figure 8 illustrates a detail fracture that grew in a rail in the

Facility for Accelerated Service Testing (FAST) at the DOT Transportation Test Center. The detail fracture is considered to be a critical

79
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80
FIGURE 8. EXAMPLE OF DETAIL FRACTURE AT TTC FACILITY

FOR ACCELERATED SERVICE TESTING

(Note: Detail fractures in revenue track generally exhibit

a smooth fracture face, rather than the "clamshell" pattern

shovn here.

)

0 200 400 600 800 1000

CUMULATIVE TRAFFIC (MOT)

FIGURE 9. TYPICAL DEFECT GENERATION BEHAVIOR

81
defect type because its location inside the rail head makes it more

difficult to detect than defects which grow in a surface cracking mode.

Figures 9 through 11 summarize the data that were used in the

simulations. Figure 9 depicts a generic curve for the cumulative num- ber of defects per mile, D generated by aging track:

3 ~ 8 3

D = CT (C = 1.25 X 10 Defect/Mile - MGT ) (1)

where T is the accumulated tonnage in MGT.

Equation 1 is an empirical fit to aggregated data for many different

defect types, and is thus not strictly applicable to detail fractures.

In this case, however, the best available data have been adopted for

illustrative purposes only. The initiation model derived from Equation

1 is the count of defects, AD, which appear during the interval, AT

between inspections:

AD = — AT = 3CT At (2)

The defect size, S, measured as a percentage of the rail head area (HA)

is assumed to increase exponentially with the tonnage accixmulated after

initiation. The regime of macrocrack growth is assijmed to span the

following range:

Initiation size, S = 1% HA

o Minimum detectable size, S = 5% HA

D

Critical size, = 90% HA

CR . - (The value of S is probably somewhat high. Figure 10 depicts that

portion of the defect growth curve between S and S_„. ^ D CR . Defect detection is simulated by a generalized detection propability

curve of the form

P(S) = 1 - exp

M

-f^j & (1 - V^^)

I

(3)

where P^^^ = P(S^j^) and M is a dimensionless parameter used to change

the shape of the detection curve and the average detection probability.

A measure of overall detection propability is obtained by averaging

P(S) across the size range:

^CR . . P „ = (S - S ) ~ f P(S)dS (4) AV CR D /

82
on<

et

LU LjJ M ZC

I—

t

OO _J

I—

I

I— <

LU O

on

0 10 15 20 25

CUMULATIVE TRAFFIC (MGT)

FIGURE 10. ASSUMED DEFECT GROWTH BEHAVIOR

o

I—oLlI hLU ^

Q I— O

>- LU I— Q.

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00OOdQ.

100

80

60

40

20

0

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"

1/ ^AV

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1 1 1

^AV

"

1 •

88%

20 40 60 80

DEFECT SIZE

(PERCENT OF RAIL HEAD AREA)

100

FIGURE 11. ASSUMED PROBABILITY OF DEFECT DETECTION

83
Figure 11 depicts two assumptions for P(S) corresponding to the

following parameter values:

P = 95% M = 0.5 P = 82%

CR AV

P = 95% M

= 0.25 P, = 88%

CR AV

Comparative analyses with these cases will be used to illustrate the

effect of improving detection probability.

SIMULATION RESULTS

The foregoing data were used to make a simulated test of the effectiveness of the following scheme for inspection interval adjustment. The

initial inspection interval is set at AT = 20 MGT, which lies within

the range of intervals currently used on typical revenue track. Whenever the detected defect count, D^, (see Figure 7) exceeds a trigger

value, D^, the succeeding inspection interval is reduced to AT' in

accordance with:

Thereafter, the inspection interval is fixed at AT' until again

exceeds D^.

The simulation begins at zero accumulated tonnage, and is continued

lantil the inspection interval decreases to 1 MGT. Figure 12 depicts

the results for the baseline case (inspection with P^^ = 82% and trigger = 0.16 defect per mile). Figure 12a shows that interval reduction begins at about 460 MGT. Thereafter, the detected defect count

oscillates as the effects of inspection interval reduction and defect

population pressure alternately dominate. The simulation reaches 1,000

MGT of accumulated traffic when the inspection interval reaches 1 MGT.

The remainder of Figure 12 depicts two possible measures of the burden

placed upon backup visual inspection resources. Figure 12b shows the

accumulated number of defects per mile which manage to escape detection

and thus enter the pool of imminent breaks and fractures . This curve would have continued to rise at an accelerating rate, had the inspection interval not been adjusted. However, the adjustment rule even- tually brings the breakout pool under control at a level of 0.58 defect/

mile. Figure 12c illustrates the number of subcritical defects which

will enter the breakout pool during each inspection interval. This

parameter rises to about 0.038 defect/mile, but thereafter decreases

rapidly as the inspection interval is reduced.

The baseline results will now be compared with two variations: a 7-percent improvement in average detection probability (Pj^y = 88% instead of

82%); or a 25-percent reduction of the trigger (D^ = 0.12 instead of

0.15 defect/mile). Figure 13 depicts the effects of these changes on

At' = AT/D. (5)

84
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86
inspection management. The inspection interval reductions begin

earlier than the baseline in both cases. Improved detection with

baseline trigger appears to maintain about the same control on defect

count as the baseline. On the other hand, baseline detection with re- duced trigger controls the defect count to a much lower level.

One might conclude from the foregoing results that trigger reduction

is the more effective of the two improvements. However, the other

simulation outputs reveal that 7-percent improvement in P^y is actually

more effective than 25 percent reduction of D^. Figure 14 shows that

the visual backup inspection burden is alleviated much more by the detection iitprovement than by the trigger reduction.

DISTRIBUTION OF DEFECTS ALONG TRACK

It is unrealistic to expect that defect occurrences will be distributed

uniformly, in a statistical sense, over arbitrarily long sections of

revenue track. Also, it is \inlikely that the segmentation of track for

defect accounting purposes can be made to correspond precisely to the

boiandaries between track sections having different inherent defect

rates. Under these conditions, it is prudent to investigate the effectiveness of defect frequency dependent inspection when defects are non- uniformly distributed along the track. Work on spatial modeling has

been started, but no complete simulations have yet been conducted.

The first step in the investigation was taken by analyzing several

years of field data from a two hxmdred mile section of revenue track.

The analysis has indicated that a Poisson distribution might describe

the count of defects detected per 10-mile segment. The family of

Poisson distributions obeys a unique linear relationship,

6^ = 62 - 3 (6)

between the shape descriptors of its probability density:

(skewness) (7)

(kvirtosis) (8)

where a'', and are respectively the second, third and fourth cen- tral moments of the distribution.

Figure 15 corrpares the Poisson model relationship with estimates of

^1' ^2 co'^P^ted from the field data, and from a simulation run with

a random number generator known to produce Poisson data. Comparison of

the simulation and field data indicates that the field data scatter can

be attributed to a finite sample-size effect. Therefore, the Poisson

distribution appears to provide an appropriate spatial model.

87
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1 1 1

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KURTOSIS, 32

FIGURE 15. SHAPE DESCRIPTORS FOR POISSON DISTRIBUTION

AND DATA HISTOGRAMS

An additional and independent rationale for ad(£)pting this model is- that

Poisson data is distributed with no bias in location, i.e. the appear- ance of a defect in one rail is equally likely as its appearance in any

other rail. This type of characteristic appeals to the statistician as a working hypothesis, in the absence of any definite evidence to the

contrary. The Poisson model will be used in the future to conduct more comprehensive analyses of field data, as well as to test the sensitivity

of closed- loop control to spatial nonuniformity

.

CONCLUDING REMARKS

Preliminary results have been presented from studies based on a simplified mode L of rail defect population dynamics. The simulation included

a closed- loop scheme for controlling the defect population by varying

the inspection interval in accordance with the findings of past inspections. The results indicate that the closed- loop concept is a promising

approach to a better allocation of inspection and maintenance resources.

The opportunity for better allocation arises because the inspection

interval might be lengthened during the early track life stage, when a backup visual inspection program can cope with the defect population.

In other words, it may be possible to transfer mobile inspection re- sources to the later stages of track life, when the defect population

begins to rise rapidly. The preliminary study has also shown that the

simulation answers can be sensitive to underlying assunptions in the

89
model. A specific illustration of sensitivity to detection probability

has been presented. The model undouhtedly possesses similar sensitivities to the rate of defect generation, the rate of defect grovth, and

the characteristics of defect distrihution along the track. Because of

these sensitivities, it is obvious that better input data will be

needed before such simulations can be used to support decisions about

the management of rail defects in revenue track. The wide diversity

of detector car equipment and operating conditions requires that detection probability bounds be established for the purposes of the simulation studies. Also, detailed information is needed about the occurrence histories, spatial distributions and growth rates of specific

types of rail defects.

The Federal Railroad Administration and the American Railway Engineering Association are currently engaged in a cooperative effort to improve track performance standards. A major component of this effort

is devoted to the collection of refined data on rail defect behavior.

The task of better characterizing defect detection probabilities is

still in the planning state. However, work on the other aspects is

underway. Defect occurrence and spatial distribution statistics are

to be developed from detailed studies of the defect histories on two

Class 1 railroads. Defect growth rates are to be measured in a series

of tests in which flawed rails removed from revenue track will be

subjected to accelerated service testing at the EXDT Transportation

Test Center. The results of these investigations will ultimately be

combined in realistic simulations to assess options for improving

track safety.

90
AUTOMATED NDE FOR DETECTION OF BRAKING

ABNORMALITIES OF TRAINS

J. D. Ferguson

Federal Railroad Administration

Washington, D.C. 20590

Abstract

Wayside instrumentation was designed, fabricated, and tested that

demonstrates the feasibility of detecting braking system performance on moving rail cars. This instrumentation is being converted into a Braking Inspection System (BIS) prototype. The objective of the BIS is

to provide a digitized output for braking alarm signals, wheel weight,

wheel temperature and braking reaction force.

Currently, "braking" or "no-braking" conditions can be detected and

presented in a digitized output. "Inadequate" or "excessive" braking is

dependent on the actual braking force magnitude. Thus, efforts are underway to discriminate the discrete braking force from a measurement

signal that includes a number of random forces not associated with the

braking action.

91

SESSION III

HIGHWAY AND ROAD BRIDGES

Chairman: J. Krugler, Federal

Administration

Highway

FAILURE ANALYSIS OF HIGHWAY BRIDGES

John W. Fisher

Fritz Engineering Laboratory

Lehigh University

Bethlehem, Pennsylvania 18015

Hans Hausammann

Fritz Engineering Laboratory

Lehigh University

Bethlehem, Pennsylvania 18015

Abstract: An evaluation of cracking that developed in two bridge structures is considered in the paper. Fatigue cracking was first observed

at the Yellow Mill Pond Bridge on the Connecticut Turnpike (1-95) in

1970. Fatigue crack growth resulted in complete fracture of a tension

flange in one of the girders. Smaller cracks were discovered in several

other beams. Fatigue cracking started at the end weld of cover plates

welded on the rolled section which forms the longitudinal girders of

the bridge structure.

The results from the bridge observations are compared with laboratory

fatigue test data on small size and full size cover-plated beams. The

laboratory tests and analytical predictions of crack growth agree well

with the observations made on the Yellow Mill Pond Bridge.

A large crack was discovered in November 1973 in a fascia girder of the

suspended span of the Quinnipiac River Bridge near New Haven, Connecticut.

Crack propagation was found to occur in different stages. A detailed

examination of the fracture is given in the paper. Fatigue cracks were

found to originate at lack of fusion areas in horizontal stiffener

splices. After the crack penetrated the web thickness it resulted in

brittle fracture of the web. Crack instability developed when the

stress intensity at the crack tip reached the material fracture toughness. The fracture toughness was estimated from J-integral measurements and from Charpy V-Notch test data.

Key words: Analysis; bridges, crack propagation; failure; fatigue;

fracture; fracture surface; fracture toughness.

1. Introduction : The web of the main girder of the Quinnipiac River

Bridge near New Haven, Connecticut, cracked after nine years of service.

Brittle fracture was arrested when one crack tip was near the girder

neutral axis and the second crack tip in the bottom flange. The tension

flange was about 40% destroyed. The investigation showed that brittle

fracture initiated from a fatigue crack that had entered the girder web.

95
Fatigue crack growth was also observed in several cover-plated beams of

the Yellow Mill Pond Bridge in Bridgeport, Connecticut. The fatigue

cracks were discovered after the bridge was in service twelve years. A

fatigue crack had propagated through the 32 ram (1.26 in.) thick tension

flange. Crack instability did not develop until the crack had propaga-|;;ed

several inches up the web.

The crack behavior of both highway bridges is evaluated and analyzed in

this paper.

\ 2. Quinnipiac River Bridge : In October 1973 a large crack was discovered in a fascia girder of the suspended span of the Quinnipiac River

Bridge near New Haven, Connecticut [11] . Figure 1 is a photograph showing the bridge profile. The crack was discovered approximately 10.4 m

(34 ft.) from the left or west end of the suspended span. The suspended

span is 50.3 m (165 ft.) long. The structure is noncomposite and the

girders are 2.8 m (9.23 ft.) deep at the crack location.

The crack initiated in the girder web and propagated to the middepth of

the girder and as shown in Fig. 2 had penetrated the bottom flange sur- face when discovered. The bridge had experienced about nine years of

service at the time the crack was discovered.

Examination of the fracture surfaces indicated that the fracture had

initiated at the web stiffener intersection. The fracture surfaces indicated that a butt weld in the longitudinal stiffener had been made at

this location but had never been completely fused. Close examination

revealed that only a surface pass had been made and the reinforcement

removed by grinding. The surfaces of this portion of the fracture were

severely corroded from their exposure to the environment.

Replicas were made at the web-longitudinal stiffener intersection, so that the fracture surface could be examined by transmission electron

microscopy. Visual examination at low magnification indicated that

fatigue crack growth was very probable in the web. The fracture surface

in the web on each side of the longitudinal stiffener indicated that a cleavage or "brittle fracture" had occurred after the crack penetrated

the web thickness. Cleavage fracture extended down the web and penetrated about 25 mm (1 in.) into the flange before it had arrested.

Examination at the web-stiff ener intersection confirmed that fatigue

crack growth had occurred. Fatigue crack growth striations were observed

adjacent to the longitudinal web stiffener break. Figure 3 shows a photograph at high magnification (49125 X) or a replica of the fracture

surface. The fatigue crack striations are apparent. Estimates of the

rate of crack propagation were made on the basis of the striation spacing. These indicated that crack growth rates between 1.8 and

—5 — 7 —6

5.1 X 10 mm/cycle (7 x 10 and 2 x 10 in/cycle) were occurring in

the region examined.

96
2.1 Material Characterization ; At the critical location the web and

the flange are fabricated from A36 steel. Tensile coupons provided a yield strength of 254 MPa (36.8 ksl) , an ultimate strength of 420 MPa

(60.9 ksl) with a 43% reduction in area and 32% elongation in a 203 mm

(8 in.) gage length. Standard ASTM Type A Charpy V-Notch (CVN) [6]

specimens were fabricated from both the web and flange material, and

compact tension specimens were made from the web.

The Charpy V-Notch tests showed that both flange and V7eb satisfied the

toughness requirements for Group 2 of the 1974 interim AASHTO Specification [2]. The average CVN Impact value for the web was 27 J (20 ft-lb.)

at 4° C (40° F) . The flange provided an average value of 47 J

(35 ft-lb.) at 4° C (40° F)

Several compact tension specimens were fabricated from material from the

web. Due to the limited material removed from the web, only eight com- pact tension specimens 51 mm x 64 mm x 11 mm (2 in. x 2.5 in. x 0.41 in.)

were made. The crack was oriented in the same direction as in the bridge

structure.

The specimens were tested at five different temperatures, one specimen

at -40° C (-40° F), two at -29° C (-20° F) , two at -23° C (-10° F) , two

at -12° C (+10° F) and one at -0° C (32° F) . The loading time for the

fracture test was about 0.1 second.

The compact tension specimen tests could not be evaluated using the ASTM

E-399 specification [8], because even at -40° C (-40° F) the available

material was too thin relative to the fracture toughness. For convenience the fracture toughness was therefore estimated using the J-integral

procedure proposed by Rice [9].

The empirical relationship suggested by Rolfe and Barsom [2] was used to

estimate the dynamic fracture toughness K-j-^ and the toughness for one second loading time were estimated by considering a temperature shift

[2]. The estimated K-values are shown in Fig. 4 and compared with the

0.1 second compact tension tests.

2.2 Stresses at the Crack Location : An equivalent stress range from

the live load can be approximated from Miner's Rule using the gross

vehicle weight distribution given in Ref. 4 assuming a value of a = 0.7

for the stress range reduction factor. This results in S^j^-j^^er ~ "^"^

(1.95 ksl) in the flange as compared to the design stress

range S ^ , =30.0 MPa (4.35 ksl). The stress at the longitudinal rDesign

stiffener is about 60% of the maximum flange stress. Hence at the longitudinal stiffener S „. - 8.1 MPa (1.17 ksl). The maximum live load rMmer

stress was estimated to be about 13.8 MPa (2 ksl). The average daily

truck traffic crossing the bridge was estimated to result in about

1,600,000 random stress cycles per year, corresponding to the

97
S - 8.1 MPa (1.17 ksi) . This corresponds to an average daily rMmer o ^ truck traffic (ADTT) of about 4300.

The dead load stress, the stress due to the weight of the structure

was 34.1 MPa (4.95 ksi) at the level of the longitudinal stiff ener on

the web

.

The fatigue crack started in the horizontal stiffener and grew into the

web. Since the two continuous fillet welds connected the longitudinal

stiffener to the web, the crack tip entered the web in a zone of high

residual tensile stress.

Residual stresses are present in the bottom flange due to the welding of

the web to the flange. It was assumed that the tensile residual stresses

are distributed over a semicircular area at the connection between the

web and the flange with a radius of half the thickness of the web, plus

the weld leg size. Over this area the stresses were also assumed semicircular and their magnitude equal to half the yield point. The dead

load stress in the bottom flange is 56.8 MPa (8.25 ksi) and the assumed

maximum live load stress 23.0 MPa (3.3 ksi).

2.3 Analysis of Crack Growth : From the examination of the fracture

surface it was apparent that crack growth had occurred in the

Quinnipiac River Bridge in a number of stages and modes. These stages

are illustrated schematically in Fig. 5.

Under normal traffic, and if about 6 mm (0.25 in.) of the 9 . 5 mm

(0.38 in.) thick longitudinal stiffener were unfused, fatigue cracking

would require between 2 and 20 million cycles of random traffic, depending on the proximity to a free surface, to propagate through the longitudinal stiffener thickness. If the crack had only been fused about

3.8 mm (0.15 in.) on one plate surface so that an edge crack resulted,

only about one million cycles of random traffic would be needed to

crack the longitudinal stiffener. Stage II of fatigue crack growth

would primarily develop after the stiffener was cracked in two. Electron

microscope studies of the fracture surface confirmed the presence of

fatigue crack growth striations during state II.

After the weld which connected the two longitudinal stiffeners together

was completely broken, the crack front moved into the web. The crack

propagated under normal random loading during this stage through the web.

An estimate of the time required to propagate through the web thickness

was made assuming the crack penetrated the web as a flat circular crack

with center of radius at the tip of the longitudinal stiffener. This

model was developed as a result of laboratory fatigue tests which demonstrated a comparative mechanism of fatigue crack growth. The stress

intensity factor K for a flat-circular crack with the crack front

approaching a free surface is

98
Where o = applied stress

a = crack length

b = plate width (width of stiffener plus web thickness)

For an initial crack size of the width of the stiffener 3.5 x 10 cycles

are needed to penetrate 95% of the web thickness. The striation spacing

measurements are in good agreement with these calculations.

Stage III was the brittle fracture of the web during a time of low tiemperature. A comparison of fracture surfaces from the compact tension

tests and the web crack indicated that the web had fractured when the

temperature was between -12° C (10° F) and -23° C (-10° F) . The estimated maximum K-value during Stage III is indicated in Fig. 4 and com- pared with the material fracture resistance. Brittle fracture initiated

in a zone of high residual tensile stresses. Once the crack became

unstable, it propagated through the zone of lower stresses in the web

and was eventually arrested in the flange.

Further fatigue crack growth (Stage IV) developed thereafter and con- tinued until the crack was discovered and repaired.

3. Yellow Mill Pond Bridge : Fatigue cracking was first observed at the

Yellow Mill Pond Bridge on the Connecticut Turnpike (Interstate Rt. 95)

in 1970 [1]. Fatigue crack growth had resulted in complete fracture of

a tension flange. Inspection of two beams adjacent to the fractured

beam indicated fatigue cracks had propagated halfway through their tension flange. Also, small cracks were visually observed at several other

beams.

The Yellow Mill Pond Bridge is located on the Connecticut Turnpike

(Interstate Rt. 95) in the City of Bridgeport, Connecticut. Construction

started in 1956 and was completed in 1957. It was opened to traffic in

January 1958. This bridge complex consists of 28 simple-span cover- plated steel beam bridges crossing the Yellow Mill Pond channel (14 in

each direction of traffic). Each bridge carries three lanes of traffic.

The position of the beams and diaphragms in span 10 are shown in Fig. 6.

The external facia beam (M88) of the eastbound bridge is skewed as four

lanes of through traffic are being reduced to three lanes.

The beams in span 10 are W36X230, W36X280, or W36X300 sections and were

rolled from A242 steel. All beams, except the interior facia beam of the

eastbound roadway (M86) , are fitted with two cover plates (primary and

secondary) on the tension flange and a single cover plate on the com- pression flange.

In October-November, 19 70, during cleaning and repainting of the Yellow

Mill Pond Bridge, one of the cover-plated steel beams on the eastbound

99
bridge on span 11 was found to have a large crack [A]. The crack had

developed at the west end of the primary cover plate on Beam 4 and Is

shown In Fig. 7. It had grown from the toe of the cover plate transverse fillet weld Into the tension flange and up 400 mm (16 In.) Into

the web . - ,. A visual Inspection showed that Beams 3 and 5 In span 11 of the eastbound

roadway which were adjacent to the casualty girder had cracks along the cover plate ends. These cracks were subsequently verified by ultrasonic

testing and a depth of penetration equal to 16 mm (0.625 In.) was measured. An indication of possible fatigue cracking was also observed at

other details on span 10 and on span 11.

Subsequently, other inspections were made at the Yellow Mill Pond Bridge

and fatigue cracks were found after removing the paint, dirt and oxides

which had accumulated at the weld toe. Inspections were made in December

1970, November 1973, June 1976, November 1976, September 1977, and November 1979 [1]. The results of these inspections are summarized in Fig. 6.

3.1 Material Characterization : The girders of the Yellow Mill Pond

Bridge were made from ASTM A242 steel with a yield strength of 395 MPa

(57 ksl) . Charpy V-Notch tests [6] were conducted on the web and flange

material. The material removed from the web shows a transition temperature of 20 J (15 ft-lbs.) at -23° C (9° F) . The flange provided 20 J

(15 ft-lbs.) at 15° C (59° F) . The difference in the transition temperature is due to the different thickness of the web and flange. There was more manganese in the flange material, which may have increased the

transition temperature. The CVN test results were used to estimate the

dynamic fracture toughness. The fracture toughness at an intermediate

loading rate and the static loading rate were estimated from the temperature shift. The estimated fracture toughness Is shown in Fig. 8 for the

web and flange material.

