NATL INST. OF STAND & TECH

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p'' NBS TECHNICAL NOTE 1088

t 9 NBS

PUBLICATIONS

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

Fitness-for-Service Criteria for

Assessing the Significance of

Fatigue Cracks in

Offshore Structures

Yi-Wen Cheng

100

U5753

No. 1088

1985

L
rwi

m he National Bureau of Standards' was established by an act of Congress on March 3, 1901. The

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NATIONAL BUREAU

OF STANDARDS

I.IBRARY Fitness-for-Service Criteria for %

Assessing the Significance of Fatigue ^^^^^^^ Cracks in Offshore Structures ^^': ^'f

Yi-Wen Cheng

Fracture and Deformation Division

Institute for Materials Science and Engineering

National Bureau of Standards

U.S. Department of Commerce

Boulder, Colorado 80303

Sponsored by

U.S. Department of Interior Minerals Management Service

12203 Sunrise Valley Drive

Reston, Virginia 22091

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

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

Issued August 1985
National Bureau of Standards Technical Note 1088

Natl. Bur. Stand. (U.S.), Tech Note 1088, 72 pages (Aug. 1985)

CODEN:NBTNAE

U.S. GOVERNMENT PRINTING OFFICE

WASHINGTON: 1985

For sale by the Superintendent of Documents, U.S. Government Printing Office, Washington, DC 20402
CONTENTS

Page

INTRODUCTION 1

AN AUTOMATED FATIGUE CRACK GROWTH RATE TEST SYSTEM 4

Abstract > 4 Introduction 5 The FCGR Test Method 5 Equipment for the Automated FCGR Test System 5 Applications 8 Constant-Load-Araplltude FCGR Test 8 Near-Threshold FCGR Test 10 Variable-Load-Amplitude FCGR Test 13 Summary 13 Acknowledgments 13 References 13

THE FATIGUE CRACK GROWTH OF A SHIP STEEL IN SALTWATER UNDER SPECTRUM LOADING 15

Abstract 15 Introduction 16 Load Spectrum 16 Simulation of Load-Time Histories 17 Experimental Procedures 18 Test Material and Specimens 18 Test Apparatus and Environment 18 Loading Conditions 19 Experimental Results and Discussion 20

Constant-Load-Amplitude Tests 20

Spectrum-Loading Tests 20

Conclusions 28 Acknowledgments 29

References 29

ESTIMATION OF IRREGULARITY FACTOR FROM A POWER SPECTRUM 31

Abstract 31 Introduction 32

Irregularity Factor 32

Evaluation of Irregularity Factor from Power Spectra 33

Direct Integration of Power Spectral Density Function 33

Estimation from Characteristic Width and Center Frequency of the Power Spectrum .... 34

iii
CONTENTS (continued)

page

Summary 37 Acknowledgments 37

References 11 Appendix 41

FATIGUE CRACK GROWTH IN AREAS OF STRESS CONCENTRATION — PLASTICITY AND SMALL-CRACK EFFECTS .. 43

Abstract 43

Introduction 44 Experimental Procedures 45 Test Material 45

Specimen Preparation 45

Loading Conditions 47 Test Environment and Crack-Length Measurements 47

Experimental Results and Discussion 47 Strain Survey 47 FCGR at Edges of a Yielded Hole in Air 48 FCGR at Edges of a Yielded Hole in Saltwater 51 Summary and Conclusions 51 Acknowledgments 51 References 54

HIGH/LOW STRESS AMPLITUDE EFFECTS ON FATIGUE CRACK GROWTH RATES OF A SHIP STEEL

IN AIR AND IN SALTWATER 55

Abstract 55

Introduction , 55

Experimental Procedures 56

Experimental Results and Discussion 57

Effects of High/Low Stress Amplitude on FCGR in Air 57

Effects of High/Low Stress Amplitude on FCGR in Saltwater 60

Effects of Tensile Overload Prior to Crack Initiation on FCGR 60

Summary and Conclusions 62

Acknowledgments 62

References 65

iv
FITNESS-FOR-SERVICE CRITERIA FOR ASSESSING THE SIGNIFICANCE

OF FATIGUE CRACKS IN OFFSHORE STRUCTURES

Yi-Wen Cheng

Fracture and Deformation Division

National Bureau of Standards

Boulder, Colorado 80303

Results of a research program to develop f itness-for-service criteria for assessing the

significance of fatigue cracks in offshore structures are presented in five papers. Each paper describes the goals and approaches to a specific task and details the results of the study. Highlights

of these results are:

1

An automated fatigue crack growth rate test system has been developed that can be used for

tests under constant- and variable-amplitude loadings.

2. A digital simulation of random signal from a power spectrum characteristic of the North Sea

environment was developed to apply random-amplitude loading. Experimental results showed that

the average fatigue crack growth rates under spectrum loading and under constant-amplitude

loading were in excellent agreement when fatigue crack growth rate was plotted as a function of

the appropriately defined equivalent-stress-intensity range. This procedure is equivalent to applying Miner's summation rule in fatigue life calculations.

3. Simplified methods of evaluating the irregularity factor of a power spectrum were derived.

4. The linear-elastic fracture mechanics analysis of fatigue crack growth at stress concentrations

appeared to be adequate in yielded regions that are caused by monotonic loading but which have

a linear stress-strain relation under cyclic loading. Small-crack behavior was observed in an ABS grade EH35 steel when crack length was less than 1.5 mm in air and less than 3 mm in 3-5

percent NaCl solution. The problem of small-crack behavior was accounted for by adding an intrinsic crack length to the physical crack length.

5. The tensile overload retardation effects on fatigue crack growth rates were similar in the

elastic regions and at the edges of a yielded hole and also in air and in 3-5 percent NaCl

solution. Tensile overload prior to crack initiation appeared to retard subsequent fatigue

crack growth in areas of stress concentration. The retardation was explained by the presence

of beneficial residual stresses and crack closure.

A subsequent report will incorporate these technical findings into a procedure for calculating

fatigue crack growth in offshore structures under realistic conditions.

Key words: fatigue crack growth; f itness-for-service criteria; fracture mechanics; offshore

structures; random-loading fatigue.

INTRODUCTION

Offshore structures are designed to withstand the cyclic stresses caused by wave loading. However, the presence of undetected defects in the as-fabricated structure may cause fatigue cracking

during service. If a fatigue crack is detected during the in-service inspection of an offshore

structure, a methodology is needed to assess its significance with respect to the safety of the facility. This need was clearly identified by the Committee on Offshore Energy of the Marine Board.

The committee's recommendations to the U.S. Department of the Interior for implementing a postinstallation inspection program for fixed steel structures included: "Develop procedures and standards for acceptance of remedial actions carried out in response to the results of an inspection." In consideration of this need, a research program is being conducted at the National Bureau of

Standards for the Minerals Management Service with the following overall objective:

To develop a technical basis for recommending remedial actions to be taken in the

event that a fatigue crack is discovered during in-service inspection of an offshore structure.

The basis