Three point bend specimen [8] were fabricated from material removed from

the web. Only four specimens could be made due to the limited material

available. The specimens were tested at temperatures of -40° C (-40° F)

,

-29° C (-20° F), and at -18° C (0° F) . They were analyzed using the

J-lntegral method proposed by Rice [9]. The specimens were tested at a loading time of 1.5 seconds which is comparable with the loading time

for the bridge structures. The results of the three point bend tests are compared with the fracture resistance estimated from the Charpy

V-Notch tests shown in Fig. 8.

3.2 Stresses at the Crack Location : The stresses in the bottom flange

near the cover plate and weld are caused by the live load, dead load and

residual stresses from rolling the beam and welding the cover plate on

the beam. The residual stress distribution for rolled W36X230 profiles

with welded cover plates was investigated in Ref . 10. The Investigation

was made using an analytical model and measurements. Stresses as high

as a /2 are present after welding the cover plate on the beam. The dead

100
load stress in the bottom flange at the crack location is 27.6 MPa

(4 ksi) . Several investigations were made to determine the stress distribution due to the traffic load.

The empirical relationship suggested by Rolfe and Barsom [2] was used to

estimate the dynamic fracture toughness K-j. and the toughness for one second loading time were estimated by considering a temperature

shift [2]. The estimated K-values are shown in Fig. 8 and compared with

the 1.5 second three point bend tests [8].

Stress History Studies : The eastbound and westbound bridges of span 10

in the Yellow Mill Pond viaduct were selected for two major stress history studies. Both studies were conducted by the State of Connecticut

and the Federal Highway Administration. The first study was conducted

in July 1971 [4] and the second from April 1973 to April 1974 [5].

It was found that slightly higher stress range events occurred over the

extended time period in 1973, then measured during the 1971 study.

A limited strain history record was obtained for the eastbound bridge of

span 10 in June 1976.

A large stress event was recorded during the stress history sample of the

west end of the eastbound bridge. Inspection of the oscillograph trace

indicated that this stress event was caused by a multiple truck presence

on the bridge. The stress range recorded at the end of the cover plate was 72.4 MPa (10.5 ksi) for Beam 7.

The histogram for the live load stress range is shown in Fig. 9. Miner

and root-mean-square (RMS) stress values, based on the frequency and

magnitude of strain events recorded in the July 1971 study and the 1976

sample were calculated. The following equations were used to calculate

these values: ^ ,„ ,

where S . is the i stress range in the spectrum and a, is the fraction

ri ° ^ 1

of stress ranges of that magnitude.

Based on average daily truck traffic (ADTT) measurements the total^number of trucks crossing the bridge was estimated to be about 21 x 10

between the opening of the bridge and the detection of the first crack

^ in 1970. The^critical detail was estimated to experience between 21 x 10

and 37.8 x 10 stress cycles.

The live load stress histograms obtained at Yellow Mill Pond are similar

to the stress histograms that have been obtained at other cover-plated

beam bridges in Michigan, Virginia, Maryland, Tennessee, and Pennsylvania.

The highest measured stress range varied from one histogram to another.

rRms rMiner

th

101
but the vast majority of these studies indicate an effective (Miner or

Rms) stress range between 6.9 MPa (1.0 ksi) and 13.8 MPa (2.0 ksi)

.

3.3 Fracture Beahavior : The fracture surface was heavily corroded and

partially covered with paint when discovered. Because of the corrosion,

no detailed f ractographic investigation could be made.

The crack started at the transverse fillet weld connecting the cover- plate to the tension flange and grew completely through the tension

flange. This fatigue behavior could be expected based on the results

from the laboratory studies [3,10]. The observed fatigue behavior is

compared with the predicted behavior in Fig. 10.

The stress intensity factor for a large semielliptical crack was determined as a function of the location on the crack front. For a crack

size of 25.4 mm (1 in.) the estimated maximum stress intensity factor

is 123 MPa /m . A comparison of the maximum stress intensity factor with

the material fracture toughness given in Fig. 11 suggests that the

static fracture resistance is applicable as no crack instability developed in the flange.

The fatigue crack completely penetrated the bottom flange and the crack

front moved up the web. At this stage the fatigue life of the cover- plated girder was completely exhausted. The load carried by the casualty

girder was redistributed to adjacent girders.

Numerous other details have experienced fatigue crack growth with most

cracks varied between 3 mm (0.125 in.) and half the flange thickness.

A. Summary and Conclusion : The cracking that developed in two bridge

structures was evaluated using fracture mechanics models for the crack

development. Fatigue crack growth was the primary cause of crack development in both bridge structures. The Quinnipiac River Bridge was found

to develop a fatigue crack from a butt weld of poor quality in a longitudinal stiffener. Fracture was predicted to occur when the fatigue

crack had nearly penetrated the girder web.

The fatigue cracks that have developed at the end of partial length cover plates on the Yellow Mill Pond girders resulted from large numbers

of stress cycles (~ 40 million) daily the past twenty years. The

fatigue resistance of this detail was eventually found to be much less

than anticipated at the time of the design. Later experimental and

mathematical models of the crack growth demonstrated that the fatigue

resistance was compatible with the field observations. No rapid fractures were found to occur as the maximum stress intensity at the crack

tip was always less than the fracture resistance of the rolled section.

102
References : 1. Fisher, J. W., Slockbower, R. E., Hausammann, H., and Pense, A.W.,

LONG TIME OBSERVATION OF FATIGUE DAMAGED BRIDGE, Paper presented

at the ASCE Annual Meeting In Portland, Oregon, April 15-18, 1980.

2. Rolfe, S. T. and Barsom, J. M., FRACTURE AND FATIGUE CONTROL IN

STEEL STRUCTURES, 1977, Prentice-Hall, New York.

3. Fisher, J. W., Hausammann, H., Sullivan, M. D. and Pense, A. W.,

DETECTION AND REPAIR OF FATIGUE DAMAGE IN WELDED HIGHWAY BRIDGES,

NCHRP Report 206, 1979.

4. Bowers, D. G., LOADING HISTORY SPAN NO. 10, YELLOW MILL POND

BRIDGE 195 BRIDGEPORT, CONNECTICUT, Research Project HPR 175-332,

State of Connecticut, Department of Transportation, 1972.

5. Dickey, R. L. and Severga, T. P., MECHANICAL STRAIN RECORDER ON A

CONNECTICUT BRIDGE, Report No. FCP 45G1-222, 1974.

6. NOTCHED BAR IMPACT TESTING OF METALLIC MATERIALS, ASTM E23-66,

Annual Book of ASTM Standard, Part 31, pp. 285-301, 1969.

7. Fisher, J. W. , Frank, K. H., Hirt, M. A. and McNamee, B. M.

,

EFFECT OF WELDMENTS ON THE FATIGUE STRENGTH OF STEEL BEAMS,

NCHRP Report No. 102, 1970.

8. PLANE-STRAIN FRACTURE TOUGHNESS OF METALLIC MATERIALS, ASTM

E.399-70T, Annual Book of ASTM Standards, Part 31, pp. 911-923,

1971.

9. Rice, J. R., A PATH INDEPENDENT INTEGRAL AND THE APPROXIMATE

ANALYSIS OF STRAIN CONCENTRATION BY NOTCHES AND CRACKS, Journal

of Applied Mechanics, Transactions, ASME, Vol. 35, 1968.

10. Roberts, R., Fisher, J. W., et al., DETERMINATION OF TOLERABLE

FLAW SIZES IN FULL SIZE WELDED BRIDGE DETAILS, Report FHWA-RD77-710, Federal Highway Administration, Office of Research and

Development, Washington, D.C.

11. Fisher, J. W., Pense, A. W., Hausammann, H. and Irwin, G. R.,

QUINNIPIAC RIVER BRIDGE CRACKING, Journal of the Structural

Division, ASCE, Vol. 106, No. ST4, Proceedings Paper 15343,

pp. 773-789, 1980.

T03
Fig. 1 Quinnipiac River Bridge

_6\</2" Stiffener

16- Inside Face (203mm x 13mm)

/4i/2"x3/e" \

Long. Stiff. ^ (114mm xlOmm)

(limm)

Left St i f f ener fPiece

U1/2 "x 3/8"

_

. UII4mmxl0mm)i]

^Web

Piece H

9'-2V

(2813mm)

Crock- Web

"Piece I

Flange Piece

T

21 6" (549mm)

I Vx22"

Flonge

(559mmx 35mm)

Fig. 2 Schematic of the Girder at the Crack Location

104
Fig. 3 Fracture Surface in Web Adjacent

to Longitudinal Stiffener Showing

Striation Markings (49125X)

-100 0 100

TEMPERATURE ,

Fig. 4 Fracture Toughness and Stress

Intensity Factor at Fracture

105
Stoge I

Initlol Crock (Unfused Weld) Fatigue Crock Growth Through Fused Port

Stoge n

Fatigue Crock Growth Through Web

Stoge HI

Brittle Fracture in Web

Arrested in Flonge

Stoge EZ

Fatigue Crock Growth

in Flonge

Fig. 5 Schematic of Crack Growth Stages

Not ImpacMd + No Crock o Crack Not

Verlli«<) By U T

« Crock v^rlfM ^ Small Crock Not VtrifM by UT

ft Root Crock

Cracks

t Brg. Pier

Fig. 6 Plan Showing Inspected Details in Span 10

106

70

60

50

40

30

20

10-

Crack Verified by U.T. Crock Verified by U.T. Crack Verified by U.T. Visual Indication of Cracking No Indication of Cracking

0.6

(4.11

1.05

(7.2)

1.50

(lO.S) 1.95

13.4)

-n-° I 2.40

16.5 2.85(19.7) 3.30(22.8) 3.75

(25.9) 4.20(29.0) Stress Range Of Primary Cover Plate , ksi (MPo)

Fig. 9 Stress Range Histogram; West Ends

of Eastbound Span 10, 1971

-95% Confidence Limits For Small Scale Beams

o Single Event

Per Truck

• Multiple Events

Per Truck

Estimated Lower Confidence

Limit For Cover Ploted Beoms

with tf > 0.8" (20mm) Cat. E'

I 10' 10'

40

10 (O

CYCLES

Fig. 10 Comparison of S ^, , with Calculated Cycles rMmer

and Fatigue Resistance of Beam Tests

108
Fig. n Stress Intensity Factor

for a = 25 mm, c = 139 mm

10^
FAILURE ANALYSIS OF DAN RYAN RAPID TRANSIT STRUCTURE

John M. Hanson

Wiss, Janney, Elstner and Associates, Inc.

330 Pfingsten Road

Northbrook, Illinois 60062

Abstract: On January 4, 1978, major cracks were discovered in three

adjacent rigid steel frames supporting an elevated curved section of the

Dan Ryan Rapid Transit. ^ Visual inspection of the cracks and subsequent microscopic examination

of the fracture surfaces established that the fractures started at the

welded junction of the bottom flange of girders that pierced the side

plates of the frame. Fatigue crack growth was found in the junctions.

Quality of the welding was poor, partly as a result of the geometry in

the joint. However, the notch from the rough flame-cut slot was determined to be sufficient to initiate the fatigue cracking.

Physical tests were made to determine the appropriate stress intensity

factor for the steel. This testing verified that the critical combination of a severe defect, high stress concentration, and cold temperature was sufficient to cause the fractures.

Key words: Failure; brittle fracture; fatigue; rapid transit; steel

frames; welding.

INTRODUCTION

On January 4, 1978, a passenger in a train passing underneath the Dan

Ryan Rapid Transit near 18th and Clark Streets in Chicago, Illinois

observed a crack in one of the rigid steel frames supporting the elevated structure. These frames are commonly referred to as bents.

Immediate action by Chicago Transit Authority (CTA) inspectors verified

the existence of the crack in Bent No. 24, and also revealed cracking in

two adjacent bents, Nos. 25 and 26. Photographs of the crack in Bent

No. 24 are shown in Fig. 1.

Construction of shoring was started in the evening of January 4, 1978.

The ground, which was solidly frozen at the time, was leveled with a layer of crushed stone. Mudsills, consisting of heavy timbers, were

placed on the stone base on both sides of each cracked bent. Falsework

towers were erected on the mudsills. As a precautionary measure, adjacent bents (Nos. 21, 22, and 23) of similar construction, were also

shored.

110 i
a) Side of bent

Fig. 1 - Views of crack in Bent No. 24

111
b) Other side, showing

origin of fracture

Fig. 1 - (Continued)

112
/

On January 14, 1978, a series of test trains with progressively increasing loads and speeds were moved over the shored structure. The

results of the test run were found to be satisfactory, and the system

was returned to service on January 15, 1978.

A Technical Committee was formed to investigate the fractures and to

develop recommendations for repair and retrofitting of the structure.

The Technical Committee, under the direction of Commissioner Marshall

Sulloway, included representatives of the Chicago Department of Public

Works, Chicago Transit Authority, Regional Transit Authority, and

DeLeuw, Cather and Company. Representatives on the Committee were

assisted by several consultants, including Dr. John W. Fisher at Lehigh

University. A report was published in January, 1979.^"'"^

BACKGROUND

The Lake-Dan Ryan Rapid Transit Line provides direct service between the

center of Chicago and the south and west sectors of the city. On the

Dan Ryan route, between 95th Street and 24th Street, the tracks are located in the median of the Dan Ryan Expressway and the future Franklin

Street Connector. At 24th Street, the tracks ascend from the median

onto a viaduct which extends to a junction with the old north-south

elevated structure at 17th and State Streets. Bent Nos. 24, 25 and 26,

where the failures occurred, are located west of Clark Street, directly

south of 18th Street. At this location, the alignment of the tracks and

structure is on a 400-ft radius.

The superstructure consists of four continuous and suspended plate

girders with a cast-in-place concrete ballasted trough. Bents supporting the superstructure are constructed with two column legs and a horizontal member spanning between them. The bents have a box-shaped

cross section. Stubs of the plate girders project 3 to 4 ft from the

sides of the boxes and are connected with the intervening girders by a bolted or link connection. The top flanges of the stub girders pass

over the top flange of the boxes, while the bottom flanges are welded to

and pierce the boxes through slots cut into the lower part of the side

plates.

Plans and specifications were prepared in accordance with criteria

established by the American Railway Engineering Association "Specifications for Steel Railway Bridges", the American Welding Society

"Specifications for Welded Highway and Railway Bridges", and the CTA

"Design Policy for Elevated Railways". The live load used for design of

the structure was the CTA Series 4000 Car for stress analysis and the

Series 2000 Car for evaluation of allowable deflections. These cars have four axles, spaced 6 ft, 27 ft, and 6 ft apart, with 8 ft between

the axles of successive cars. The axle loads are 24.5 kips and 17 kips

for the Series 4000 and 2000 cars, respectively. However, trains com- posed of Series 2000 cars had been used exclusively on the Dan Ryan

113
Line since its opening. The design also considered the following additional loadings:

Dead Load: Weight of track and ties (400 Ibs/lin ft of

track) plus weight of steel, concrete deck and

track.

Impact

As specified for electric locomotives.

Long. Force: 15% of live load without impact.

Wind and lateral

forces450 lbs /ft minimum.

Wind on train: 300 lbs/ft on one track applied 8 ft above top

of rail.

Nosing: 7000 lbs at top of rail in either lateral

direction at any point of span.

Structural steel conforming to the "Tentative Specification for Structural Steel for Welding, ASTM Designation: A36" was used in the structure, with an allowable tensile unit stress of 20,000 psi.

Construction began in April, 1968 and was finished in June, 1969. The

Dan Ryan Line was opened to traffic in September, 1969. On an average

weekday, 467 trains pass over this viaduct. The trains vary in length

from eight cars during the rush hours to as few as two cars at night.

The initial field examination of the fractures indicated that they

originated at the welded junction of the girder flange tips to the box

side plates. Chevron markings were observed on the fracture surfaces

pointing towards the welded junctions. The fracture surfaces on Bent

Nos. 25 and 26 appeared to be lightly rusted, while the fracture surface

on Bent No. 24 was more heavily rusted.

During the course of the field examination, the top pin and one of two

link plates connecting a girder to the stub extending from the west side

of Bent No. 24 were found to be missing. Despite a search, these pieces

were not found.

Measurements were made of the location and width of the cracks in each

of the bents. The cracks penetrated the bottom flange plate and both

side plates of all three bents, arresting at the edge of, or with only

very little penetration into, the top flange plate.

EXAMINATION OF FRACTURES AND TESTING OF STEEL

114
In Bent No. 24, the crack extended through and also into the web of the

girder. The trajectory of the crack across the bottom plate was inclined approximately 20 degrees to a normal to the longitudinal axis of

the bent, and the crack developed a short spur on the west side just

above the bottom plate, suggesting the presence of a torsional force

component in the bent when the fracture propagated.

A test program was carried out on pieces of 3/A-in. thick steel plate

taken from the sides of Bent Nos..24, 25 and 26, and a 1 1/2-in. thick

plate taken from the bottom of Bent No. 25. The pieces taken from the

sides of the bents included the origin of the fracture. These pieces

are shown in Fig. 2.

Samples were cut from the pieces and used for the following tests:

1. Visual and microscopic examination of the origins of the

fractures

2. Tensile tests

3. Chemical analysis

4. Charpy V-Notch tests

5. Compact tension tests

6. Metallographic examination

Examination

A sketch depicting the major features of the origin of fracture in Bent

No. 24 is presented in Fig. 3. The area which was not welded is shown

as the flame-cut surface. The vertical height of the flame-cut surface was about 3 1/4 in. A heavy multiple-pass weld was used to join the

outside surface of the box to the girder flange. On the inside of the

boxes, the weld was minimal. Paint was noted in gas holes and other

unfused portions of the fracture surface.

A smooth fine-grained surface was observed along the outside of the weld

on the outside of the box. Subsequent microscopic examination verified

that this surface was the result of fatigue crack propagation through

the weld. The major portion of the fracture surface, primarily above

and below the area of the welded junction, is coarse grained and contains chevron markings, typical of brittle fracture.

A detailed microscopic examination was carried out on the fracture surfaces of Bent No. 26, and a limited examination on Bent Nos. 24 and 25,

in order to establish the mechanism of crack growth and fracture. The

piece removed from Bent No. 26 was more extensively examined because it

115
Fig. 2 - View of pieces cut from side plate

of Bent Nos. 24, 25 and 26

116
Outside of box

Gas hole

Point

Study region

Flame cut

surface at

approx. 45

to surface

of side plate

Gas hole

Brittle fracture

Fatigue crack

from outside

surface

Slag

Unfused weld

Abraded

surface

Brittle fracture

Fig. 3 - Schematic of origin of fracture in Bent No. 24

117

III
Fig. 4 - Fatigue crack growth striations

of Study Region No. 2, Bent No. 24

(By John Fisher, Lehigh University, X13,000)

Fig. 5 - Cleavage fracture pattern at top

of Study Region No. 2, Bent No. 24

(By John Fisher, Lehigh University, X4000)

118
was not as abraded and dirty as the pieces from Bent Nos. 24 and 25.

The examination was made by both scanning electron microscopy and transmission electron microscopy, by personnel at Lehigh University.

Replicas of the surface were made at selected Study Regions, shown for

Bent No. 24 in Fig. 3. These replicas were examined with the transmission electron microscope. The examination of Study Regions 1 and 2 on Bent No. 24 indicated that fatigue crack growth initiated near the

inside of the flame-cut surface and extended toward the outside. The

growth was observed to extend at least 0.1 in. from the inside to outside in Region 1, and about 0.25 in. from inside to outside in Region 2.

The crack growth rate was observed to be about 8 x 10"^ in. per cycle

near the outside surface and increased to about 20 x 10"^ in. per cycle

near the flame-cut inside region. Near the top of Region 2, evidence of

cleavage fracture was apparent. Photographs of the fatigue striations

and cleavage fracture are presented in Figs. 4 and 5.

Striations indicative of fatigue crack growth were observed in the Study

Regions on Bent Nos. 25 and 26. The examination of the piece removed

from Bent No. 25 also indicated that fatigue crack growth developed from

the flame-cut area. These striations were observed to extend about 0.25

in. from their origin. The striations indicated a crack growth rate of

less than 10 x 10"' in./ cycle. Thereafter, ductile dimples and cleavage

facets were observed.

On Bent No. 26, the striations were most closely spaced near the outside

surface. Toward the inside, the striations turned and tended to move

down the side plate.

The scanning electron microscopy served to confirm the observations made

by the transmission electron microscopy, but did not add substantial new

information.

In summary, the examination of each origin of fracture showed that

fatigue crack growth originated adjacent to the flame-cut areas and also

at the exterior weld surface. The striations were all very finely

spaced, indicating that very low rates of fatigue crack growth were

experienced under cyclic loading. After the fatigue cracks extended and

coalesced, a brittle (cleavage) type of crack extension developed.

Testing

Tensile strength, yield point and percent elongation in 2 in. were

measured using round and flat coupons cut from the pieces taken from the

bents

Machining and testing of these coupons were performed in accordance with

the methods prescribed by ASTM E8. The coupons were oriented so that

the direction of fracturing in the test would be the same as the

119
direction of fracturing observed in the plates. The steel was found to

conform to the minimum mechanical property requirements of ASTM A36.

A chemical analysis was also performed on the steel. This analysis was made by the wet chemical method and atomic absorption techniques in

accordance with ASTM E350 and other pertinent ASTM specifications. All

samples conformed to ASTM A36 steel. The low silicon and aluminum

contents indicated that the steel was not killed or deoxidized.

Charpy V-Notch specimens were machined from the test pieces. Three

specimens from each piece were tested at each of the following temperatures: -50, -20, 10, 40, 70, 100 and 150°F (only two specimens from the

piece from Bent No. 26 were tested at 150°F) . The machining and testing were performed in accordance with ASTM E23. The results of these tests are shown in Fig. 6. They show good toughness with impact energies

greater than the present American Association of State Highway and

Transportation Officials' (AASHTO) minimum requirement of 15 ft-lbs at

40°F for Zone 2. Minimum service temperatures in Zone 2 are greater

than -30°F but less than -1°F. Transition temperature behavior for this

steel increases rapidly in the range between +40 and +70°F.

Compact tension tests were performed to obtain a measure of the critical

stress intensity factor, K^,, which would produce a brittle fracture

under the conditions existing at the time the cracking occurred in the

bents. The specimens for these tests were cut from the piece taken from

Bent No. 25, and tested in general accord with ASTM E399.

Three specimens were tested at each of the following temperatures:

-30°F, -5°F, and +30°F. The method of loading produced a fracture in

the specimens tested at -30°F and -5°F in about 1 second. Because of

the increased toughness of the specimens tested at +30°F, a loading time

between 2 to 4 seconds was required to cause fracture.

It was anticipated that valid Kj^ values, the critical stress intensity

factor determined in accordance with ASTM E399, might not be obtained

with specimens taken from the 3/4- in. thick side plates. Therefore,

displacement was measured in two ways: at the mouth of the opening using

an oscillograph, and also in line with the loading pins using an x-y

recorder.

In all cases, cracking occurred at the maximum load. Of the nine

specimens tested, one set of data was lost because the displacement gage

ran out of travel before the specimen fractured. None of the load

versus mouth opening displacement curves met the ASTM E399 requirement

for linearity, and therefore, the measured displacements in line with

the loading pins were used to indirectly obtain Kj^, by the J integral

method.

120
sqi - U 'NAD

121
A plot of the Kxc values versus test temperature is presented in Fig. 7.

It is evident that there is a significant transition in the toughness of

the steel at about 0°F. For the purposes of this investigation, the

measured values of K-j-^ are considered to be a conservative but realistic

estimation of K^, the critical stress intensity factor associated with

the conditions causing the cracking in the piers.

A metallographic examination did not reveal any defects in the steel.

The appearance of the grain structure near the surface was considered to

confirm both the fatigue and fracture modes of crack propagation.

STRESS ANALYSIS

In the design of the Dan Ryan structure, a three-dimensional space frame

analysis was used in order to include the effect of the curvature of the

girders. The space frame model consisted of two bents, the four superstructure girders framing between bents, the stub girders cantilevering

from the bents to the hinge points and the diaphragms between girders.

The drop-in spans between hinges were analyzed separately as grids.

Reactions determined from the grid analysis were applied to the space

frame

.

Various loading configurations on both single and double tracks were

used to arrive at the maximum stress conditions. The magnitude of

tensile and compressive stresses in the bents and girders was on the

order of 20,000 psi. Magnitude of shear stress was 12,500 psi. These

levels of stresses were permitted by the specifications of the American

Railway Engineering Association in effect at that time.

Following the discovery of the cracks, the original designer undertook

an independent investigation of the stresses in the bents, using an advanced three-dimensional space frame analysis.

Stresses in all bents under the Series 2000 Car loading were found to be

less than the allowable stresses, with the exception of Bent No. 24

adjacent to the girder where the fracture occurred. At this location,

the reanalysis indicated the side plate where the fracture originated

was overstressed in shear under dead load, live load (two tracks)

,

impact and centrifugal force. Specifically, the allowable shear stress

was 12.5 ksi, and the calculated maximum shear stress was 14.9 ksi. The

tensile stress range at this location for live load on two tracks plus

impact and centrifugal force was approximately 3.2 ksi.

With the exception of the moderate overstress in shear of the web of

Bent No. 24, the results of the reanalysis confirmed the original

stress analysis.

In an effort to more accurately assess the stress conditions near the

origins of the fractures, a computer analysis was performed by the

122
CM

400 -

300 -

uH

200

I 00

-50 -30

Test Temperature °F

Fig. 7 - Kj.^ Test Results (By Materials Research

Laboratory, Inc., Glenwood, Illinois)

123
original designer using the NISA (Numerically Integrated Elements for

System Analysis) program developed by the Engineering Mechanics Research

Corporation.

Because of the large size of the model, an attempt to minimize the

strain energy was not made. Therefore, the numerical values of displacements and of stresses obtained were regarded as approximate. The

welded joints between the box side plates and the bottom flanges of the

four curved girders were assumed to be completely fused; i.e., the

thickness of all elements was taken as constant throughout the region of

each element. Similarly, constant elastic properties of the material

were specified for all elements.

The structural model for the initial set of computations consisted of

Bent Nos. 24 and 25, and the girders extending from Hinge 18 to

Hinge 19, as illustrated by Model A in Fig. 8a. The model contained the same structural elements and support conditions as that used in the

original analysis, to allow comparison of results. Model A consisted of

1583 four-node shell elements defined by 1590 corner nodes.

Uniformly distributed dead load was simulated by statically equivalent

loads applied at the nodes along the top rows of girder web elements.

Additional concentrated loads were applied at the end nodes of all four

girders to account for the weight carried by the hinges in Span Nos. 24

and 26. Live load node forces were equivalent to train wheel loads

(without impact) . Two trains were placed in a position that leads to

maximum tensile stress at the bottom flange of Bent No. 24 at the intersection with the exterior girder. The Series 4000 Car was used for live

load.

Following the analysis of Model A, a second model representing the area

of the intersection of the exterior girder with Bent No. 24 was prepared, as shown in Fig. 8b. Model B contained 165 four-node shell

elements and 180 corner nodes. The boundary conditions specified for

Model B were the displacements obtained from the analysis of Model A.

Analysis of the deformations in the area of the intersection between the

box side plate and the bottom flange of the girder indicated that

stress concentrations due to out-of-plane bending of the side plate are

present. In order to determine the order of magnitude of these stresses

a third model was prepared. Model C was programmed with a finer mesh

than Model B, as illustrated in Fig. 8c.

Model C is a detailed representation of the side plate of Bent No. 24 in

the vicinity of and including a segment of the bottom flange of the

girder. This model contained 212 elements defined by 571 nodes. The

elements used in this model are eight- and six-node isoparametric shell

elements. All nodes along the boundaries of Model C coincide with the

corresponding nodes of Model B. The displacements of these boundary

124
125
b) Model B

t Girder 6127

Girder GI27

c) Model C

Fig. 8 - (Continued)

126
nodes obtained from the analysis of Model B were used as the prescribed

boundary conditions for Model C.

Stresses acting at the middle surface, at the outside face, and at the

inside face of the web were obtained from the finite element analysis.

The membrane stresses acting at the middle surface of the side plate were low (less than 10 ksi) , even at the level of the girder flange.

However, theoretical peak stresses due to out-of-plane bending of the

side plate exceeded the nominal yield stress for A36 steel, both in

compression on the outside face, and in tension on the inside face. The

live load stresses were, in general, lower than the dead load stresses.

Still, peak stresses in excess of 20 ksi were indicated. The peak live

load stresses are of the same sign as the corresponding dead load

stresses so that stress reversal does not occur.

DISCUSSION AND FINDINGS

Following the occurrence of the fractures in Bent Nos. 24, 25 and 26, an intensive review of the design, a careful inspection of these and other

bents and piers, and a thorough testing program on representative

samples of the steel were performed. The results of the design review

and the testing program were summarized in the preceding chapters.

Substantial variation was found in the quality of welding of the junctions between the girder flanges and bent side plates in Bent Nos. 21

through 26, and the details of some joints did not conform to the

details shown on the shop drawings. Numerous welding deficiencies were

apparent. Cracks believed to be caused by fatigue were found in at

least one junction in Bent Nos. 21 and 22 and in junctions other than

where the fractures originated in Bent Nos. 24, 25 and 26. In Bent

No. 23, the girder flange bears on top of the bottom plate of the bent,

creating a different condition than in the other bents. No cracks due

to fatigue were found in Bent No. 23.

Brittle fracture of structural steel members usually occurs when a critical combination of defect, stress intensity, and cold temperature

exist simultaneously. .

In Bent Nos. 24, 25 and 26, the defect was present in the form of a condition at the junction of the girder flange and bent side plate which

was very much like a "built-in crack". It was a difficult area in which

to weld, particularly in Bent No. 24 where the girders are skewed at an angle of approximately 45 degrees with respect to the box.

The analysis verified that high stress levels would be present at these

locations, under both dead load and repeated applications of live load.

High residual stresses due to welding would also be present. It should

be noted that calculated and residual stresses which exceed the yield

point will be redistributed by yielding of the steel. Thus, the cyclic

127
train loading in a highly sensitive area caused fatigue cracking, which

over a period of time increased the size of the defect and created a severe stress concentration at the tip of the crack.

Prior to the discovery of the fractures, temperatures in the Chicago

area were extremely cold on several occasions. There were five occasions when the temperatures dropped below 0°F and the coldest temperature of the winter season to that time, as recorded at Midway Airport,

was -7°F on December 10, 1977.

It seems likely that the first fracture occurred in Bent No. 24. The

analysis shows that the level of shear stress was higher at the point of

origin in this bent than in Bent Nos. 25 and 26. A link was missing on

one of the girders located adjacent to Bent No. 24. The influence on

the fracture caused by the missing link is not known, but regardless of

the disengagement of the link, conditions that would cause fatigue

cracking were present. Visual observations after the fractures were

discovered also indicated that a heavier coating of rust was present on the fracture surface of Bent No. 24 than the other two bents.

The laboratory testing confirmed that the combination of stress con- centration at the point of origin, fatigue cracking, and cold temperatures was sufficient to cause the fractures. However, the steel was

found to meet current toughness requirements recommended by the American

Association of State Highway and Transportation Officials for steels

used in bridges.

It was concluded that the extremely severe stress concentration at the

junction of the girder flange and bent side plate was the primary cause

of the brittle fractures which occurred in the Dan Ryan Rapid Transit

structure.

REFERENCE

(1) Final Report on Causes of Fractures in Bent Nos. 24, 25 and 26,

Dan Ryan Rapid Transit, Department of Public Works, Chicago, 1979.

128
ACKNOWLEDGEMENTS

This paper is based on the report prepared by the Technical Committe

and published in January, 1979. The Technical Committee consisted o

Marshall Suloway

Commissioner of Public Works

Department of Public Works: Louis Koncza, Chief Engineer

Chicago Transit Authority:

Regional Transportation

Authority

DeLeuw, Gather & Company;

Thomas L. Wolgemuth, Mgr.,

Engineering

Dan Brescia, Mgr. of Engineering

Department

William G. Horn, Staff

Vice President

The Technical Committee was assisted in the conduct of the investiga

tion and preparation of their report by:

Department of Public Works:

George Eng

John M. Hanson

Deputy Commissioner

Consultant, Wiss, Janney, Elstner

& Associates, Inc.

Regional Transportation Authority:

B. J. Ford

Don Buck

John W. Fisher

Anton Tedesko

Hannskarl Bandel

Division Director, Transportation

Director of Safety

Consultant, Lehigh University

Consultant

Consultant

DeLeuw, Gather & Company;

G. M. Randich Senior Vice President

The examination of the fracture surfaces and material testing was performed by Wiss, Janney, Elstner and Associates, Inc., Lehigh University, Materials Research Laboratory, Inc., and Taussig Associates

Inc. The analysis was made by DeLeuw, Gather & Company.

129
BRIDGE WELDING AND FRACTURE CONTROL

Carl E. Hartbower

Federal Highway Administration

Washington, D.C. 20590

ABSTRACT: In the 1970 's, at least twenty bridge girders suffered

brittle fracture and over a hundred girders were found to be cracked;

fortianately (or providentially) there was no catastrophic collapse.

However, with an apparently increasing incidence of weld cracking

and fatigue cracking, the probability of a catastrophic collapse

occurring in the 1980 's is very real indeed.

The range of occurrence of weld cracking in steel bridges has reached

alarming dimensions. Why? Is there a problem with the design of

welded connections? Is there a shortage of qualified welding engineers? Is there a problem with quality control and quality assurance

in fabrication? Is there a people problem and/or deficiencies in the

flaw-detection techniques? Is there a need for nondestructive testing engineers? Is there a base-metal weldability problem? Is there

a problem with cosmetic corrections and weld repairs? The answer is an emphatic yes to each of these problems and more!

The situation is so serious that without prompt positive action on the part of all concerned, there will be more and more bridges made

of concrete even where steel makes better engineering sense.

The paper takes a second look at some of the above problems using

bridge failures from the 1970 's as examples and then suggests some changes that will have to be made in the 1980 's if the problems we now face are to be solved.

KEY WORDS: Fracture control; nondestructive inspection; quality

control; welded steel bridges.

STATE-OF-THE-ART TECHNOLOGY, BRIDGES

Cracked welded-plate-girder bridges. The incidence of brittle fractures and weld cracking in steel bridges has reached alarming

proportions in the 1970 's. Over twenty bridge girders have developed

brittle fractures, and in-field inspections of bridges imder traffic

have revealed hundreds of rejectable indications using ultrasonic

testing (UT) . Millions of dollars have been spent in bolt-splice

repair of defective butt welds in fracture-critical members. The

130
Silver Memorial Bridge at Point Pleasant, West Virginia, the

Paducah Bridge on 1-24 between Kentucky and Illinois, and the Fremont

Bridge in Portland, Oregon, are among the major bridges known to have

UT-rejectable indications at the time of this writing. Intensive

investigations are in progress to determine why and when the cracking

occurred.

In just one State, inspection of 44 two-girder bridges with intermediate floorbeams and stringers revealed 24 bridges with cracked

members. The cracks are occurring near the top flange at the floorbeam-to-girder connection plates; the general problem is believed to

be out-of-plane deformation in the girder webs caused by end moments

in the floorbeams. In the 24 cracked bridges, there were 290 cracked

locations in negative-moment areas.

In one bridge which in 1973-74 had been given "in-depth" inspection

including UT, in 1977 was found to contain 132 UT-rejectable indications of length %-in. or greater in 51 flange groove-weld butt

splices. Many of the indications have since been confirmed to be

weld cracking and the fact that many of the cracks were found in

compression members as well as tension members suggests that the

cracks were in the welds at the time of fabrication. Investigation

is continuing. Another bridge contains a great number of fatigue

cracks at the toe of fillet welds connecting cover plates to flanges.

Some of the cracks were as long as 2 in. and over ^^-in. deep as determined by UT. The recorded stresses were small (RMS stress

range only about 1.3 ksi); a tremendous traffic count is said to

explain the fatigue-crack growth. But there may be another problem

which has not been adequately addressed - premature initiation of

fatigue crack growth as a result of weld-toe cracking occurring at

the time of fabrication and undetected by quality control (QC) or quality assurance (QA). Microcracking at the weld toes intensifies

the stress-concentration factor inherent in the toe geometry and is

exceedingly difficult to detect in nondestructive inspection.

Why have there been over twenty brittle fractures in bridge girders in

the 1970' s and why has there been so much cracking in bridge girders

in the 1970' s? Why are so many cracks found after the bridges are erected and under traffic? These and related questions are examined

in the following paragraphs using as examples some actual bridge

problems from the 1970' s.

The base-metal weldability problem . Consider two examples. The

Bryte Bend Bridge, Sacramento, California, June 13, 1970, developed

brittle fracture of a main-load-carrying tension flange. The

flange that fractured came from a heat of steel (C4913) that occurred

in ten locations in the bridge. The steel was ASTM A517 grade H (modified), 2-% in. thick. Not only was the steel seriously lacking

131
in toughness but also it was highly susceptible to hot cracking.

It was a weld hot crack that initiated slow crack growth during shop

handling and erection of the girder. Subsequent inspection revealed

numerous cracks in the standing bridge where cross bracing was welded

to the A517-H flanges. There was also a serious cracking problem

when the fabricator attempted to qualify his welding procedures by

test; transweld cracking occurred in the submerged-arc-welded (SAW)

butt splices, cracking that sometimes ran several inches into the

base metal

A second bridge containing the same steel, the Silver Memorial Bridge

at Point Pleasant, West Virginia, replacing the bridge that collapsed

killing 46 people in December 1967, has experienced serious weld

cracking problems. This bridge is highly significant to the question of whether base-metal weldability is a problem. Two companies

were contracted to fabricate the truss members for this bridge; the

fabricator on the West Virginia side of the river used A514 grades

F and H from the same steel producer supplying steel for the Bryte

Bend Bridge, and the fabricator for the Ohio end of the bridge used

A514 grades F and H from a different steel producer. Ultrasonic

(UT) inspection of the groove-weld butt splices revealed numerous

UT-rejectable indications, some so large that over a million dollars

in bolted-spl ice-plate retrofitting was deemed necessary by the State

of West Virginia. The statistics of defect location and incidence

rate clearly showed that the A514 grade H chemistry as supplied by one of the steel producers was causing a significantly greater

amount of weld cracking than the same ASTM type and grade supplied

by the second producer. Clearly there was a weldability problem

with the A517 grade H Steel supplied by one of the producers. The

AWS Dl.l Structural Welding Code lists A514/517 as prequalified

steel irrespective of grade and, therefore, under the AASHTO bridge

specifications there is no requirement for weldability testing this

steel.

The design-of-welded-connections problem . Consider two examples.

When the tied arch at Paducah, Kentucky was completed, long behind

schedule, weld cracking in the form of lamellar tearing had been

so costly for the fabricator, he sued the State of Kentucky

claiming over a million dollars in extras. The court found the

design to be a major contributor to the problem and assessed

damages at 80:20 in favor of the fabricator. In other words, the

court found that 80 percent of fault was with the design and since

the designer was an agent to the State, the State had to assume

80 percent of the cost. The remaining 20 percent was found by the

court to be the result of bad welding. This problem was discovered

before the bridge was opened. Now, just a few years later, the bridge

is closed at the time of this writing, having been found to contain numerous UT-rejectable planar-type indications in the A514 tie-girder

groove-weld butt splices and corner joints. The bridge has been in

service less than 10 years.

132
When the Lafayette Street Bridge in St. Paul, Minnesota developed a brittle fracture, failiare analysis revealed a gusset-to-web-tostiffener connection which according to the contract plans was to be

welded from both sides and without coping, so that the three welds

were intersecting. With the gusset located only six inches above

the flange, the fabricator found the joint design impossible so he

used a backing bar, and then failed to get complete penetration.

The ensuing fatigue cracking in the gusset-to-stiffener connection

entered the girder web and triggered brittle fracture.

The "cosmetic-corrections" and weld-repair problem . Bridge after

bridge with weld cracking has been found to have weld repair asso- ciated with or directly causing the cracking. The first crack found

in the Silver Memorial Bridge was approximately 1 by 4 in. in a multiple weld repair. At least two repairs were made in the location

that cracked. The crack that initiated brittle fracture in the

Neville Island Bridge in Pittsburgh was in a massive repair of an electroslag weld. When weld repairs crack it should not be suprising considering the high restraint that is inherent in a local ex- cavation, and considering the poor welding techniques often used,

sometimes without knowledge of QC management.

"Cosmetic corrections" are particularly insidious because it is com- mon practice for the corrections to be made without preheat and without benefit of nondestructive inspection. Often the welder decides

what is to be done and how it is to be done. Consider, for example,

minor undercut sometimes even within the limits acceptable under the

Structural Welding Code (AWS Dl.l); if the welder chooses to make the

"cosmetic correction", a low-energy-input stringer bead is typically

used with perhaps 30,000 J/in. and no preheat. If the flange

happens to be 2 or 3 in. thick, that makes no difference to the

welder, but to the steel it can mean untempered martensite. Then

there are "cosmetic corrections" to oxygen-cut flange edges; if a gouge is produced in cutting, this is sometimes corrected by a short

(tack-like) stringer bead across the edge of the plate . Again, this

is typically done with low-energy input and without benefit of preheat. If sabotage were the intent, it could not be more effective!

Since it is not a "repair weld", it may be done without knowledge

of QC/QA. Such welding is likely to result in toe cracking and

where the weld is transverse to the principal stress , the toe

cracking is likely to initiate premature fatigue cracking and/or

brittle fracture. Unfortunately, there are advantages in the

concept of "cosmetic corrections," namely no paper work and no delays

with inspection, preheating or post heating, so it will not be easy

to eliminate the practice.

133
In the AASHTO Guide Specifications for Fracture Critical Nonredundent

Steel Bridge Members , September 1978, there are several provisions

specifically on repair welding. Repairs are placed in two categories:

noncritical and critical. When repair involves removal of weld or base metal, surveillance by QC/QA is essential. Engineering decisions

are involved as to the orientation of the excavation, the choice of

electrode, the joint configuration (including angle and root opening),

stringer-bead placement and preheat/interpass/postheat temperature.

All too often when little attention is given to these considerations,

weld repairs crack, necessitating multiple repairs in a given location. When hydrogen is the cause of cracking, if the repair itself

does not crack, it is not unusual to find additional cracks nearby in

weld locations previously accepted by QC/QA. Weld repair and cosmetic

corrections are a major problem, as proven by cracks found in standing

bridges.

The quality-control/quality-assurance problem . QC is virtually

nonexistant in some shops and QA is often a passive hands-in-thepockets operation, underpaid, understaffed, and underskilled. Standing bridges with cracking or brittle fracture always raise the

question: When did the cracking initiate , was it in the fabrication

shop or was it the result of fatigue? To answer the questions one should be able to refer to the fabrication QC/QA inspection records.

Case after case can be cited where inspection was by radiography

but the film was so poorly exposed or so poorly processed that it is

virtually useless. Where was the QA inspector who was responsible

for reviewing and approving the film? The advantage, we are told, in

radiographic testing (RT) is in having a permanent record, but what

good is the record if the techniques are so poor that the record

cannot be read? In the case of one bridge that suffered brittle

fracture, the RT files were lost. In another case, the file was found and rejectable indications were visible in the film; apparently

no one had read the radiographs.

Then there are numerous cases where AWS Dl.l has been violated and

apparently no one, neither QC nor QA, saw the violation. AWS Dl.l,

Section 9 specifically prohibits tack welds left in the finished

bridge. The tie-girders of a large tied-arch bridge have recently

been found to contain man holes cut in the webs of the tie girders

where in closing the manholes after erection, a backing bar was tack

welded on the inside of the box. The opening was closed in the final

stage of erection. AWS Dl.l, Section 9 requires that backing bars

be spliced and inspected so as to assure that there will be no cracked

groove-weld butt joints. The tie girders of a large award-winning

tied arch under traffic have been found to contain numerous cracked

butt splices in the steel backing corner welds. The defective welds were in violation of AWS Dl.l from the day they were welded.

Where was QC/QA in that bridge? The Fort Duquesne Bridge in

134
Pittsburgh has a problem with lamellar tearing, together with grooveweld toe cracking, sometimes across the entire 48- in. width of flange.

The design called for a cruciform joint configuration in 2-1/2 in.-

thick ASTM A517 steel. The lamellar tearing can be seen by amy pedestrian walking the bridge at the time of this writing. Why didn't

QC/QA see the cracking at the time of welding?

In the 1970 's, a footnote in the Structural Welding Code (footnote

27 page 73 of AWS Dl.1-75, or page 103 of AWS Dl.1-79) placed

responsibility for inspection on the owner. For some members of the

AWS Dl.l Committee, it was incredible that the fabricator should not

be held responsible for continuing evaluation of the quality of his

work, and yet under footnote 27 the "Inspector" was an agent of the owner. Moreover, in the 1970 's there was no distinction between QC

as a responsibility of the fabricator and QA as a prerogative of the owner. For reasons not altogether clear to the author, the prolonged

arguments against clarifying the QC/QA functions included adament

rejection of the terms QC and QA in spite of the fact that in other

industries the terms are commonplace and well defined. Anyone

interested in the compromise reached by the AWS Dl.l committee needs

only turn to AWS Dl.1-80 (1980) paragraph 6.1. A few State

Departments of Transportation (DOT) including New York and California

have long rejected the provisions of AWS Dl.l in this regard; in

November 1979, FHWA issued a Technical Advisory clearly defining the

roles of QC and QA, recommending that QC be specifically the

responsibility of the fabricator and QA the prerogative of the owner

(the State DOT) . Moreover, it was recommended that QC for fracturecritical members (FCM's) include both RT and UT as complementary

test methods, and that QA no longer be a passive watch-dog operation

but use hands-on auditing of QC by UT.

In the late 1970 's, intensive reinspection of several bridges xonder

traffic containing fracture-critical members (FCM's) revealed

instances of both cracking and rejectable workmanship defects (lack

of fusion, slag and porosity) at the ends of submerged-arc -welded

(SAW) groove-weld butt joints. Many of the bridge members had been

siibjected only to RT. In the 1970' s AWS Dl.l had no provision for

minimizing scattering off the cassette into the image area, which

tends to obscure defects at the plate edge; when the end of the butt

splice corresponds to the start of welding, there appears to be a greater probability of defects, perhaps associated with inadequate

or improper starting tabs or preheat. Linear-elastic fracture

mechanics shows that the stress intensity associated with an edge

crack can be dangerously high and, therefore, is likely to initiate

premature fatigue crack growth and/or brittle fracture.

Another type of cracking seriously neglected by QC/QA and yet

dangerous from the standpoint of promoting premature fatigue crack

) 135
growth is fillet-weld toe cracking. This type of cracking is insidious

because inspection for such cracking is usually only done by magnetic

particle testing (MT) which may produce a buildup of iron powder at

the geometric discontinuity inherent at the interface of the plate

surface and weld crown. Dye-penetrant and/or eddy-current inspection

would be far more effective for detecting weld toe cracking but are seldom used. MT with prods must be done before painting (bare metal

contact); both MT with prods and AC-yoke MT must be done in two

directions if both toe cracking and transverse weld cracks are to be

detected.

The critical-crack-size/nondestructive-testing problem . If linear

elastic fracture mechanics (LEFM) is to be utilized, nondestructive

testing must be capable of measuring both the length and depth of any

given indication. Among the weld defects that are of greatest concern are the planar defects such as lack of fusion, incomplete

penetration and cracking. The current AASHTO Guide Specifications for

Fracture Critical Nonredundant Steel Bridge Members (September 1978)

specifies Charpy V-notch impact minimum values for qualifying the

welding procedures used in all weld joints in "fracture critical"

members. The Charpy requirements for weld metal joining A36, A588,

and A572 steels is 25 ft-lb at minus 20°F and for weld metal joining

A514/517 steel is 35 ft-lb at minus 30°F.

Based on the Barsom/Rolfe relationship

Ki^2 = 5(CVN)E

where CVN is the weld-metal Charpy ft-lb value and E is Young's

modulus, one can estimate the critical stress intensity corresponding

to the specified Charpy values, and from the critical stress intensity

estimate the critical crack sizes. Table 1 gives the calculated flaw

dimensions for two flaw shapes when the planar indication is embedded

in the weld. Note that the critical depths in the weld-residualstress field are less than one inch. With the ultrasonic testing

techniques presently used by the bridge-building industry, there is

a serous problem in quantitatively sizing a flaw when the flaw depth

is less than the UT-beam width.

Without the ability to measure both the length and depth of a given

planar indication, it is not possible to utilize LEFM. Furthermore,

even with quantitative flaw sizing, without a realistic estimate of

the largest flaw that may escape detection in nondestructive testing,

it is not possible to specify the required toughness. Furthermore,

to allow for possible error in flaw measurement and to allow for some crack growth before instability, the nondestructive techniques

136
and personnel skills must assure that the largest flaw escaping detection is no more than say 50 percent of the critical crack size. To assure that defects will be found that are smaller than the critical

crack depths listed in Table 1, either greater toughness must be

specified or nondestructive testing techniques and personnel skills

will have to be improved.

It is the shockingly large cracks that have been found in standing

bridges that attest to a problem both with nondestructive testing

and welding as presently conducted.

10-YEAR FQRCAST

Can the incidence of weld cracking during and after fabrication be

reduced? To do so during fabrication will require the services of a welding engineer and careful consideration of sometimes neglected but

vital details such as (1) cracked tack welds (weld metal and/or

weld heat-affected zone), (2) consistent implementation of lowhydrogen practice, (3) control of "cosmetic corrections" and weld

repairs, (4) control of cooling rates to preclude both brittle weldtransformation products in the heat-affected zone and loss of toughness in the weld metal, (5) use of welding consumables that provide

tough, crack-resistant weld metal, (6) control of welding procedures

to minimize planar-type defects, (7) designer review of welded

connections at the shop-drawing stage, and (8) attention to the

inspectabi 1 i ty of welded connections.

And to markedly reduce the incidence of cracking after fabrication

it will be necessary to have the services of a nondestructive

testing engineer (at least an ASNT Level-III person) and qualified

ASNT Level-II inspectors capable of finding and sizing buried weld

defects.

To realize these requirements in the 1980' s, some changes will be

necessary. For the last two decades our colleges and universities

seem to have been oblivious to the need for solidly educated welding engineers and welding metallurgists.

Attempts to get educators' attention by published articles comparing

U.S. and USSR curricula and the numbers graduated in welding engineering has been of no avail. With an average of only about 20 BS

welding engineers in the USA per year for the last 20 years, as compared with 2000 each year in the USSR, and currently about 800

majoring in materials science and metallurgical engineering each

year in the USA and Canada as compared with 8000 each year in the

USSR, it is not surprising that welding engineers are seldom found

in plants fabricating bridge girders. The FHWA for the last 2 years

has advertized for a welding engineer, offering a qualified person

137
$30,000 per year plus excellent fringe benefits, and is still looking.

Only two schools in the USA offer a BS degree in Welding Engineering,

Ohio State and LeTourneau (Longview, Texas). The author knows of no college or university in the USA offering a BS in Nondestructive

Testing. If in the 1980' s we are to minimize the incidence of weld

cracking, and when occasionally it does occur, if we are to be

assured of finding and sizing the cracks, it will be necessary to have

an adequate supply of well-trained welding engineers and nondestructive testing engineers. This is true not only for more efficient,

better weld fabrication but also if we are to regain the lead in

research and development. Without the latter we will be coasting;

Omer Blodgett of Lincoln Electric tells his design classes that in

todays rapidly advancing technology, to coast is to go down hill.

Assuming that our college and university administrators see the

light and deliver well-trained welding engineers and nondestructive

testing engineers in sufficient quantity, let's look at some of the

changes needed to reduce the incidence of weld cracking as enumerated

above.

(1) Control of tack welds . AWS 01. 1 as presently written is ambiguous and imcomplete in this connection. In bridges, tack welds are

not allowed in the completed structure and, therefore, must be

removed or overlayed with a continuous weld. When tack welds are not

completely consumed by the final weld, cracks that are produced in

the tack weld or in the heat-affected base metal under the tack weld

are initiation sites for fatigue or brittle fracture in service.

Tack welding is inherently a low-energy-input procedure and, therefore,

is inherently prone to produce weld heat-affected-zone (HAZ) cracking.

Moreover, the high hardness that attends low-energy-input welding

procedures in hardenable steels, like A588 and A514, makes hydrogen

cracking more likely. If preheat is used in tacking, the chances

of both HAZ hardening and hydrogen cracking will be reduced. To

accomplish this AWS 01. 1 par 3.37 will have to be revised to read:

3.37.1 Tack welds should be subject to the same quality

requirements as the final welds except that (1) minor

(superficial) discontinuities such as undercut and unfilled

craters need not be removed before final welding.

3.3.7.2 Tack welds, whether incorporated into a final weld

or not, shall be subject to the same procedural and

consumables requirements as the final welds -

138
(1) when cracking and/or porosity are visible in the

surface of a tack weld, the tack weld shall be

removed before final welding;

(2) whether incorporated into a final weld or not,

tack welds shall be preheated to 150°F or in

accord with Pable 4.2, whichever is higher.

3.3.7.3 Tack welds which are not to be incorporated into

final welds shall be removed and the tack-welded surfaces

ground smooth and flush, and the surfaces inspected with ' MT and/or DT.

There are many changes to be made in the Structural Welding Code;

the above is cited in detail as an example of what will have to be

done in the 1980 's to tighten the present specification.

(2) Use of low-hydrogen practice . The use of low-hydrogen electrodes

is only the first step in avoiding delayed (hydrogen) cracking. Lowhydrogen practice involves control over moisture and contaminants

(such as hydrocarbons) at every step of fabrication. Weld plate

surfaces and welding wire must be free of rust and hydrocarbons.

Rust may contain chemically combined water which can not be driven

off by preheating. Grease and oil and drawing lubricants (on wire)

are hydrocarbons which break down in the arc to form hydrogen.

E6010 and E7010 electrodes should never be allowed where low-hydrogen

practice is specified. After consumables (electrodes and fluxes,

including FCAW wire) are exposed to the ambient shop environment for

times in excess of the code-stated limits, low-temperature (300 to

400° F) baking in a low- vacuum chamber will permit multiple rebaking

without damage to the consumables. Properly used holding ovens for

flux and electrodes positioned near the welding station are essential

to low-hydrogen practice. Preheat/interpass and postheat temperatures

should no longer be arbitarily specified but calculated based on the

critical cooling rate to avoid undesirable weld-transformation

products. Post-heating of welds and weld repairs will be done with

temperature and time calculated to remove all but very low levels of

residual hydrogen. Carbon equivalent calculations will be made for

each heat of steel to assure that an adequate preheat/interpass

temperature is used to avoid excessive HAZ hardness.

Obviously low-hydrogen practice will only be implemented in the 1980's

with a sufficient number of (1) certified and qualified welding

inspectors and (2) welding engineers.

(3) Weld repairs and cosmetic corrections . The latter constitute

a situation that is presently out of control. Shop personnel are making engineering judgements as to what corrections need to be made

139
and how the corrections are to be executed. If the incidence of

cracking is to be reduced in the 1980' s, cosmetic corrections as such

will be prohibited and treated as weld repairs. Weld repairs will be

given extraordinary attention by both welding and nondestructive

testing engineers, with mandatory weld-qualification-procedure testing for repairs made to fracture-critical members. All critical weld

repairs will be post heated with temperature and time calculated to remove all but a very low level of residual hydrogen.

(4) Control of weld-transformation products . With knowledge of the

minimum cooling rate for avoiding low-toughness, high-temperature

transformation products and the maximum cooling rate to avoid the

martensite transformation (hard, brittle HAZ), the welding engineer

will calculate the energy input and preheat temperature that will

result in cooling rates within an allowable range.

When steel heats are found to have high hardenability, with carbon

equivalent (CE) in excess of say 0.60 based on the IIW formula, weld

transformation to martensite will be avoided by interrupting the

cooling cycle just above the Ms temperature (start of the martensite

transformation) using the chemical composition for calculating the

Ms temperature. The duration of the postheating will be based on the

time to isothermal ly transform to lower bainite.

(5) Tough crack-resistant weld metal . As long as human beings are making welds, there is the chance of error and/or omission and

problems with workmanship. Thus, the weld deposit is the most likely

place to find defects, defects ranging from relatively innocuous

porosity to planar defects like lack of penetration and cracks. If

such defects are to remain dormant or at worst grow very slowly, the

weld metal must be tough so that the critical crack size is no smaller than the plate/weld thickness.

(6) Sizing of defects . Ultrasonic testing will have to be updated

both with regard to instrumentation and certification of personnel.

Inspectors will be certified for assignment to fracture-critical

members on the basis of a performance test administered by the State

Department of Transportation. Automated scanning devices will be

utilized together with computer analysis of the recorded data. Band C-scan testing will be done to determine (1) the cross-sectional

position of a reflector and (2) plan-view contour maps of echo

amplitude. Two-probe time-delay testing will be used for defect

sizing; a diffracted longitudinal -wave pulse of ultrasonic energy

from the defect extremity will be accurately timed enabling the

position of the extremity to be defined in the vertical plane.

140
(7) Crack-tough steel for bridges . Fracture-critical members (FCM's)

will be fabricated using steels purchased with guaranteed minimum

Charpy V-notch impact values at the lowest anticipated service

temperature (LAST) and a guaranteed ASTM E208 nil -ductility transition

(NDT) temperature 30°F or more below the LAST. Crack-arrest properties at the LAST will be specified. When the designer anticipates

problems with lamellar tearing, Z-di recti on properties will be

specified. Several new steels will be used (new only to bridge

construction) such as ASTM A633 grades A thru E, incl, A678 grades

A, B and C, and A710 grade A classes 1, 2, and 3. The latter, with

0.07 carbon maximum, will be specified where FCM's are to be welded

in the field; otherwise, no welding will be allowed during erection

of FCM's. The Charpy V-notch test will be used only in the steel

mill for checking plate compliance with the minimum specified properties. The precrack Charpy impact test will be used by QC/QA to

verify that the NDT temperature is at least 30°F below the LAST.

Crack toughness will be specified for both weld heat-affected zone

(HAZ) and fusion zone. The microstructure in the HAZ with lowest

crack initiation/propagation resistance will be determined using

composite precrack Charpy impact testing as a requirement of weldprocedure-qualification testing for FCM's.

Just about everyone will admit that it is better and less costly to

do a job right the first time. But short cuts (cutting corners)

can be very very tempting from the standpoint of cost and delivery

schedules. But when a fabricator runs into weld cracking problems,

it frequently means extraordinary cost escalation and loss of

production time. When cracks escape QC/QA and are found after a bridge is in service, the cost is usually passed to the tax payer.

Millions of dollars are lost each time this happens, in inspection,

retrofit and manhours lost in redirected traffic. And in the event

of a catastrophic collapse with possible loss of life, the cost is

incalculable.

141
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THEORY AND DESIGN OF INSTRUMENTATION

FOR BRIDGE INVESTIGATION

D. A. Bracher

ManTech International Corp.

Arlington, Virginia 22202

D . A. Garrett

National Bureau of Standards

Gaithersburg, Maryland 20234

C. O. Heller

CADCOM Division

ManTech International Corp.

Annapolis, Maryland 21401

The U.S. Department of Transportation is currently responsible for

approximately 500,000 bridges. Approximately 100,000 bridges per year

require maintenance.

According to Escalante and Itcr^^, the rate of deterioration of concrete

structures has increased considerably over the past 10-15 years: The

1969 annual total cost estimated for restoration and protection of

bridges on the interstate road system was $2.6 billion.

Commercial instrumentation is available for evaluating the deterioration

of small metal probes that can be imbedded in concrete in close proximity to the metal of interest. This may be adequate for future concrete

bridge footings, but does not address the problem of testing older

footings

.

An "Acoustic Crack Detector" mounted on a back pack is currently used

with marginal success. Radiation gaging has been tried and found to

give inaccurate diagnostic information.

The fabrication of a panacean detection instrumentation is highly un- likely for the diagnosis of bridge footingsy However, theoretical calculations have indicated that hybrid instruiiientation using both

radiation and electromagnetic interrogation for measurement of the

integrity of bridge footings is feasible.

This paper will address the theory of design of such instrumentation.

1/ NBS Special Publication 550: A Bibliography on the Corrosion and

Protection of Steel in Concrete

143
Development of standard tests and testing methods for the evaluation of

construction materials has been of interest since the turn of the century. Congress created the National Bureau of Standards in 1901. The

American Society for Testing Materials (ASTM) was formally organized and

legally chartered by the State of Pennsylvania in 1902. The two of the

first four original technical ASTM Committees formed were Committee A-1 on steel, and Committee C-1 on Cement.

The objective of this set of experiments is to theoretically derive the

mathematical formulations necessary for the optimization of radiation

wavelengths employed for gauging of composite materials: concrete re- inforced by steel. Firstly, consider the Beer-Lambert Law applied to a single component matrix. — ^ /< 1 = 1.

Where

-o

I = Transmitted Radiation Intensity

= Incident Radiation Intensity

^yA^ = Linear Attenuation Coefficient (cm ''")

X = Matrix Thickness (cm)

A

f = Matrix Density

N - Number of Atoms Mole

^

6 = Total Cross Section in Barns

A = Effective Atomic Weight in Grams

144
145
Example Instrumentation to Prove the Theory

A gamma ray transmission instrument was developed to detect blockage in

molten material transfer lines,

Periodically, problems are encountered with blockages in the lines used

to transfer molten material from one location to another in industrial

plants. The lack of productivity during the downtime required to

locate the blockages can represent a significant economic impact to the

companies affected. These periods can last as long as a week and can represent a dollar production loss of $50,000/hr.

Blockages occur as a result of impurity buildup or freezing inside the

transfer lines. Various compounds present as impurities comprise the

blockages since these compounds often possess higher melting temperatures than that of the molten material. The impurities rapidly solidify

restricting the material flow. Finally, complete blockage occurs.

( See Figure I)

.

A blocked line can be cut on a "best guess" basis and a blockage is

located by inserting a tape into the line. Oftentimes the lines must

be cut and probed in several places before the location of the blockage can be determined. It is possible to displace impurities with the

tape and temporarily remove the obstruction. The blockage can reoccur

whenever material flow is re-established in the lines.

An instrument for accurately and rapidly locating blockage based on the

gamma ray transmission technique is shown in Figure II. A highly collimated beam of gamma rays from a small radioactive source is transmitted

through the blocked transfer lines. General trends and localized anomalies in the blockage deposits are detected by observing variations in

the transmitted gamma ray beam intensity. The beam intensity represents

a quantitative measure of the thickness of blockage deposits present in

the pipe. (See Figure III).

Now consider concrete that has been attacked by sea borers. Sea borers

produce channels in the concrete which result in damage to the

structural integrity of pilings below the water line. Also consider the

semi-random placement of the steel reinforcing rods shown in Figures IV,

V.

We feel that a cross fertilization of technology is needed to solve

these problems: Radiation transmission combined with f erro-magnetic

measurements aided by a mini-computer should yield results defining the

integrity of the structure under test.

1/ D. A. Garrett, GAMMA RAY TRANSMISSION GAGING FOR THE DETECTION OF

BLOCKAGES IN MOLTEN MATERIAL TRANSFER LINES Presented at the

Spring Conference/Seminar of the American Society for Nondestructive Testing April 2-5, 1978, New Orleans, Louisiana

146
147
148
149
Figure V

150
SESSION IV

PIPELINE TRANSPORTATION SYSTEMS

Chairman: M. Lauriente, Department

of Transportation

DUCTILE FRACTURE ANALYSIS OF PIPELINES

F. Erdogan

Lehigh University, Bethlehem, PA 18015

Abstract: In this paper after briefly discussing the evolution

fracture failure in pipelines and reviewing the related basic concepts,

the problem of ductile fracture is considered. It is assumed that the

thickness-to-radius ratio of the pipe is sufficiently small so that

structurally the pipe may be treated as a cylindrical shell. It is also

assumed that the pipe wall contains a part-through crack around which

yielding takes place through the entire wall thickness. In the analysis

the plastic deformations are approximated by a perfectly plastic layer

and the bending theory of shallow shells is used to solve the problem.

The questions regarding the fracture initiation along the crack front,

the progressive growth of the crack, the plastic necking of the net

ligament, and the pipe rupture leading to leak or break are then con- sidered. The concept of crack opening stretch is used in the analysis

and the discussion of ductile fracture.

Keywords: Pipeline fracture; ductile fracture; part-through crack;

crack opening displacement; crack initiation; progressive crack growth;

plastic necking instability; leak vs. break.

Introduction ' '

• Depending on the thermo-mechanical behavior of the material and the

nature of the applied loads and the environmental conditions, in the

design of pipelines, tank cars, and a great variety of other shell

structures it is often necessary to consider fatigue or corrosion crack

propagation and fracture among the possible modes of failure. This

requires, in addition to the application of standard failure theories

specified by the existing design codes, the treatment of the problem

of acceptance and safety from the viewpoint of fracture mechanics. In analyzing the fracture failure in such structures one is likely to

encounter nearly all types of fracture ranging from the subcritical

growth of initial cracks involving practically no plastic deformations

to the final rupture of the pipe wall resulting from the net ligament

plastic necking instability. Generally, in pipes, containers, and

pressure vessels the fracture failure may evolve in the following manner: (a) A fatigue or a corrosion fatigue crack may initiate around a localized imperfection, generally a surface flaw, having the worst possible combination of geometry and loading conditions, (b) Next, this

dominant flaw may grow subcritically taking roughly a semielliptic

shape. At the initial stages of this phase the cylinder wall would be

mostly elastic and the plastic deformations would be confined to a relatively small region along the crack front only (Figure la and b).

(c) As a consequence of increased crack depth the net ligament through

153
the wall would become fully plastic and the rate of subcritical crack

propagation would increase markedly (Figure 1 c). (d) Finally, under

the peak load the net ligament would become unstable and rupture. Depending on the combination of factors which consists mainly of relative

magnitudes of the fracture resistance of the material and the crack

driving force, the resulting through crack would either be arrested or would continue to grow in an unstable fashion. These phenomena are respectively known as the "leak" and the "break".

In the initial phase of the fracture process outlined above a predictive analysis of the problem of fatigue crack propagation can be

carried out provided the stress intensity factor along the crack front

is known. This requires the solution of the three dimensional elasticity problem for a vessel with a part-through surface crack. At the

present time the problem appears to be analytically intractable. The

existing solutions of the problem are based on either the alternating

method (see, for example, the review articles by F.W. Smith, and R.C.

Shah and A.S. Kobayashi in [1]), or the finite elements [2-4], or the

method of boundary integral equations [5].

At the third and fourth stages of the fracture process (Figures 1 c,

d, and e) it is understood that prior to net ligament rupture the pipe

wall around the crack would undergo large scale plastic deformations.

Consequently, the elasticity solutions would not be adequate to study

these phases of the fracture phenomenon. In a large group of important

structures in which the thickness to radius ratio is relatively small,

the last phases of fracture usually involve relatively long and deep

surface cracks with full plastic deformations along the entire net ligament and through the container wall around the crack tips. Particularly

in structures such as the great variety of straight and bent piping and

pressurized containers used in transportation in which fluctuating

bending stresses are superimposed on stresses caused by the internal

pressure, relatively long surface cracks before rupture appear to be a rule rather than an exception. By making two major assumptions the

fracture problems arising in the failure analysis of such structures

may be brought under control and treated analytically. The first

assumption is that because of the relatively long crack length and

small thickness-to-radii ratios the structure can be adequately modeled

as a "shell". The second assumption involves the use of a modified

version of the conventional plastic strip model to account for the

plastic deformations in the crack region. In a structural component

undergoing ductile fracture which involves progressive crack growth, it

is generally agreed that the size of the energy dissipation zone Rd

shown in Figure 2 is an increasing function of the crack length. Consequently, in the case of slow progressive ductile fracture the fracture

resistance of the solid increases with the increasing crack length

(hence, the concept of "crack extension resistance curve" or the Rcurve). In progressive ductile fracture one may generally distinguish

four different regions in the material, namely the basic energy dissipation zone Rq, the wake region of the residual stresses Rj^, the crack

154
Figure 2. Progressively growing ductile fracture

155
initiation zone Ri , and the remaining elastic zone (Figure 2).

There is some experimental evidence to the effect that at the tip of a growing crack the "crack opening angle" or the "crack opening displacement" remains constant. Even though this notion has not been fully

documented and universally accepted, it may still be used to support

the conjecture that in order for the fracture initiation to take place

at the crack front, some local strength parameter must reach a critical

value- and, in the absence of a more suitable load factor, the crack

opening stretch may be considered as the measure of this parameter.

This is a necessary condition for fracture propagation and is related

to the formation, growth and coalescence of holes in the small fracture

initiation or "fracture process zone" Ri at the crack front. On the

other hand continuous (stable) growth and instability of fracture

require that the condition of global energy balance be satisfied. The

latter constitutes essentially the sufficient condition for fracture

propagation. The R-curve approach to ductile fracture is a typical

application of this concept.

The type of ductile fracture resulting from progressively growing cracks

would occur usually in structural components having a through crack or having a part-through crack with a relatively small crack depth to wall

thickness ratio (Figure 1 d and f). However, for some crack-structure

geometry and material combinations the crack opening stretch may reach

and far exceed the local critical value and yet the condition of global

energy balance may not be satisfied. Some plate and shell structures

of high toughness materials having relatively long and deep surface

cracks may fall into this category. In this case the net ligament

would generally undergo plastic "necking", and the net ligament rupture

may result from a "necking instability" under increased membrane loads.

In the following sections the problem of ductile fracture in pipes

containing surface cracks will be discussed by using these concepts

with the crack opening stretch as the primary correlation parameter.

It should perhaps be pointed out that the concept of "tearing modulus"

recently introduced by Paris appears to be a very effective way of

characterizing and classifying structural materials with respect to

their ductile fracture resistance. However, the concept is confined to

certain types of J-controlled progressive fracture and is not applicable

to net ligament plastic necking instability.

Ductile Fracture of Cylinders with a Part-Through Crack

Let a relatively thin-walled pressurized cylinder contain a surface

crack which is sufficiently deep and long so that under the given pressure the net ligament and the cylinder wall around the crack are plastically deformed. Using a plastic strip model which is basically a generalization of the conventional Dugdale model used in plane problems,

and the bending theory of shells one can calculate the crack opening

stretch along the leading edge of the crack [6]. Figures Ic and If

show the geometries of the part-through and the through crack, respectively. A sample result for the crack opening stretch 6 calculated

156
at the tips of an axial through crack on the mid-surface of the cylinder

is shown in Figure 3 where No=PoR is the membrane resultant in hoop

direction, ay is the flow stress representing the yield behavior of

the material, and the shell parameter A and the normalization factor d-] are given by

X = [12(l-v2)]^ a//Rh , d^ = 4aaY/E (1)

R, h, and 2a being the mean shell radius, the wall thickness, and the

crack length, respectively. The curve X=0 corresponds to the flat plate

results.

One can use such results in three different ways. First, if one assumes

that the critical crack opening stretch can be used as the local crack

initiation criterion, then based on Figure 3 and on similar results

obtained from a part-through crack solution a set of design curves may

be prepared which would give the load carrying capacity of the cylinder

against crack initiation. These curves are shown in Figure 4 where the

varying parameter 6/d-| correspond to the prescribed critical crack

opening stretch ratio (which is a property of the material). Note that

the results are general and may be used for any material for which the

critical stretch falls between 0.4 d] and 12 d]. Some sample results

for a part-through crack are shown in Figure 5. Here Sq is the maximum

crack opening stretch which is at the deepest penetration point (or the

midpoint) of the crack. The dots on the right hand side of the figure

correspond to the load carrying capacity of the cylinder having a through crack of the same length 2a. The fact that they are generally

somewhat greater than the extrapolated values obtained from the partthrough crack solution may indicate that in some cases at least in principle leak before break is possible.

One may also use the crack opening stretch concept to model the fracture

process in a cylinder with a fully yielded net ligament and to determine the failure pressure. Even though it is very difficult to predict whether a given fully yielded net ligament will undergo progressive

fracture propagation or plastic necking, experiments indicate that for

a progressive crack growth characterization of a fracture process the

crack opening stretch must be lower than a certain value. For example,

the results given in [7] suggest that for progressive fracture propagation the crack opening stretch at the leading edge of the crack needs to

be less than approximately 0.04 (h-d), h-d being the thickness of the

net ligament. On the other hand the results of the tests on pressurized

steel cylinders reported in [8] show that for conditions near the rupture value of pressure the opening stretch 6c at the deepest penetration point of the crack is generally greater than 0.1 (h-d). Therefore, in cylinders with a part-through crack it is possible to encounter both a progressive crack growth in the fully-yielded net ligament

and the plastic necking of the net ligament. In the safety analysis

of this type structures the most important problem is quite clearly

the estimation of the instability value of the pressure. There seems

to be no simple and unique methods currently available to calculate

157
6

7 d

0.0 0.2 0.4 0.6 0.8 1.0

Figure 3. Crack opening stretch at the tip of an axial through

crack in a cylindrical shell.

Figure 4. Load carrying capacity of a pressurized cylindrical shell

with an axial through crack against crack initiation.

158
0.00.0 0.5 1.0

Figure 5. Load carrying capacity of a pressurized cylinder with an axial part-through crack.

this pressure. For relatively thin-walled plate and shell structures

having a part-through crack with a fully yielded net ligament the two

methods outlined below appear to be very promising.

The first method refers to the calculation of the instability pressure

for the progressively growing part-through crack. Examining the crack

opening stretch curves given by Figure 3 and similar results obtained

for the part- through crack (see, for example, [9]), it may be observed

that for a given crack geometry when the membrane stress resultant

No"Pof^ reaches a certain value, a slight increase in the pressure Pq

may cause a very large increase in the crack opening stretch 6. Physically this implies some kind of a plastic instability. Thus, one may assume that the instability load corresponds to having an arbitrarily high relative slope in the 6 vs. Nq curves. Fixing this slope,

for example, at an arbitrarily high value of 20 in the units shown in

Figure 3, for each crack depth d one may obtain a curve giving the

instability load as a function of the crack length 2a (or the shell

parameter x). For an axial crack in a pressurized cylinder these

results are shown in Figure 6. One may note that, aside from the dimensions of the crack and the cylinder, the only information needed to

obtain these curves is the value of the "flow stress" ay characterizing

the yield behavior of the material. Figures 7 and 8 show the comparison of the theoretical results given in Figure 6 with the results of

burst tests obtained at Battel le Laboratories [8] for steel pipes (R = 18 in; h = 0.4 in.). Considering the complexity of the problem the

agreement appears to be quite good.

159
0.2

Figure 6. The instability load Nq = Rpo for a pressurized cylinder

with a progressively growing part-through crack (d<h) or

a through crack (d=h).

Figure 7. Comparison of calculated (Figure 6) and experimental [8]

instability loads in steel pipes with an axial through

crack.

The second method concerns the calculation of the instability load in

cylinders having a part-through crack for which the net ligament undergoes plastic necking. In this case as the pressure is increased,

unlike the progressively growing part-through crack problem in which

there is no significant change in the initial crack depth do until the

crack initiation and propagation begin, the crack depth do also in- creases. For a given pressure below the instability value, the crack

160
1.0

0.8

ha; 0.6

0.4

0.2

0

0 I 2 3 X ^ ^ ^

Figure 8. Comparison of calculated (Figure 6) and experimental [8]

instability loads in pipes with an axial part-through

crack.

depth reaches an equilibrium value, there is a stress redistribution

around the crack region, and the stress in the net ligament becomes the

plastic stress corresponding to the prevailing strain. The fracture

instability of the net ligament is clearly related to the question of

whether or not there exists an equilibrium value of the crack depth d.

Under the given pressure if d has an equilibrium value, the net ligament may undergo some plastic necking but no rupture would take place.

On the other hand, if d has no equilibrium value then fracture will occur and the (smallest) corresponding pressure is the fracture instability pressure.

In order to investigate whether or not an equilibrium value of the

crack depth d exists for a given pressure one may express d as follows

(see [6]):

d = d^ + a(e^)6^ (2)

where dg is the initial crack depth, a is a function of the average net

ligament strain e^' ^""^ is the average stretch in the net ligament.

£2 and 6z are functions of the current value of the crack depth d. &2

is calculated from the elastic-plastic shell theory and hence is a known function of d. The nonlinear equation (2) can be solved for d

by a successive approximation technique by assuming d(0) = dg and

d(N+l) = do + a(£2(N)) 6^(N) , N= 0,1,2,... (3)

Now, if d has an equilibrium value it can be obtained from (3) as the

limit for N^. Clearly for the given pressure if the successive

161
—\—,—I—>—I

'— 8 lO i2 14 ^6

Figure 9. Results of the successive approximation calculations to

determine the burst pressure (N is the number of iterations).

approximation formulated by (3) is convergent giving a finite value

which is less than the thickness h, then there would be a load redistribution around the crack accommodating a stable net ligament thickness.

On the other hand the divergence of the successive approximation technique (as defined by d exceeding h) would mean the fracture due to net

ligament plastic necking instability.

As an illustration of the technique to estimate the failure pressure in

cylinders we consider the tests carried out on full scale line pipes at

Battel le [8]. Following is the relevant information regarding the tests.

Dimensions: R = 18 in, h = 0.403 in, dg = 0.201 in,

2a = 3.8 in, o^^ = 64.6 ksi

;

Test Pressure p^Cksi): 1.0 1.2 1.25 1.29

6g (mills): 29 60 80 Failure

Here 6q is the measured crack opening displacement on the outside sur- face of the cylinder and at the center section of the crack corresponding to the pressure levels shown. In applying the technique it was assumed that

ay = + 2 ksi, a(e^) = 0.46 + 0.54 (1 - ^) . (4)

The calculation was carried out for the pressures 1.0, 1.2, 1.25, and

1.267 ksi. The result is shown in Figure 9. The ordinate in the

162
figure is the crack opening displacement 5q on the middle surface of

the cylinder and at the center section of the crack. Which one of the

quantities Sq, 6q and d is used as an ordinate in the figure is, of

course, immaterial. It is seen that the calculations for the first

three pressures are convergent meaning that these pressures are below

the instability or failure value. For pg = 1.267 ksi the calculations

diverge, implying that this pressure is equal to or greater than the

failure value. The figure also shows, for each pressure, calculated

values of the average net ligament strain and 69. Note that the

iteration technique predicts the failure pressure quite accurately.

However, particularly at smaller pressures, the predictions for 6e are not good. This is somewhat understandable, as the model is based on large scale yielding and large plastic strains which may not in fact

exist at lower pressures.

Acknowledgement: The research on ductile fracture of pipelines at

Lehigh University is being supported by the Department of Transportation

under the Contract DOT-OS-82007

.

References:

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

J.L. Swedlow, ed. ASME, New York, 1972.

2. S.N. Atluri, K. Kathiresan, A.S. Kobayashi , and M. Nakagaki

,

"Inner surface cracks in an internally pressurized cylinder analyzed by a three-dimensional displacement hybrid finite element

method", Proc. of the Third International Conference on Pressure

Vessel Technology, Part II ASME, pp. 527-534, 1977. *

3. S.N. Atluri and K. Kathiresan, "Outer and inner surface flaws in

thick-walled pressure vessels". Trans, of the 4th Int. Conf. on Structural Mechanics in Reactor Technology, Paper G5/4, CECA, CEE,

CEEA, Luxemburg, 1977.

4. J.J. McGowan and M. Raymond, "Stress intensity factor solutions for

internal longitudinal semi-elliptic surface flaws in a cylinder

under arbitrary loading", paper presented at the 11th National

Symposium on Fracture Mechanics, ASTM, Blacksburg, VA., June 1978.

5. J. Heliot, R.C. Labbens, and A. Pel 1 issier-Tanon , "Semi-elliptic

cracks in the meridional plane of a cylinder subjected to stress

gradients", paper presented at the 11th National Symposium on Fracture Mechanics, ASTM, Blacksburg, VA. , 1978.

6. F. Erdogan, G.R. Irwin, and M. Ratwani, "Ductile fracture of cylindrical vessels containing a large flaw", ASTM-STP 601, pp. 191-208,

1976.

163
\

7. J.D. Landes and J.S. Begley, in Fracture Toughness , ASTM STP 514,

pp. 24-39, 1972.

8. J.F. Kiefner, W.A. Maxey, R.J. Eiber, and A.R. Duffy, "Failure

Stress Levels of Flaws in Pressurized Cylinders," in ASTM- STP 536,

pp. 461-481, 1973.

9. F. Erdogan, "Ductile Fracture Theories for Pressurized Pipes and

Containers," Int. J. Pressure Vessels and Piping , Vol. 4, pp. 253-

283, 1976.

164
RADIOGRAPHIC VARIABLES AND WELD FLAW ANALYSIS

Robert C. Placious

Radiation Physics Division

National Bureau of Standards

Washington, D.C. 20234

Abstract: An alternative to existing workmanship codes, where failure

predictions are based on a priori experience, is a fi tness-for-purpose

criteria where failure predictions are based on fracture mechanics analysis. In this alternate "code", critical flaw dimensions form the coor- dinates for a decision curve separating failure-no failure regions. The

critical flaw dimensions under this system are the length and throughthe-wall depth of the flaw, at least for the particular case of pipeline

girth welds. This requirement to accurately dimension flaws puts additional quality control on the radiographic inspection procedures. The

flaw depth, in particular, cannot be directly determined by scaling the

radiograph. We need to measure the image contrast presented by the flaw

and relate this to metal thickness reductions. Contrast however, is con- siderably influenced by the radiographic variables. Some of these are under the radiographer's control; some are not. The variables over which

some degree of control can be exercised include x-ray kilovoltage and

filtration, film and screen selection, and film processing techniques.

The variables over which the radiographer has little control are mainly

of a geometric nature - that is, they are related to the exposure geometry. These include x-ray source size, source to film distance and material thickness for example. These geometric quantities also determine the

radiographic definition or "unsharpness" of the image. The limits on accuracy of flaw depth analysis due to these variables will be analyzed

in this paper.

Key Words: Weld flaw inspection; radiographic nondestructive testing;

flaw analysis from radiographs; pipeline radiographic inspection; flaw

depth determinations.

Introduction

The existing workmanship code for liquid and gas pipelines is A. P. 1.1104

[1]. This has been the standard for most pipeline construction in the

United States for many years. The code undergoes continuous review by

responsible representatives from API, AGA, AWS, etc. A detailed radiographic procedure is given in the code. With the exception of undercut

however, flaw depth measurements are not specified in the code.

A newer technology - fracture mechanics - may form the basis of an alternate construction code. If this approach is followed, flaw dimensioning.

165
particularly the length and depth of the flaw are important parameters of

the analysis. To the extent that such a "code" is used the need for more

accurate flaw sizing becomes significant.

Radiography, of course, is not the only nondestructive testing technique

that is applicable to weld flaw detection and dimensioning. Each of the

available techniques have advantages and disadvantages in this application. By and large, however, only radiography and ultrasonics predominate in field use. Perhaps this is because they are the most suitable

for internal flaw detection and analysis.

Table I is a listing of the advantages and disadvantages of several non- destructive inspection methods that have been used in the field for pipeline weld flaw detection and evaluation. ; , Table 1

FIQI) fCTlOK FOR FUHOWiflCraiZATlCN

(TIL". DCBITY Ont£)

oj/wiTATiVE ULimanics

U) ESTAD113ED FItLD TOCLDUt

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at Goinw ((jRicafai;jj) to m\

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Tirt ojattiic DOC icT DO (Gi uj i-^injismi

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a) TOT JtHSlTIVt mi RT OR L'TTT) mf(Z FLAWS

<2) DOES I4)T RJUK HIQtY SfaLLED

Q) DOES ICT RBlilft HldLY SWLLID

(1 ) FLAW M5T BE AT OR (fAR ai^fli

(2) DteS tCT ttATlJRE FUiW DtPm

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(2) lor nwi iCAi. roR sweiiffi large surface

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(3) STIER ru; RAH ItASURtlM TtWJ FL'W

A review of this table shows why radiography remains as the method of

choice for nondestructive inspection of girth welds. In the remainder of

this paper we will describe some of the characteristics of the radiographic process and their application to weld flaw analyses.

Girth Weld Radiography Details

Large pipelines use a panoramic x-ray source placed internally on the

pipe axes and usually aligned in the plane passing through the center of

the girth weld. Radiation is emitted in a 360° azimuthal direction.

The beam cross section is fan shaped. Films and screens are placed in

a light tight cloth package externally around the girth weld. The ex- posures, processing, and interpretations are carried out in the field.

166
In some cases the reading is made while the film is still wet. Frequently, these interpretations are made by the radiographer and perhaps a welding engineer or contractor, neither of whom are necessarily "readers"

Recent studies reveal that a certified reader when combined with good

radiographic quality control increases the probability of detecting a flaw - particularly cracks - the most difficult flaws for radiographic

detection. [2J The reader certification should be carefully controlled

and guidelines for this are given in reference 2. The problem of reader

certification should be taken into account along with the problems of

image quality control to be discussed here.

Image Quality

Table 2 lists the most common x-ray variables associated with the radiographic procedure. These variables will be discussed as they relate to

flaw detection and sizing which of course are intimately related to the

degree of image quality achieved.

The radiographer has some degree of control over all of these variables

except probably items (1) and (5). The notation R and C indicate that

the parameter is identified more with resolution (R) or contrast (C) although in the final analysis these two quantities cannot be completely

uncoupled. Resolution and contrast are the two main factors contributing

to overall image quality. These will be addressed in a general way in

the text following.

Image Resolution

There are several ways to define or describe image resolution. In optics

there is usually done in terms of the degree of separation of the intensi

ty distribution of two light sources. We choose to define it, for present purposes s in terms of "unsharpness" . This is sometimes given the

more formal designation - radiographic definition. We note of course

that the modulation transfer function is another way to define or measure

this quantity but unsharpness has the advantage of general acceptance in

industrial radiographic facilities. Therefore, we will continue to use this term in this paper.

Table 2

X-Ray Factors Affecting

Image Quality

1.

2.

3.

4.

5.

6.

7.

X-Ray source size

Kilovoltage & beam quality

Film density of image

Processing

Exposure Geometry

Film/ screen type

Operator care & experience

(R)

167
Unsharpness refers to how well a sharp edge of an object is imaged by the

radiographic system. Scanning across the image with a precision microdensitometer produces a trace which is referred to as the edge spread

function. The width of the edge spread along the "X" axis (abscissa) is

defined as the "unsharpness" of the image £3]. The degree of "edge

spread" is a function of several quantities. These include x-ray source

size, specimen thickness and source to film distances (geometric factors)

and film type, presence or absence of intensifying screens, film/screen/

specimen contact and object or x-ray source motion. The geometric quantities are generally not controllable. The latter quantities to a large

extent are controllable. Exposure geometry is usually dictated by the

system to be radiographed. In a later section, under Flaw Detection and

Sizing, the relationship of unsharpness to dimensioning will be considered.

Image Contrast

The radiographer can exercise some control over contrast. Film contrast

can be improved by proper choice of kilovoltage, film, and exposure.

Consider Figure 1 for example, where the ordinate (gradient) is the con- trast produced on the film as a function of image density. A denser

0 0 5 10 15 2 0 2 5 3C 3 5 4 0 DENSITY

Fig. 1. Density vs. Gradient for several

film types. Reference "Radiography in

Modern Industry"

80

0

168
radiographic image has better contrast for a given film type than the same image at a lower density. For this reason, most industrial radiographs are exposed to achieve densities of 2 or greater.

Another factor affecting image contrast is beam quality which relates to

the kilovoltage applied to the x-ray tube. A more accurate designation

of beam quality is the photon energy spectrum produced by the x-ray tube.

The effect of beam quality on contrast is quite easily seen if one con- siders a typical exposure chart for radiographing a material. On these

charts or graphs exposure is plotted as a function of material thickness

for different kilovoltages (see Figure 2). Taking the exposures required

Fig. 2. Typical exposure chart for steel

with kilovoltage as a parameter.

Reference - "Radiography in

Modern Industry".

for two steel thicknesses for example, we can construct a table as

fol lows :*

Radiography in Modern Industry, Third Edition, pg. 68.

169
kV

160

200

Steel

Thickness

3/4"

1"

3/4"

1"

Exposure for

a Given Density

18.5 mAM

70.0 "

4.9 "

11.0 "

Relative

Intensity

3.8

1.0

14.3

5.8

Intensity

Ratio

3.8

2.5

The higher intensity ratio at the film implies a greater contrast for

the lower kilovoltage but at a considerable cost in exposure needed.

Processing also affects contrast as is shown in Figure 3. In general,

it is wise to follow the film manufacturer's recommendations on processing since they will try to optimize the contrast sensitivity for their

own films. In general, the base and fog level of the film is generally

affected adversely by greater processing time or temperature and this can affect image quality or x-ray sensitivity.

2 40

2 20 O

<

200

4 6 8 10 TIME (MINUTES)

14

Fig. 3. Relative speed, average gradient, and fog level as a function of manual processing time for a typical film.

Reference "Radiography in Modern Industry . Since the factors just described also affect total exposure time for a given specimen, this must be taken into account in the overall selection process.

170
Flaw Detection and Sizing

Weld flaw detection and analysis is considerably improved by reader

training and experience. Scanning girth weld radiographs requires a high

degree of concentration in order to reduce the probability of not detecting a flaw because of a momentary attention lapse. Flaw detection is

also complicated by the weld bead ripple in the image which can mask a serious flaw or render its sizing more difficult. Such a flaw can be

missed on a field radiograph, particularly if its exposure geometry is

such as to place it in a shaded region of the weld image. Figures 4 and

5 are illustrations of the failure to detect a flaw - in this case a crack - on the field radiograph. This flaw was later revealed in a laboratory radiograph and in radiographs taken after the weld root and crown overbead had been removed.

Fig. 4. Photo of typical sec- tion of a field radiograph.

There is a crack in this area

.

Fig. 5. Crack in weld sec- tion shown in Fig. 4.

This was not revealed

until root and crown overbead were removed.

Most of the other types of weld flaws are more easily detected and sized

by radiography. In fact, for three dimensional flaws, such as slag and

porosity, radiography is the method of choice. There are several procedures that have been developed for flaw sizing. These include:

1. Comparative Visual Reference System [3J [5J

2. Direct Densitometry [4J [5]

3. Penetrameter/Flaw Visual Reference Method [5]

171
The comparative visual reference system utilizes a film step wedge which

relates image contrast to change in metal thickness. The details of producing the step wedge are discussed in reference [3]. The probable

accuracy of the method is discussed in [5]. Direct densitometry is used

when the flaw is large enough in projected area to exceed the size of the

densitometer aperture. One can relate four densitometry readings to the

flaw depth by using the formula cited in the reference. Two of the readings are taken on and adjacent to the penetrameter/shim image and two are taken on and adjacent to the flaw. The third method listed above is

really a variation of method number 1. It requires a very experienced

reader who can relate the contrast of the penetrameter/shim image, relative to the base metal, to the contrast of the flaw relative to its

surroundings. Fortunately, most flaws relative to their surroundings

fall in the same range of contrasts as is produced by the penetrameter/

shim thickness contrast relative to the base metal.

The basic problem with the comparative reference system is the calibration of the film step wedge. This should be done at the same time and

location as the field radiographs but generally this is not the case. Later exposures in the laboratory attempt to simulate the field conditions but some are not duplicable. Therefore, to assess accuracy of the

methods used when the original exposure conditions are not known, re- quires a series of empirical tests similar to the following recipe.

For a given geometry there are four parameters to assess; kilovoltage

(beam quality), processing, film type and x-ray exposure (beam current/

exposure time product). Therefore, each parameter is assessed independently to the extent this is possible. For example:

1. Fix processing, film type, x-ray exposure and vary kilovoltage.

Assess contrast changes on image.

2. Fix kilovoltage, processing, x-ray exposure, and vary film type.

Assess contrast changes on image, etc.

Some of these steps can be eliminated if reliable documentation of the

exposure conditions in the field is available. However, once these

assessments have been made and the reader of the field radiograph has determined the contrast for a flaw (relative to its surroundings) it is

possible to estimate how much this contrast would change due to the variables cited above.

Unsharpness, which was previously considered in the section on resolution,

is a geometric condition which compounds the analysis described above.

In the case of girth weld radiography with an internal source this quantity can be significant and must be considered in the analysis. For ex- ample a 40 inch diameter pipe with 5/8 inch wall thickness when radiographed with an x-ray source whose size is 5 mm (.2") will yield an un- sharpness value of .006" based on the formula given in ASTM E-94 [6J.

This can be a serious problem for sizing defects that are (e.g.) of the

order of .040" in width, length or depth. Note that API 1104 specifies

172
a 2T image quality on pipeline inspection unless otherwise specified.

This implies the ability to detect a flaw of the order of .025" for a wall thickness of 5/8" or .050" if 4T quality is acceptable. Yet the

unsharpness can be of the order of 10 to 25 percent of these values

simply due to geometrical restrictions. The only parameter available to

reduce the geometrical unsharpness is the source size which is generally

an option not available on field x-ray systems.

This is a brief discussion of some of the problems connected with flaw

detection and sizing. Our goal is to reduce the variability presently

existing in the radiographic inspection of pipeline girth welds when flaw

sizing is a critical need.

References

1. "Standard for Welding Pipelines and Related Facilities", API Standard 1104, Fourteenth Edition, January 1977, American Petroleum

Institute. New York, NY.

2. F.E. Sugg, NASA Tech. Briefs 2, 3, 1977.

3. Non Destructive Testing Handbook, Vol. II, R.C. McMaster, Editor,

page 19.1, 1959, The Ronald Press Co., New York, NY.

4. A.R. Hagedorn, "Radiographic Detection of Internal Undercut in Pipeline Girth Welds", Pipeline Research Summary, 1979, PR-15-95, page 13

5. "Consideration of Fracture Mechanics Analyses and Defect Dimension

Measurement Assessment for the Trans-Alaska Oil Pipeline Girth

Welds", Vols. I and II, PB260-400 & PB260-401 , National Technical

Information Service, Springfield, VA 22161.

6. ANSI/ASTM E-94-77, "Standard Recommemded Practice for Radiographic

Testing", ASTM, Philadelphia, PA 19103.

173
DEVELOPMENT OF WELDING CONSUMABLES FOR ARCTIC PIPELINES

M. Nakabayashi

Union Carbide Corporation

Linde Division

Ashtabula, Ohio 44004

Abstract

In order to meet the severe requirements of Arctic pipe quality, a series of new submerged arc welding fluxes and wires were developed over

the last several years for commercial line pipe production. Additional

efforts are now underway at Linde' s Materials Technology Laboratory to

improve upon these new fluxes and wires.

The current specification calls for weld metal impact requirement

of 70 ft. lbs at 0 °F. In order to achieve this in practical commercial

applications, pipe manufacturers are seeking weld metal impact values

of the order of 80 to 85 ft. lbs. at 0 °F. Such high impact properties

of weld metal have to be met in multiple wire, high current, high speed

submerged arc welding conditions which is typical of the current pipeline seam welding practice.

In order to attain such a goal it is imperative that control of

weld metal microstructure (achieved by controlling heat input and

chemical composition of weld metal) and substantial reduction of inclusion level in the weld metal are essential. Meeting the high weld metal

impact property requirements is further complicated by high dilution

from the base metal (60 to 70 percent is typical) which introduces some undesirable elements into the weld metal.

This presentation covers various experimental investigations from

slag-metal reactions to weld metal mechanical properties of HSLA line

pipe steels for Arctic application.

174
SESSION V

MOTOR CARRIERS

Chairman: K. Picrson, Bureau of Motor

Carrier Safety

1978 ROADSIDE VEHICLE INSPECTIONS^

Presented byKenneth L. Pierson

Bureau of Motor Carrier Safety

Department of Transportation

Washington, D.C. 20590 '

Foreword

The safety inspection of vehicles and drivers, conducted during

periods when actual highway transportation operations are underway, is one of the activities of the Federal Highway Administration's Bureau of

Motor Carrier Safety intended to increase the safety of interstate

commercial motor vehicle transportation on the Nation's highways. This

report is a compilation of the results of the Bureau's roadside vehicle

inspection activities conducted at various locations throughout the

country during calendar year 1978.

Introduction - .

The Federal Motor Carrier Safety Regulations (FMCSR) and the

Hazardous Materials Regulations (HMR) are applicable to the transportation operations of motor carriers of passengers or property engaged in

interstate or foreign commerce. The Regulations are designed to promote public safety by minimizing the risks associated with these transportation operations and thus contribute to efforts to reduce highway

accidents. The Federal roadside safety inspection program is designed

to identify vehicle defects and driver conditions which could cause ac- cidents, and to remove vehicles and drivers (that are deemed imminently

hazardous) from the highways.

Roadside vehicle/driver safety inspections may be conducted by one Federal Motor Carrier Safety Investigator working with State authorities,

or the inspections may be on a larger scale, with several Federal and

State personnel working in one or more locations. Occasionally, other

elements of the Department of Transportation as well as other Federal

agencies participate--such as the Department of Defense, the U.S. Coast

Guard, the Federal Bureau of Investigation, the Nuclear Regulatory

Commission, and the U.S. Customs Service^.

^This paper is reprinted from a publication of the Bureau of Motor

Carrier Safety, Federal Highway Administration, Washington, DC 20590.

^Certain vehicles and drivers, because of either the type(s) of cargos

or area of operations are exempted from certain Federal safety requirements; e.g., "exempt intracity" operations.

177
Vehicles may be inspected on either a selective or random basis.

The selective method involves a visible and audible prescreening by the

Federal inspector to determine if there are obvious signs of defects.

If during this cursory examination a deficiency is noted, a thorough

inspection of the vehicle and driver documentation is performed. The

selective prescreening inspection procedure is intended to identify as many defective vehicles as possible during the inspection period. The

prescreening inspection procedure is not intended to produce statistically valid measurements about the number or percent of defective

vehicles on the highway. Vehicles may also be inspected on a random

basis to provide statistically valid data for use in determining Federal

Motor Carrier Safety Program emphasis areas, and for statistical reports.

Inspectors are instructed to perform thorough inspections on vehicles in

accordance with a preplanned vehicle selection procedure designed to

insure that the vehicles are selected on a random basis.

Regardless of the selection procedure used, vehicles found to have

defects sufficient to warrant being classified as imminently hazardous

to public safety on the highway are declared out-of-service, and may

not be operated until the serious defects have been repaired. If a vehicle is inspected and found to have lesser defects that are not

imminently hazardous, those defects are noted on the inspection report

and may be repaired as soon as possible away from the inspection site.

In addition to the inspection of the vehicle, the required credentials and documents of the driver are also inspected. Most drivers^

are required by Federal regulation to have a valid operating license

issued by the State authority, a valid medical certificate, and a current driver's daily log.

The driver's log contains entries showing the number of hours

worked, the types of duties performed, the mileage traveled, and other

identifying information. The purpose of the daily log is to assure compliance with Federal regulations limiting the driver's hours of

service, and to reduce the risk of accident involvement due to excessive hours on duty or driving. Drivers who are found to be in violation of Federal hours of service limitations at the time of the roadside inspection are placed out-of-service.

In all cases where the inspection report notes defects or driver

violations, the motor carrier is required to certify within 15 days to

the Bureau of Motor Carrier Safety that the defects have been corrected.

^Drivers operating in "exempt intracity" operations or who are driving

vehicles with a gross weight of not more than 10,000 pounds and not more

than two axles are exempt from some or all of these requirements.

178
The Bureau of Motor Carrier Safety (BMCS), in continuing its effort

to improve highway safety, conducted 27,601 roadside vehicle inspections

during calendar year 1978. The total number of vehicles inspected

during this period represents an increase of 87 percent over calendar

year 1976, and an increase of 50 percent over the number of inspections

performed in calendar 1977. All vehicles inspected during 1978 were selected for inspection using the selective prescreening procedure.

Inspections of Interstate Property Carriers--Trucks

Ninety-four percent (25,965) of all vehicles inspected were operated either by for-hire or private carriers of property. For-hire

carriers of property are classified into two types, "authorized" and

"exempt." Authorized carriers are those who perform transportation

services authorized by a certificate or permit issued by the Interstate

Commerce Commission (ICC). Exempt carriers are those who transport

for hire those commodities that are exempt from ICC economic regulations.

In most cases these commodities are agricultural products, including

produce, livestock, and grain. Private carriers of property are those

who engage in transportation as an adjunct to another business enterpri se

An additional classification of carriers, "other," used in the

report includes those transporting the mails operating under contract

with the U.S. Postal service, carriers in foreign commerce, and carriers

whose classification was unknown at the time of the inspection.

Distribution of Out-of-Service Defects by Carrier Type

Of the total 25,965 property carrying vehicles inspected, 10,423

(40 percent) were found to have one or more serious mechanical defects

which resulted in the vehicle being marked and declared out-of-service

by the BMCS inspector. Table 1 displays the distribution of the out-ofservice actions by carrier type, and the average number of defects per

vehicle.

Table l--Out-of-Service Actions by Carrier Type

T^Carrier Total Vehicles

Inspected

Total Outof-Service

Average Number of

Defects per Vehicle

Authorized

Private

Exempt

Other

11 ,424

9,990

1 ,871

2,680

4,805

3,667

809

1 ,142

3.37

3.38

3.85

3.86

179
The average number of defects discovered per vehicle inspection in

calendar 1978 was 3.46. In calendar 1976 the average number of defects

discovered per vehicle was 3.23, and in calendar 1977 the average number

of defects per vehicle was 3.31.

Defects Discovered by Type and by Classification of Carrier

The 25,965 property carrying vehicle inspections disclosed a total

of 89,998 defects, 16,226 of which are classified as being imminently

hazardous and out-of-service. The greatest number (36,385) of defects

discovered was in the lighting/electrical systems of these vehicles, and

2,191 (6 percent) of these lighting/electrical defects were serious

enough to cause the vehicle to be placed out of service. The largest

number of out-of-service defects was found in the vehicles' service

brake systems. Of the total of 17,293 defects discovered in the service

brake systems, 8,128 (47 percent) were out-of-service. The largest

percentage of out-of-service defects within the service brakes related

to brake hoses or lines (33 percent), followed by inoperative brake

energy warning devices (21 percent), leaking brake chambers (10 percent),

and cracked brake drums (4 percent). Table 2 shows the distribution of

the out-of-service and less serious defects by defect type and carrier

classification.

Inspection of Interstate Passenger Carriers--Buses

Six percent (1,636) of the vehicles inspected were operated by

carriers of passengers. Passenger carriers are classified into two

types, "authorized" and "exempt." As in the case with property carrying

vehicles, authorized passenger carriers are those who perform a for-hire

service authorized by a certificate or permit issued by the ICC. Exempt

or other carriers of passengers are those performing a for-hire service

which is exempted from the requirements relating to operating authority

by certain parts of the Interstate Commerce Act, and those whose classification was unknown at the time of the inspection. - -^^

Distribution of Out-of-Service Defects by Carrier Types

Table 3 displays the distribution of out-of-service actions by

carrier type and the average number of defects per vehicle.

Table 3--0ut-of-Service Actions by Carrier Type

Carrier Total Vehicles

Inspected

Total Outof-Service

Average Number of

Defects per Vehicle

Authori zed

Exempt/Other

1 ,398

238

122

26

0.79

1.06

180
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181
The average number of defects discovered per vehicle inspection in

calendar 1978 was 0.83. In calendar 1976 the average number of defects

was 0.96, and in calendar year 1977 the average number of defects per

vehicle was 1 . 08.

Table 4 shows the distribution of the out-of-service and less

serious defects by defect type and carrier classification.

Table 4--Distribution of Defects by Type and

Carrier Classification

All Carriers Authorized Exempt & Other

Description of Out of Out of Out of

Violations Total Service Total Service Total Service

Parking Brakes 4 3 3 2 1 1 Service Brakes 1 0 1 138 71 19 1 6

Bus Requirements^ 249 1 178 1 71

Emergency Equip. 168 128 40

Exhaust System 13 11 11 9 2 2

Fuel System 3 1 2 1 1 Window Construction 119 107 12

Hazardous Matls. 1 1 Lighting/Electrical 414 7 354 6 60 1 Power Source 19 72 14 6 5 1 Steering System 4 1 2 2 1 Suspension 5 4 5 4

Ti res-Non-Steeri ng

Axle 46 4 36 4 10

Tires-Steering Axle 6 1 5 1 1 Wheel

s

80 14 61 14 19

Other Parts/

Accessories^ 73 4 63 4 10

TOTAL 1 ,361 142 1 ,108 122 253 20

^Bus Requirements i ncl ude standee line , notices to passengers, and light

on emergency door.

^Engine of vehicle unable to start without external assistance.

^Other Parts/Accessories include speedometer, horn, defrosters, mirrors,

seat belt installation for the driver, and windshield wipers.

Driver Violations

Inspection of the property and passenger carriers' drivers during

calendar year 1978 disclosed 24,481 violations of the FMCSR. The

largest percentage of the violations (52 percent) related to failure to

182
comply with one or more of the requirements of Part 395 of the FMCSR,'

Hours of Service of Drivers. The largest single type of hours-ofservice violation was the failure to maintain a driver's daily log

(4j627 violations), followed by failure to have a current daily log

(3,521 violations), and "other" log violations such as incomplete

entries and failure to have the previous 30 days' logs in possession.

Although private carriers of property accounted for only 36 percent

of the total number of vehicles inspected (FMCSR administered by BMCS

are not applicable to private carriers of passengers), 47 percent of

the total driver violations discovered are attributed to drivers of

private property carrying vehicles. In addition, these drivers

accounted for 56 percent of the most frequently found violation--failing

to maintain a driver's daily log. Examination of drivers' daily logs

for the 30 days prior to the time of the inspection disclosed 2,117

instances where the driver had violated Federal regulations which prescribe the maximum number of hours drivers may drive or remain on duty.

Of the total number of hours of service violations, 1,522 instances had

occurred on tours of duty prior to that during which the inspection was performed. Twenty-eight percent (595) of the drivers who had violated

the Hours of Service Regulations were discovered to be operating

vehicles in excess of the maximum number of hours permitted at the time

of the inspection and were consequently placed out-of-service by the

BMCS investigator at the time and place of the inspection.

Table 5 shows the distribution of driver violations by type and

carrier classification.

Vehicle Defects and Accident Involvement

Motor carriers of passengers and property are required by the FMCSR

to report to BMCS those accidents which resulted in (a) estimated

property damage to the vehicles, cargo, and fixed objects of $2,000 or more, or (b) a personal injury, or (c) the death of any person. In

calendar year 1978, 1,597 accidents were reported to BMCS as having

occurred as a result of a mechanical defect or failure. Of the 1,597

accidents involving mechanical defects reported in 1978, those defects

cited most often were defects in the brake system (496 or 31 percent),

followed by tires (429 or 27 percent), steering (83 or five percent),

wheels (82 or five percent), and coupling devices (80 or five percent).

In 1978, those accidents attributed to mechanical defects resulted in

124 fatalities, 1,305 injuries, and $19,983,560 in property damage.

Inspection of commercial vehicles in operation during calendar year

1978 showed that the five most common mechanical defects which caused

accidents during that year (brakes, tires, steering, wheels, and

coupling devices) accounted for 31 percent of the total number of

defects discovered during BMCS roadside inspections, and resulted in

70 percent of the total number of out-of-service actions taken by BMCS

staff members to remove defective vehicles from the highways.

183
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184
Concl usion

Since 1975 the percentage of vehicles inspected by the BMCS and

found to have imminently hazardous mechanical defects has increased

approximately three percent per year. The majority of these serious

defects can be readily detected by visual/audible inspection, and the

elimination of these defects can have an appreciable impact on the

Federal effort to reduce commercial vehicle accidents. In addition to

a commitment to increase the Federal inspections of commercial vehicles

in operation during 1979, BMCS recently instituted rulemaking to revise

the Federal Safety Regulations to address commercial vehicle drivers'

failure to prepare logs and inadequate inspection, repair, and maintenance procedures.

The largest single violation on the part of commercial vehicle

drivers is the failure to maintain a driver's daily log. The failure

to maintain this record of a driver's hours of service enables those

drivers who are operating in excess of the maximum number of hours

permitted to defeat attempts by Federal inspectors to determine the

actual hours of service of the driver. Effective June 18, 1979, all

drivers who are required to prepare logs will be placed out-of-service

for a minimum of eight hours if a currently-maintained driver's log is

not available for inspection during the course of Federal driver-vehicle

roadside checks.

The present hours of service limitations contained in the Federal

Regulations have been the subject of extended analysis and seven public

hearings held in major cities throughout the United States. The next

step in the rulemaking process, the publication of a Notice of Proposed

Rulemaking, is planned for the Fall of 1980.

In addition to making the hours of service rules more effective.

Part 396 of the Federal Motor Carrier Safety Regulations relating to the

inspection, repair, and maintenance of interstate commercial vehicles

has been extensively revised to insure more adequate inspection and

maintenance procedures. This revision of the Regulations will become

effective on April 1, 1980.

The Bureau of Motor Carrier Safety intends to continue its efforts

in deterring poorly maintained vehicles and fatigued drivers from

traveling the highways by expanded safety program monitoring, increased

enforcement, and the training of State personnel in inspection procedures.

185
REGULATION, LEGISLATION AND THE TEAMSTERS

An Equation for Highway Safety

R. V. Durham

International Brotherhood of Teamsters

Washington, D. C. 2 0001 '

'

Introduction

"In 1978, 5,075 Americans were killed in accidents involving heavy

trucks, an alarming increase of 53 percent over 1975. During the same period, the number of all highway fatalities increased 12.6 percent,

in itself a significant climb.

"Likewise, fatalities to occupants of heavy trucks numbered 1,010,

an increase of 41 percent from 1975. Fatalities to occupants of passenger cars rose only 7 percent during that period. This shocking mortality rate makes truck driving one of the more dangerous occupations

in America ."^

These words of Senator Charles Percy in introducing the "Truck

Safety Act" last year echo the concern raised by all facets of the

transportation community over the epidemic increase in heavy truck

fatalities. The government, vehicle manufacturers, carriers and organized labor have all recognized the crying need for concrete methods of

reducing this ominous trend.

Scope of Problem:

The problems created by heavy truck accidents are not limited to

heavy trucks. Of the 50,000 highway fatalities recorded in 1978,

17,578, or nearly 35 percent, involved a truck or bus of any size or weight. Heavy trucks represent a significant fraction of that total.

What are the reasons for this rapid increase in fatal accidents

involving trucks? In complex transportation issues like this one, there

are as many different theories as there are people available to give

them. However, of particular interest to this group is the nature and

extent of equipment failure as a factor in these highway accidents.

Whether an accident was caused by a vehicle defect is often difficult to determine, especially in the case of a driver fatality. Estimates of defect involvement range from a high of 30 percent down to one or two percent. A Department of Transportation study of large

truck accidents in Texas found that in 1973, about 6.4 percent of the

vehicles involved were reported by the police to have defects that were

connected with their accident.'- This is opposed to a two percent defect

rate for passenger cars from the same data.

The study went on to point out, "the most numerous defects reported were brakes, tires and wheel failures-about 70 percent of all defect

accident cases. Note that all of these are failures in relatively

visible components of the vehicles."^

186
This high rate of defect involvement in accidents is alarming

evidence of improper or inadequate maintenance procedures on the part

of carriers. Even more startling is the recent experience of the

Bureau of Motor Carrier Safety (BMCS) in conducting roadside inspections

of heavy trucks

.

The Bureau has the responsibility of enforcing the Federal Motor

Carrier Safety Regulations. Among other things, these standards set out

the minimum requirements for necessary parts and accessories and for

inspection and maintenance of heavy commercial vehicles. Upon discovering a vehicle in a roadside inspection which does not meet specific

criteria for safe operation (for example, a leaking air brake hose)

,

the BMCS inspector will place the vehicle "out of service" until the

defect is repaired. This means that the vehicle may not be moved

(except to tow it away from the inspection site for repairs) , until it

is placed in safe operating condition.

In the latter half of 1978 through March 1979, BMCS inspected over 2,500 vehicles during special spot inspection throughout the country.

Of this number, 46 percent were placed out of service.^ It must be

remembered that vehicles selected for inspection are usually first given

a cursory visual check. If the vehicle has a noticeable defect, it is

then moved to a safer location for a more thorough inspection. Therefore, the argument is made that only the "worst" vehicles are inspected.

But during that same period, a total of 307 vehicles were selected

for inspection completely at random. Of that group, fully 34 percent

were placed out of service. And in a recent roadcheck of over 1,7 00

vehicles traveling in the Mississippi River area, 44 percent of those

inspected were deemed unfit for service.^

Defects in the brake system accounted for the greatest number of

out of service violations, followed by tire and electrical system

deficiencies. In the Mississippi inspections, faulty brakes made up

55 percent of the total out of service violations.

What is the significance of these figures? They mean that for each

truck passed on the highway, there is a better than one in three chance

that it has a serious defect in a vital safety related system. If

that truck is involved in an accident as a result of the defect ( or

for any other reason) , it is likely that another vehicle will be involved, usually a passenger car.^ And in accidents of that nature, cars come out on the short end of the stick.

In collisions between cars and tractor-trailers resulting in a fatality, one study has shown the car occupant death rate to be ten

times higher than the truck rate.^ Another review of several studies

has shown that for each truck occupant fatality in crashes involving

trucks, there are 30 to 40 other highway users killed. •^'^ The bottom line is that trucking accidents are not only of grave

concern to our organization as representatives of the truck driver,

but also must be the concern of all parties interested in highway

safety

.

With the scope of the problem evident, the question now is, how

to -keep defective and unsafe vehicles off the road? The title of this

paper refers to three distinct groups: the Executive Agency, the

187
Legislature, and the Union. All three play a vital role in arriving

at solutions to the problems presented by mechanical failures. None

by itself can adequately deal with all aspects of the problem.

To further explore this concept, it is helpful to examine a few

specific examples in which the interaction of these groups is leading

to concrete methods of preventing mechanical failures due to vehicle

defects.

The Executive Agency;

A recent Department of Transportation (DOT) study on preventive

maintenance isolated four factors necessary for the alleviation of com- mercial vehicle defects r'^'^

1. Detection of Defect

2. Communication to Responsible Party

3. Repair

4. Reporting to the Detector that Repair has been Effected.

A post-trip inspection by the driver and his completion of a vehicle condition report were, at the time of the study (197 5) , the

sole required general inspection duties of the carrier. Accordingly,

the DOT study recommended both a pre and post-trip inspection, preferably conducted by the driver, as a method of detecting defects and

communicating them to the carrier, (steps 1 and 2 detailed above).

The DOT study also suggested that the driver's post-trip inspection report, complete with notation of maintenance activity, be carried

on the power unit for the review of the next driver.

In light of these recommendations , and in response to our experience as representatives of hundreds of thousands of truck drivers,

the Teamsters Union formally petitioned BMCS for a revision of the

Federal Motor Carrier Safety Regulations concerning maintenance, inspection, and repair.

Specifically, we asked BMCS to require a pre-trip visual inspection

in addition to the already mandated post-trip inspection; to require

that the post-trip inspection report be reviewed by maintenance personnel and that listed defects be corrected, with the action taken noted

on the inspection report; and that the completed inspection report be

then placed on the power unit for the review of the next driver.

The chronology of subsequent events is a case study in the oftentimes frustrating regulatory process. BMCS responded to our June, 1976

petition by issuing an Advanced Notice of Proposed Rulemaking in April

of 1977 . This Notice explained the scope of the problem and asked

interested parties for comments and suggestions.

After lengthy and laborous study of the numerous comments received

in response to the Notice, BMCS promulgated a Final Rule on July 2,

1979, to take effect on August 31, 1979.-^^ The final rule incorporated

the major recommendations of our petition. However, the industry was concerned that aspects of the new rule would raise operating costs without enhancing safety. To allow time for consideration of their protests, BMCS delayed the effective date four months to December 31,

188
1979. ^ A minor change was made in the wording of the regulation*, and

BMCS again delayed the effective date to allow carriers to bring their

operations into compliance, this time establishing April 1, 1980 as

1 Pi the new date.

By this time, all parties appeared satisfied with the present state

of the regulation. We felt that our petition was adequately addressed

and most carriers found they could comply with only minimal changes in

their recordkeeping and maintenance operations.

However, less than a month before the rule was to take effect, the

United Parcel Service (UPS) went into Federal Court asking for both a review of the regulation and a stay of the effective date pending

review. '^ UPS felt that the requirement that the post-trip inspection

report be placed aboard the power unit would be prohibitively expensive

and unnecessary for their type of operation, and that the driver's need

for that information would be best served by allowing him daily access

to the report at the dispatch site. The Department of Transportation

responded and the Teamsters intervened in the suit, pointing out to

the court both the logistical problems presented by the UPS proposal

and the well documented advance in highway safety which the new regulations would bring about.

The case is still pending in the Circuit Court of Appeals, but

the court did deny the UPS motion for a temporary stay. Therefore, the new regulations took effect the first of this month.

This rulemaking is an example of how the Teamsters Union and governmental agencies can work hand in hand to try to solve some of the

problems presented by mechanical failures in trucks. We have also

worked closely with various agencies on a wide range of issues, including steering, tire loading, air brake systems, and overall vehicle integrity. But regulation is only one entry in our "highway safety

equation"

.

Role of Legislature;

Some problems are best addressed through legislation. An example

of such a problem is the limited scope of the authority of BMCS . The

authority of BMCS in enforcing the Federal Motor Carrier Safety Regulations is basically restricted to placing violating vehicles out of

service, although they may levy civil penalties for drivers who fail

to properly maintain a daily log as required.

This method of enforcement has proven to be inefficient and ineffective. Even using their resources as best possible, BMCS was able

to inspect less than one percent of three million commercial vehicles

under its purview in 1978.-^^ This low inspection rate provides no credible deterrent for unscrupulous carriers who try to increase profits

at the expense of safe vehicle maintenance. Faced with only the pro-

* The requirement that drivers "certify" that repairs had been made was deleted because "the certification of repairs must be done by a mechanic

with skill and experience in the technical areas" and because the re- quirement would "place an unfair burden on drivers" (44 Fed. Reg. 76525)

.

189
spect of occasionally repairing a vehicle placed out of service by a BMCS inspector at a roadside station, some carriers adopt a "breakdown" maintenance policy, repairing vehicles only when they actually

cease to operate.

The authority of BMCS in enforcing their regulations is limited

by their enabling legislation, passed in 1937. Therefore, in order for

them to obtain improved enforcement authority, it was necessary to

repair to the legislature.

Over two years ago, the Teamsters Union entered into discussion

with Senator Charles Percy (R.-Ill.) concerning the feasibility of

introducing legislation to allow BMCS to more effectively enforce their

own regulations.

The fruit of those discussions was S. 1390, the Truck Safety Act

of 1979. The bill in part allows the Secretary of Transportation

(presumably through BMCS) to levy civil penalties of up to $10,000

for non-compliance with federal safety regulations. In addition, it

authorizes the Secretary to prosecute persons who knowingly and willfully violate the standards, and provides a maximum jail term of one year and a maximum fine of $2 5,000 upon conviction of such violators.

The bill addresses several other important issues which are not of

great interest here, such as the working conditions of drivers. But one other provision of the bill stands to substantially reduce the number

of defective vehicles on the highway. It would protect trucking em- ployees from discipline, discharge or discrimination if they refuse to

drive an unsafe vehicle or if they report safety violations to state or federal authorities.-^^

There presently exists a similar provision in the National MasteiFreight Agreement, negotiated between the Teamsters and the carriers.

While over dne-half million trucking employees are covered by that

agreement or one with a similar provision, there are thousands of drivers who don't enjoy the benefits of a Teamster contract and need the

protection provided by this piece of legislation.

The Truck Safety Act passed the Senate by an overwhelming margin

in March, and is currently under consideration in the House. Its passage there will mark a great stride toward safer trucks.

A number of other issues are before the legislature which may have

an impact on the vehicle defect problem, including the relaxation of

economic regulation in the trucking industry and length and weight laws.

Accordingly, the Teamsters Union attempts to put forth the concerns of

our membership whenever legislation affecting their interests is con- templated . The Union Role:

The hand of the government, either executive or legislative, is

sometimes necessary to reach solutions to our problems. But often

problems can be solved without the aid of the state, through collective

bargaining

.

Long before the introduction of legislation providing protection

against discrimination for employees who refuse to drive unsafe vehicles,

the National Master Freight Contract had such a provision. To take

190
another example, just as a BMCS inspector can place a vehicle "out of

service", many contracts allow a Teamster business agent to "red tag"

an unsafe vehicle with much the same effect.

The collective bargaining process can deal with issues at a speed

seldom seen in either the legislature or the executive branch. One

instance of this ability centers around the controversy over the safety

of the so-called "cab-under" tractor-trailor configuration.

When the Teamsters Union first learned of the development of a truck design featuring the driver's compartment under the cargo unit,

the initial reaction among officers with truck driving experience was

that such a vehicle would be unsafe for a variety of reasons , including

lack of forward visibility and vehicle crashworthiness . However, those

conclusions were based on a "truck driver's common sense" approach to

the issue, with little real engineering basis.

Accordingly, we enlisted the aid of the Highway Safety Research

Institute of the University of Michigan, requesting an analysis of the

safety and handling properties of the "cab-under" configuration. The

o 1 resulting study was the basis for contract negotiations for the

latest National Master Freight Agreement. That contract now contains

a clause reading: "No driver shall be required to drive a tractor

designed with the cab under the trailer."

While the Teamsters also petitioned the National Highway Traffic

Safety Administration for a rule prohibiting operation of such a vehicle

we were able to adequately protect our membership from a hazardous

vehicle without reliance on the government. Since Teamsters will not

drive the "cab-under", its commercial application is extremely limited,

and only a few units are now in operation.

The losses incurred through accidents resulting from vehicle

defects, in terms of both human resources and property damage, are too

great for society to ignore. This paper has sought to briefly explore

the existing and future potential for a three-pronged attack on faulty

and defective vehicles.

Through the cooperation and coordinated efforts of the Congress,

the Department of Transportation, and Organized Labor, great strides

have been made and will continue to be made toward the ultimate goal of

safe highways.

191
APPENDIX "A"

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193
FOOTNOTES

1. Congressional Record - Senate , June 21, 1979, p. S.8259.

2. National Highway Traffic Safety Administration, U. S. Department

of Transportation, Fatal Accident Reporting System, 1978.

3. McDole, Thomas L. , et al , Effect of Commercial Vehicle Systematic

Preventive Maintenance on Specific Causes of Accidents.

Ann Arbor, Michigan: The University of Michigan, Highway Safety

Research Institute, July, 1975, p. 31.

4. Ibid . , p. 6.

5. Bureau of Motor Carrier Safety, "Major and Special Emphasis

Roadside Vehicle Inspections, August, 1978 - February, 1979," p. 1.

6. Bureau of Motor Carrier Safety, "Mississippi River Roadside Inspection, May 14-18, 1979," p. 4.

7 . Ibid . , p. 5.

8. Perchonok, Kenneth and Ranney , Thomas A., Analysis of Truck , Tractor/Trailer Accident Data , Buffalo, New York, Calspan Corp.,

June, 1975, p. 3.

9. Baker, Susan, et al

,

"Fatal Tractor-Trailer Crashes: Considerations

in Setting Relevant Standards , " Paper presented at the Fourth

International Congress on Automotive Safety, July 14-17, 197 5.

10. Perchonok, op . cit

.

, p. 11.

11. McDole, op . cit

.

, p. 11.

12. Durhame, R. V., letter to Dr. Robert Kaye , June 2, 1979.

13. 2 Fed. Reg. 18103.

14. 44 Fed. Reg. 38523.

15. 44 Fed. Reg. 50042.

16. 44 Fed. Reg. 76526.

17. United Parcel Service v. United States , 80-1268.

(D.C. Cir . , 1980)

.

194
18. Congressional Record - Senate , June 21, 1979, p. S.8259.

19. Senate Bill, S.1390, Section 11(b).

20. National Master Freight Agreement, Article 16, Section 1.

21. Bunch, Howard M. , et al , "An Evaluation of the Safety and Handling

Properties of the Cab-Under Truck Tractor Vehicle," Ann Arbor,

Michigan: The University of Michigan, Highway Safety Research

Institute, June, 1978.

22. National Master Freight Agreement, Article 16, Section 6(h).

1^5
STRESS SYSTEMS RELATED TO FRACTURE OF

DUCTILE AND BRITTLE MATERIALS

Donald J. Wulpi

International Harvester CompanyFort Wayne, Indiana 46801

In order to understand how various types of fractures are caused, one must understand the forces acting on the materials as well as the

characteristics of the materials themselves. All fractures are caused

by stresses, and the "weakest link" theory applies; that is, damage

will occur when the stress on the weakest element exceeds its strength.

Also, by understanding the ways in which single load, or monotonic,

fractures are caused, one can then better understand fatigue fractures,

which are the result of many thousands or millions of load applications.

To understand the forces, it is necessary to study the stress systems

acting on the part. It is very useful to study the stress systems acting on a cylindrical member, such as a rod or shaft. A variety of

stresses can be applied to a cylindrical member; also, the same principles can be applied to non-cylindrical parts. Shafts, and shaftlike parts, are very common parts and are widely used in construction

of many assemblies and machines.

Stress systems are best studied by examining free-body diagrams which

are simplified two-dimensional models of complex three-dimensional

stress systems which are internal in the shaft. The figure shows the

orientations of the normal stresses (tension and compression) and the

shear stresses (sliding) which are 45° to the normal stresses. The

free body diagrams of shafts in the pure types of loading—tension,

torsion, and compression--are the simplest; they can then be related

to more complex types of loading.

Tension Loading

When a shaft or similar shape is pulled in tension, it gets longer and

narrower, just as does a rubber band when it is pulled. Similarly,

the square in the free body diagram is elongated in the direction of

the tensile stress and is contracted in the direction of the compressive stress, as is shown at the left in the sketch. Note that the

shear stresses are at 45° angles to the axial tensile stress and the

transverse compressive stress. Also, note that there are two sets of

shear stresses, each perpendicular (90°) to the other, diagonally

between the normal stress directions.

Ductile materials, by definition, are those that deform because the

shear stress exceeds the shear strength before any other type of

damage takes place. That is, the shear strength is the "weak link"

in the system, and is the controlling factor. Therefore, under a

196
tensile force, the internal crystal structure of the metal shears, or

slips, microscopically on the millions of shear planes in the metal,

resulting in lateral deformation, commonly called "necking down",

prior to final fracture. This necking phenomenon occurs in the plastic

regime, which means that the deformation is irreversible. Of course,

plastic deformation is characteristic of any fracture of a ductile

material. An interesting fact is that fractures of ductile metals in

pure tension originate at or near the center of the shaft, toward the

late stages of the fracture process. Many tiny internal cracks join

by the process of "microvoid coalescence" to form a rough, jagged

fracture surface. As the crack progresses outward, it eventually

reaches a region near the surface that forms a 45° wall, or "shear

lip", around the periphery of the fracture. This forms the familiar

"cup and cone" fracture that is typical of tensile fractures of

ductile materials. Incidentally, whenever a 45° shear lip is seen on any fracture, it locates the end of the fracture, or the last part to

fracture.

Brittle materials, by definition, are those that fracture because the

tensile stress exceeds the "cohesive" strength before any other type

of damage can take place. Now, the cohesive strength is the "weak

link" in the system, and is the controlling factor. Brittle materials

always have a fracture direction that is perpendicular (90°) to the

tensile stress, and have little or no deformation because fracture

takes place before the metal can deform plastically, as in ductile

materials. Thus, a tensile fracture of a brittle material has a direction that is essentially straight across the shaft. It has a bright, sparkling appearance when freshly fractured, characteristic

of brittle materials. Since the magnitude of the controlling tensile

stress is essentially uniform across the shaft, fracture can originate

at any location in the cross section, in the absence of stress

concentrations

.

Torsion Loading

When a shaft is twisted in pure torsion, the same stress system

exemplified by the free body diagram—rotates 45° in one direction or

the other, depending upon which way the shaft is twisted. When

twisted as shown in the figure, the stress system rotates 45° counter

clockwise. Note that the normal stresses (tension and compression

components) are now at 45° to the shaft, while the shear stresses

are now longitudinal and transverse. Each pair of stress components

are mutually perpendicular to each other.

In a ductile material, the shear strength is again the "weak link"

when the shear stress exceeds the shear strength. Again, plastic, or permanent, deformation occurs although it is not obvious, as is the

necking down effect in tension. The shape of a cylindrical part, such

as a shaft, is not changed by torsional deformation. An example will

explain why

:

197
Imagine that the shaft consists of an infinite number of infinitely

thin disks. When the pack of disks is twisted, each disk slips a very

small amount with respect to its neighbors. However, the diameter of

each disk does not change, because of the slippage on the transverse

shear plane. Eventually, fracture occurs on this transverse shear

plane, which is essentially the sliding face between two adjacent disks.

In pure torsion, the final rupture will be at the center of the shaft,

although it will be offset toward one side if a bending stress is also

present. Defonnation also occurs on the longitudinal shear plane, but

this does not noiniially cause fracture unless the material is weak in

the longitudinal direction.

A brittle material in pure torsion will again fracture perpendicular

to the tensile stress component, which is now at 45° to the axis of the

shaft. This spiral type fracture is characteristic of brittle metals,

including case hardened shafts, which are subjected to torsional loading. Other brittle materials, such as glass or chalk, will also fracture in a spiral manner if twisted.

The elastic stress distribution in pure torsion is maximum at the sur- face and zero at the center. Thus, fracture normally originates at

the highest stressed region—the surface— in pure torsion. This is

true for all of the stress components: normal (tension and compression)

as well as the shear stresses as is shown in the figure.

Compression Loading

When the shaft is loaded in axial compression (assuming a shaft that

does not buckle) the stress component system again rotates 45° so that the compressive stress component is now axial, while the tensile

stress component is transverse, as is shown in the figure. Now the

shear stress components are again 45° to the shaft, as they were during

tension loading,

A ductile material in compression does the reverse of what it did in

tension: it becomes shorter and thicker under the action of the slippage on the shear planes. In short, it bulges when squeezed by the

compressive force. This is characteristic of metals being cold headed,

pancake forgings, and other ductile materials under axial compression.

A brittle material in pure compression will, as always, fracture

perpendicular to the maximum tensile stress component. Since the tensile stress component is now transverse, the brittle fracture direction

is now longitudinal, or parallel to the shaft. Brittle materials, such

as extremely hard metals, glass, chalk, and rock will split or shatter

longitudinally when loaded in compression. Indeed, this is the principle of rock crushing.

The elastic stress distribution in pure compression again is the reverse

of that in tension: uniform across the section (assuming no stress con- centration) , except in the compression direction, as shown in the figure.

198
Bending Loading

When a part is loaded in bending, the convex surface has a free body

diagram stress system similar to that shown here for tension.

Conversely, the concave surface will be stressed in compression, and

will have a stress system as shown for compression. Approximately

midway between the two surfaces will be the neutral axis where all

stresses are zero, while the surface stresses are at a maximijm. Thus,

failure can be expected to originate on the convex side of the bend

where the tensile stress exists.

Fatigue

The preceding discussion was concerned with single load deformation and

fracture. Fatigue is unique in that the magnitude of the multiple load

applications are usually not high enough to cause plastic deformation.

That is, the stresses are relatively low. However, upon many, many

repetitions microscopic changes take place in the crystal structure

that leads to formation of fatigue cracks.

The essential thing to remember is that the propagation, or growth, of

fatigue cracks over a large period of time, is in exactly the same direction as the rapid crack growth of brittle materials under the

same type of loading. That is because fatigue cracks propagate in a direction that is perpendicular to the principal tensile stress —the

brittle fracture direction.

Thus it can be seen that fatigue cracks will propagate essentially

straight across a shaft or other surface stressed in tension, such as the convex side of a shaft or other part in bending. Similarly, a shaft under repetitive torsional loads will form fatigue cracks 45°

to the axis of the shaft. Reversed torsion—twisted alternately in

both directions —will form two sets of fatigue cracks, each at approximately 45° to the shaft, or approximately 90° to each other. This is

the action that forms the classic "starry" fatigue fractures of a splined shaft in which many cracks, each at 45° to the axis, occur during reversed torsional fatigue. If axial compressive forces cause

fatigue cracks, they will be parallel to the axis of the shaft;

however, this type of loading is rare.

Summary

By keeping in mind the principles outlined here, the analysis of single

load and fatigue fractures can be better accomplished. The principles

are always the same; however, confusion can be caused by misinterpretation and uncertainty as to the type of loading, particularly under

combined loading conditions.

199
200
PERSPECTIVES ON DIAGNOSTIC SYSTEMS

FOR THE TRUCKING INDUSTRY

V.H. Miles

American Trucking Associations, Inc.

Washington, D.C. 20036

The paper is based on a survey which was conducted among the

members of the American Trucking Associations Maintenance

Council. The primary purpose of the survey was to obtain an

indication of how state of the art technology could be best

applied to resolving the trucking industry's maintenance

problems. Additionally, the survey was designed to assess the

industry's receptiveness to new and sophisticated maintenance

aids. This paper, therefore, presents a summary of the results

of the survey and the conclusions we have drawn. The key point

of the paper is that the maintenance of several vehicle systems

such as electrical including the charging and starting system,

engines, and brakes constitute a major cost factor to the

industry. This cost can most likely be reduced significantly by

the use of state of the art diagnostic systems. While the

trucking industry appears to need this capability, it will be up

to the manufacturers to develop the concepts and demonstrate

their value.

Key words: Automated test equipment; diagnostics; technology in

truck maintenance; truck maintenance aids.

AS many of you may be aware, the trucking industry is

rather conservative when it comes to technological innovation. We

believe, as do managers of other successful businesses, that the

first objective of a company is to make a profit and since we

operate on a very narrow profit margin there is a propensity to

invest only in the tried and proven. There is very little

individual research and development work related to vehicles

sponsored by motor carriers. In fact, only a few trucking

companies are large enough to have engineers assigned solely to

such R&D functions. This then means that it is not only

incumbent upon the manufacturers who supply us with our

equipment to conceive and develop new concepts, but they must

also demonstrate the concepts' value to us. In the area of

diagnostics this is a major challenge. When it comes to offering

the industry sophisticated computer controlled systems which

instantaneously provide vehicle systems analyses, tell mechanics

what to do to effect repairs, issue the parts, and generate the

201
required record, 1 have found a level of skepticism among most

truckers that only my education in the "Show me" state could

have prepared me for. This situation does not, however, diminish

our interest in finding better and more efficient ways to operate,

but it does significantly affect our rate of progress and the

manner in which we proceed.

To improve our maintenance capability, the industry has

established working groups such as the ATA Maintenance Council

and the Technical Advisory Group to affect coordination with

government and the manufacturer; to aid in identifying our

needs or, stated another way, our problems; and to continue this

coordination in the development and testing of the solution. This

process obviously leads to evolutionary, rather than revolutionary, changes in the industry. The development of diagnostic

systems for commercial trucks has and, no doubt, will continue

to adhere to this principle.

Our interest in the subject of diagnostics is not new. It arose during the early seventies when it became clear that motor

vehicles would become much more complex as governmental rulemaking began to dictate engineering designs. Antilock braking

systems and emission controls confirmed this contention. To help

deal with this situation, the position 1 currently serve in was

created and a Task Force of representatives from trucking and

the manufacturers was formed. The purpose of this Task Force

was to look into the areas of diagnostics in general and antilock

brakes in specific. While the antilock issue has been settled in

the courts, there remains a nucleus of capable, forward thinking

people who have been exposed to the state of the art in

technology for diagnostics. These people have formed a Diagnostic

Task Force for the purpose of continuing their efforts to determine the applicability of this technology to the trucking industry

and to promote its acceptance where appropriate.

Following the premise that the onus for new product

development is on the manufacturers, the Diagnostic Task Force

has centered its efforts on first trying to determine where the

industry can best benefit from diagnostics, secondly , on attempting to assess the industry's receptiveness of new and more

sophisticated maintenance aids, and thirdly, to promote the

awareness of a broad spectrum of the industry's maintenance

managers and executives to what the state of the art in

technology has to offer.

To accomplish our first two goals, we recently conducted a survey among the trucking firms which participate in the Maintenance Council. The survey questionnaire was quite comprehensive, consisting of sixty-three questions, but was not designed

to derive exact analytical results. Realizing that fleet executives

202
have very limited time for such matters, we asked that they

provide the best data that was readily available and that where

time consuming searches and computations were required, an

estimate would suffice. Therefore, I must characterize the results

of this survey as ballpark indications.

With this inherent limitation, we sent out approximately 410

surveys to which we received 50 responses. 1 was somewhat

disappointed initially with the limited response, but after some

discussions with others who have conducted surveys within the

industry 1 found that a 12% response is considered good.

Again, 1 must stress that this low response is not a reflection of a lack of interest in improved maintenance capabilities within the industry, but rather a combination of factors.

First is the strenuous workload of the maintenance mianagers who

must give priority to their daily operations. Second, as 1 eluded

to earlier, it must be recognized that many of our present

maintenance directors started as helpers or mechanics and while

they know their trucks to the nth degree, they may not have the

scientific training which would allow them to understand and

appreciate the benefits of complex advance technology systems.

Last and certainly not the least important is the detailed twelve

page survey which might tend to discourage all but the most

zealous from responding.

Although there may be some bias, since prople most interested in the subject are most likely to respond, we do feel that the

results are reasonahh^ r<=presentative of the industry. At any

rate, it's the best information we have. The size of the firms

varied from one as small as twenty-five vehicles to a firm

owning 85,000 vehicles. A total of approximately 140,000 vehicles

are represented. 1 should note that for the purposes of the

survey a vehicle could be a pickup & delivery vehicle, a straight truck, a tractor, or a trailer. The firms also varied in

type of business and the manner in which the vehicles are

utilized. Approximately 1/3 of the firms were common carriers,

1/3 were private carriers, and the remaining 1/3 consisted

mostly of leasing companies and specialized haulers.

In structuring the survey we tried to anticipate the more

profitable ways of using diagnostics. Among these were:

1. Identifying pending failures in the incipient stage so that repairs could be effected before costly damage

occurred.

2. Simplifying maintenance by readily identifying the

problem and thus reducing manhours expended.

203
3. Reducing lost revenues and operational delays, by

precluding enroute breakdowns.

4. Enabling the trucking firms to meet ultimate EPA

in-service standards by maintaining their vehicles within

prescribed parametric limits.

To assess the correctness of these conjectures, we designed

our questions to acquire an indication of the industry's maintenance and operating philosophy, identify their maintenance

problems, and assess their preference for diagnostic systems.

In examining the industry's maintenance and operating

philosophy we found that on the average all vehicles, with the

exception of trailers, are retained for approximately five years.

This equals about 160,000 miles on P & D vehicles, 190,000 miles

on straight trucks, and 550,000 miles on tractors. Trailers are

generally kept for 10 years and several firms retain them

throughout their useful life.

This rate of turnover is a significant indicator of a fleet's

need for an extensive maintenance investment and also indicates

how quickly a firm might acquire newer vehicles which may be

more readily amenable to the use of automated diagnostics. Let

us take a look at tractors as an example. As 1 just mentioned,

the average trade-in point for the original owners of tractors is

550,000 miles. The average engine life before a major overhaul

is required is approximately 300,000 miles. Transmissions last

for about 350,000 miles and differentials and axles require

overhaul or replacement at about 315,000 miles. As you can

surmise, the vehicle is essentially rebuilt at about the 300,000

mile point and traded before the second major overhaul is

imminent

.

With surging increases in the cost of money, the declining

availability of capital, and the slowing of the economy, this

vehicle turnover pattern could change. It may prove to be wiser

or possibly necessary for many carriers to accomplish that

second round of vehicle rebuild. This then would extend the

average original ownership of a vehicle to approximately 900,000

miles or about nine years. An alternative approach would be to

try to extend the overhaul interval through greater emphasis on

preventive maintenance and the use of more sophisticated prognostics and diagnostics. Additionally, 1 would anticipate, and

we are seeing to some degree, a more intense emphasis on

"spec'ing" for greater overall reliability.

Actually, all carriers have a scheduled Preventive Maintenance Program, However, the schedules appear to be as diverse

204
as there are trucking firms. According to our survey each

tractor receives a PM inspection on the average every 11,400

miles. This compares well with a report by Burlington Fleet

Services, a data processing service for trucking firms, which

states that linehaul tractor PMs are conducted on an average of

every 12,111 miles. Further, it should be noted that most fleets

use a graduated inspection program where through a cycle of two

to three inspections, the extent of the inspection increases as a

function of the mileage interval increase. For example, some

companies have A, B, and C inspections which are conducted

every AOOO, 12000, and 2A000 miles respectively. The details of

these inspections vary from 190 items on an A inspection to 254

on a C.

While each fleet may have its own distinct schedule and

inspection procedure, one point is certain—that all PM checks

are time consuming and their productivity is limited by the

ability of the mechanic and capability of his test equipment. For

this reason, we believe that the development of a system to be

used to accomplish quick, revealing PM checks would be one of

the most cost effective applications of diagnostic systems. In

fact, a prototype of such a system for diesel engines has been

developed by HaiiiLitun lest Systems and is presently undergoing

in-service testing with the New York Transit Authority. In this

test program the bus engines are tested each day as the bus

proceeds through the checklane. A computer printout is produced

which lists the measured engine operating parameters and flags

any that exceeds predetermined limits. This type of system

inherently provides the capability to track the history of each

parameter. Such a running historic record in turn provides the

maintenance manager with the information necessary to predict

failure and schedule maintenance to accomodate his operational

schedule yet preclude catastrophic failures.

The possible cost benefits of such a system is indicated by

the fact that each tractor, 'as an example, is out of service for

maintenance 14 days a year. This is about a 5% out 'of service

rate which must be reflected in lost profits. Since profits from

each tractor average about $170 a day, when multiplied by 14

that comes to a $2380 loss annually. While this out of service

rate could never be reduced to zero, a reduction of only two or

three days could result in significant cost savings, especially

for the large fleets. Even for a company with only 100 tractors,

a two day reduction in the out of service rate could result in a

$34,000 a year increase in profits.

Another area where we expect a payoff in the use of

advanced diagnostics is in mechanic productivity. From our

survey we found that when a problem occurred, the mechanics

205
spend about 18% of their time locating the cause of the problem.

Further, approximately 7-5% of the vehicles which have had

maintenance performed return within 30 days for repair of the

same or a relates uroolem. The ultimate diagnostics systems

which would be acceptable to the industry should allow the

mechanic to use a quick and simple procedure to identify and

isolate the cause or causes of a problem and make a lasting

repair. As you will see when we discuss the preference portion

of our survey, one of the maintenance managers' primary goals

is to be able to " fix it right the first time ."

This leads us to ask about the capability of the mechanics

and in particular to attempt to assess their adaptability to new

techniques. Presently, most mechanics perform all general

maintenance functions. Only a few firms have specialists. Our

survey indicates that most firms perform all their maintenance

including major overhauls. The exceptions are primarily body

and framework and precision machining. The average experience

level of the mechanics is 10.65 years, although this varies from

as low as two years with one company to as high as twenty-five

years with another. About 55.6% of the mechanics have received

factory or technical training and 80% of the companies provide

periodic training for their mechanics. Although most companies

have only relatively rudimentary test equipment, several

companies have invested in more elaborate test gear, especially

test equipment for the electrical starting and charging system.

In fact, nine of the 50 responders own dynometers. Every

com.pany reported that they have not experienced any bias on the

part of mechanics against new test gear. Actually, most

mechanics are both willing and able to use the equipment.

These findings lead us to believe that the commercial

trucking industry will readily adapt to more modern test equipment and techniques. In reaching this conclusion, we are not

disregarding tne difficulty some automotive manufacturers have

experienced in trying to promote new approaches such as plug-in

diagnostics. The fact is that the motivation of the auto service

mechanic and his shop manager for maximum efficiency is

probably not as great as it is in a company shop which

maintains its own vehicles. Also, through "spec'ing," most fleets

keep a reasonably standardized fleet. Thus, the fleet mechanics

are not required to work on as large a variety of vehicles.

While attempting to identify the major maintenance

problems, we found that by far the most frequently identified

problems were electrical. Among these, the most prevalent problem was burned out light bulbs. Now, a burned out bulb is easy

to identify and doesn't require any special diagnostic system.

However, it can be difficult to replace, create a safety hazard.

206
and result in enroute delays and fines. Since this does not

require a diagnostic remedy, why even mention it? The reason is

that it points out a consideration which we must keep in mind

and one that 1 believe is the key to this conference. That is,

that improvement in the reliability of an item which frequently

fails may very well be more cost effective than providing a better way to maintain it. In the case of lights, we believe that

with the reliability achieved, in today's sealbeam headlamps that

the technology must be available to provide long lasting lamps

for the other lighting needs on commercial vehicles.

The other major electrical problems involve alternators,

batteries, starters, wiring and connectors. Development of a quick and simple capability for a pretrip check of the starting

and cnarging systeiu would most likely result in a significant

savings by reducing highway breakdowns. Better than half of

our road service calls for all types of commercial vehicles are

in response to electrical system problems. The average tractor

requires 2.7 road service calls per year with an approximate

cost of $90 per response. This equals $243 per year per tractor.

Considering that in 1977 there were 1.2 million tractors registered in the U.S., it is clear that reducing electrical system

problems will have high potential for cost benefits.

The second major problem area identified was the engine.

The major causes of engine failures were worn rings, bushings

and injectors. Certainly all metal parts are destined to wear

and failure, but effective diagnosis of impending premature

failures with immediate corrective maintenance will very likely

cause an increase in the average life of these engines and the

possible concomitant extension of the use of the vehicle. We are

all familiar with the old cliche "But for a nail,* etc." well, in

the real world of trucking an appropriate expression could be

"But for a $10 ^art a $_000 engine was lost." The maintenance

managers are concerned with this problem. For this reason

approximately 75% of those responding to our survey are using

oil sampling analysis to attempt to anticipate failures.

The third most significant area of complaint was the brake

systems. Brake system maintenance is needed on the average,

every 164,000 miles. This requirement is attributed primarily to

air line failure and lining wear. Here, as with the light bulbs,

the tendency is to conclude that diagnostics m.ay be of little

benefit and emphasis should be placed on designed-in reliability,

however, there may be some imaginative approach such as a plug-in system for checking valves, timing, and balance which

might contribute to lining life. Here again it will be up to the

ma^nufacturer to show us that there is a better way—that we can

move away from the hammer and screw driver type of inspections.

207
While survey results indicate that manual transmissions for

diesel tractors are lasting 350,000 miles, we are not convinced

that this is a true representation of the life of these components.

Some of our fleets have been experiencing an unusually short

operational life with some of the new transmissions which are

designed to mate with the high torque rise engines. Many of the

problems we have seen are typical new product failures which we

believe will soon be eliminated. Thus it is possible that the

average life of transmissions which have been debugged are a

lot longer. As an indication of this, two of our responders

reported an average transmission life of 700,000 miles. Ideally,

we would want to see drive train component last approximately

600,000 miles, which exceeds the average trade-in point. 1 am

not aware of any diagnostic test equipment that has been

developed for the drive train. Here again, the emphasis might

best be placed on reliability rather than repair.

One additional area of concern is tires. Although we did

not gather sufficient data relative to tire failures to draw any

conclusions, the trucking industry is interested in the benefits

derived from properly inflated tires such as fuel savings,

reduced tire wear, as well as safety. For this reason we support

the efforts of NHTSA and the manufacturers to develop tire

pressure monitors. We ask only that before these devices are

mandated that their reliability be adequately proven. To aid in

this process, the ATA Maintenance Council has been fleet testing

the commercially available devices and will continue these tests

into the foreseeable future. The results of the tests are available

to NHTSA. It is in our best interest as well as the government

rulemakers that we not repeat the billion dollar mistake that we

encountered with anti-lock brakes.

We concluded our survey with a series of questions designed to determine the industry's preference for diagnostics.

From these questions we found that the most desired

application of diagnostics is to detect problems in the charging

and starting system. The second most preferred application is for

the engine, followed by one for the brakes.

When asked to select their primary reason for purchasing

diagnostic equipment, the answer selected by a sizable majority

was: to reduce n-^nec^ssr.ry repairs and parts replacements. The

other choice in the order they were rated were:

o Improve mechanic productivity and reduce training.

o Reduce maintenance costs by reducing road calls.

208
o Improve fuel economy.

0 Comply with Clean Air Standards.

The relative emphasis given to Clean Air Standards simply

reflects the fact that the maintenance manager must view his

more immediate problems with greater seriousness. At present

except for some rudimentary smoke opacity limits, there are no

Clean Air Standards for in-service diesel trucks. If emission

standards do ever come into being it will most likely be more

feasible and reasonable to enforce them by placing acceptable

limits on the performance parameters of the engine. Monitoring

these parameters will not only alert the vehicle owner when a

limit is exceeded but should allow immediate diagnosis of the

problem. We believe that this approach will more satisfactorily

prevent air pollution from heavy duty vehicles as well as aid

the fleet operator in maintaining his vehicles and in gaining

maximum utilization from his fleet.

When asked whether they would want an on-board system,

off-board system, or both, only one responder selected on-board.

The remainder split evenly between off-board and both.

From some of the comments received both in the survey and

in our discussions with members of the industry, we believe that

the apparent bias against on-board electronics is a result of the

antilock episode. Due to this stigma, any on-board application of

electronic devices will require explicit proof of reliability. In

this regard, we are working through the SAE to develop an

Electical Equipment Environmental Recommended Practice for

Trucks similar to the J1211 standard for automobiles. 1 believe

as more electronic sensors and controls are added to commercial

vehicles that use of a certain level of on-board sensing as well

as diagnostics will be inevitable.

1 can foresee a number of advantages to having on-board

sensors. Not the least of these would be the time savings in test

system hook-up and the possibility for instantaneous readout. A

prototype 32 parameter system has been developed by Rockwell

and has received extensive testing in the government and

industry sponsored heavy-duty vehicle fuel economy test program.

To the best of my knowledge there are no plans at this time to

market this system, but a much simpler system is on the market.

Again, 1 believe that proven reliability and cost will be the

driving factors in gaining industry acceptance of such a system.

In assessing how the maintenance managers preferred to

use the equipment, we found that three-fourths of those responding felt that the equipment should remain in the shop rather

209
than be portable. They also preferred that the equipment be in a single unit, but be movable within the shop. Interestingly,

only five wanted to use diagnostic systems in the check lane. 1 relate this lack of interest in a more innovative approach to my

experience in the Air Force. As history has shown the horse

soldiers were not overly enamoured with the tank, nor did the

leaders of the Army and Navy rush forth to embrace Gen.

Mitchell's concept of Air Power. Well, this sort of perception of

change is still with us. Today's fighter pilot is not overly

receptive to long range standoff weapons launched from big

aircraft. The point is, as Alvin Toffler detailed in his book,

Future Shock, that most people are more content with the status

quo. The introduction of sophisticated diagnostic equipment will

change the mode of operation of most companies but it will

probably be for people not intimately associated with the daily

operation of a fleet maintenance facility to develop and show the

advantage of these changes. The New York Transit Check Lane

Test Program which 1 previously mentioned is a good example.

The growth of the use of computers and data processing in

the trucking industry will help facilitate this change. Although

computers are extensively used in the accounting side of the

trucking industry, their use is still budding within the maintenance area. Approximately half of the responders have

computer-based maintenance data systems, but of those who do

not, more than half are converting or plan to convert to

computers in the near future. The remainder felt that their

company was too small to make the use of a computer economical.

Not too astonishing then was the response of more than

half of the companies that they wanted a diagnostics system

which would feed information directly into a central maintenance

data system.

Finally, the one area that surprised me the most was the

response to the question concerning how much they would be

willing to pay for a diagnostic system. They could choose

between less than $5000, $5000 to $10000, $10000 to .<1^20000, and

over $20000. Most of the responders felt the system should cost

between $5000 and $10000. However, two were willing to pay in

excess of $20000. Of course the value of these figures is no

better than the commitment of the responder. But it does indicate

that contrary to my previous impression, maintenance managers

are willing to invest in expensive equipment if the cost effectiveness is there.

1 believe the industry is on the threshold of making

dramatic changes in their maintenance concept. With the expanded use of diagnostics by the engine producers and the

210
development of electronic feedback systems for pollution control

and fuel economy, the field use of diagnostics will be a natural

follow-on. The interest is there. The challenge is for the

manufacturers to plan the development of these maintenance aids,

now, as the changes in vehicle design are evolving so that they

can provide us with systems that meet our needs when those

needs arise.

211
INVESTIGATION OF DEFECTS IN

MOTOR CARRIER ACTIVITIES

Robert F. Hellmuth , National Highway Traffic Safety Administration ; Washington, DC 20591

Congress recognized in writing the National Traffic and Motor Vehicle

Safety Act of 1966 that safety standards could not possibly cover everything that could be wrong with an automobile. It therefore

created the Office of Defects Investigation, whose task it is to

discover and investigate potential safety-related defects in motor

vehicles and motor vehicle equipment.

The NHTSA receives some 5,000-5,000 owner complaints every month. In

previous years, most of these complaints were in letter form, but as our nationwide toll-free Hotline becomes better known, the trend is

shifting from letters to phone calls. The "Auto Safety Hotline" is

accessible, on a toll-free basis, to all vehicle owners and to those

desiring information (or wishing to comment) on vehicle safety

defects. This toll-free number (800-424-9393) can be used by callers

from the 48 contiguous States (Alaska, Hawaii, Puerto Rico, Guam and

the Canal Zone cannot use this number)

The purpose of the "Auto Safety Hotline" is to solicit accurate and

timely information from the public on vehicle safety problems and to

help us gather potential safety defect data. Reports from the Auto

Safety Hotline are used as an early warning to identify problems and

trends in specific vehicles and items of equipment and to assist in

defect investigations and recall actions. Copies of this data are also furnished to the appropriate manufacturers.

In addition to the Hotline we have some 2,500 repair shops across the

Nation involved in a Parts Return Program where suspect parts are furnished for our engineering analysis to identify the existence of

safety-related manufacturing and design defects. In the 7 years this

program has been in operation, we have received some 7,000 suspect

parts.

All such inputs are coded for computer storage when received. The

type of information stored includes owner name and address, vehicle

make, model, year, VIN, components involved, fault, and whether an accident or injury resulted. The letter is then microfilmed,

primarily so that we have a permanent copy and secondarily for

logistical purposes. At the present time, NHTSA has over 500,000

inputs in our computerized complaint file. The physical storage and

retrieval problems of this many documents necessitates our microfilming and use of computer systems.

212
An initial screening of all incoming inputs separates out those with

significant potential safety problems. These are then further screened

and analyzed by the Engineering Analysis Division to define the defect

and determine if any are candidates for a formal investigation. This

analysis is performed by an engineer assigned to that particular item,

and usually includes contacting the vehicle owner, checking our computer file for other similar complaints, determining if a service

bulletin or previous recall campaign applies to the subject problem,

contacting the manufacturer, and in some cases testing as necessary.

Once the defect has been defined and it appears to be significant, it

is discussed at a defects review panel meeting. The group, which

consists of engineers, legal counsel, and consumer services, decides if

a formal investigation is warranted. If so, the matter is turned over

to the Defects Evaluation Division for processing. At the time an investigation is opened, the manufacturer is immediately notified by

telephone, and a press release is often prepared which outlines the

basis of the case and describes the alleged defect.

The Defects Evaluation Division is charged with the management and

completion of the formal case investigation. The case investigator

builds the case, typically through owner interviews, information

requests to the manufacturer, R. L. Polk surveys, additional testing

by our Ohio Test Facility or other contractors and any additional

information received from the public from consumer advisories and press

releases. It is not unusual for the number of consumer complaints on

a particular item to increase a hundred-fold after the initiation of

a formal investigation is announced in the media.

If an initial determination is made, the manufacturer and other

interested parties, whether consumer groups or private individuals, are given the opportunity to present their views in a well publicized

public meeting. After the meeting, our Administrator makes a final

determination based upon the investigation file and the comments from

the public meeting. If the Administrator's decision is that a safety-related defect does exist, the manufacturer is ordered to

notify vehicle owners and provide remedy in accordance with the Act.

If a manufacturer disagrees with the determination, he can resort to

the judicial process for relief.

213
EFFECTIVENESS OF NHTSA ' S DEFECTS PROGRAM

IN THE MOTOR TRUCK AREA

In order for NHTSA to effectively exercise its authority in the area

of safety-related defect recalls, the cooperation of the vehicle

driver, vehicle owner or fleet, the Bureau of Motor Carrier Safety,

Teamsters, and the vehicle and component manufacturers is needed. This

cooperation involves communication at six levels

:

1. Vehicle drivers, if they do not own the vehicle,

should communicate problems they are experiencing

on the road to the vehicle owner, the vehicle owner's

maintenance personnel, or possibly to the NHTSA.

2. Vehicle owners or fleets should report suspect defects

to the NHTSA or the vehicle or equipment manufacturer.

3. BMCS field inspectors should report thru channels, defects

detected

.

4. Vehicle and equipment manufacturers must report any

known safety-related defect to the NHTSA.

5. Teamsters safety personnel should report suspect defects

to the NHTSA.

5. Vehicle and equipment manufacturers must correct any

safety-related defects in a timely manner after a recall campaign has been established.

214
APPENDIX

MECHANICAL FAILURES PREVENTION GROUP

31st Meeting

Failure Prevention

in

Ground Transportation Systems

ATTENDANCE LIST

Swaroop Amaravadi

Dept. of Metallurgy, U-136

University of Connecticut

Storrs, CT 06268

John Anderson

Spectral Dynamics

Box 152

Parsippany, NJ 07054

D. A. Bracher

Mantech International Corp.

2341 Jefferson Davis Highway

Arlington, VA 22202

James P. Bruner

Research & Technology, Armco Inc.

P.O. Box 600

Middletown, OH 45043

Harry C. Burnett, Jr.

Executive Secretary-MFPG

National Bureau of Standards

Washington, D.C. 20234

Rustam Chinoy

Mechanical Technology Inc.

968 Albany/Shaker Road

Latham, NY 12110

Douglas B. Chisholm

HRS-11

Dept. of Transportation

Washington, D.C. 20590

Ivan Debernardo

Dofasco

P.O. Box 460

Hamilton, Ontario L8N-3J5

R. V. Durham

Internatl. Brotherhood of Teamsters

25 Louisiana Ave., N.W.

Washington, D.C. 20001

Robert J. Eiber

Room 11-4-072

Battelle-Columbus Laboratories

505 King Avenue

Columbus, OH 43201

Francois Envent

Mantech International Corp.

Suite 930

2341 Jefferson Davis Highway

Arlington, VA 22202

Fasil Erdogan

Dept. of Mechanical Engineering

Lehigh University

Bethlehem, PA 18015

Daniel Fairchild

Fram Corp.

55 Pawtucket Avenue

East Providence, RI 02916

J. Ferguson

Federal Railroad Admin., RR D-33

400 7th Street S.W.

Washington, D.C. 20590

217
John Fisher

Associate Director

Fritz Engineering Laboratory

Bethlehem, PA 18015

John L. Frarey

Shaker Research Corp.

Northway 10 Executive Pike

Ballston Lake, NY 12019

Thomas P. Groeneveld

Battelle-Columbus Laboratories

505 King Avenue

Columbus, OH 43201

John M. Hanson

Wiss, Janney, Elstner & Assoc.

33 Pfingsten Road

Northbrook, IL 60062

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Center for Materials Science

National Bureau of Standards

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National Bureau of Standards

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December 6, 1968

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June 25-26, 1969

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BIBLIOGRAPHIC DATA

SHEET (See instructions)

PUBLICATION OR

REPORT NO.

NBS SP 621

2. Performing Organ. Report No. 3. Publication Date

October 1982

4, TITLE AND SUBTITLE

Failure Prevention in Ground Transportation Systems . Proceedings of the

31st Meeting of the Mechanical Failures Prevention Group, held at the

National Bureau of Standards, Gaithersburg , MD, April 22-24, 1980.

5. T. Robert Shives and William A. Willard, Editors

6. PERFORMING ORGANIZATION (If joint or other tlian NBS. see instructions)

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

7. Contract/Grant No.

8. Type of Report & Period Covered

Final

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street. City. State, ZIP)

Center for Materials Science, National Bureau of Standards, Washington, DC 20234; Office

of Naval Research. U.S. Dept. of the Navy, Arlington, VA 22217; National Aeronautics and

Space Administration, Goddard Space Flight Center, Greenbelt, MD 20771; Naval Air Systems Command, U.S. Dept. of the Navy, Washington, DC 20360

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11. ABSTRACT (A 200-word or less (actual sumnnary of most significant information,

bi bliography or literature survey, mention it here)

If document includes a si gnificant

These proceedings consist of 18 submitted entries (16 papers and 2 abstracts) from the

31st meeting of the Mechanical Failures Prevention Group which was held at the National

Bureau of Standards, Gaithersburg, Maryland, April 22-24, 1980. The theme of the

symposium was failure prevention in ground transportation systems. Areas of interest

included rail vehicles and structures, highway and road bridges, pipeline transportation systems, and motor carriers.

12. KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semi colon s) bridges; diagnostic systems; failure; failure detection systems; fracture; fracture

control; ground transportation; motor carriers; pipelines; rail structures; rail

vehicles; reliability; transportation systems.

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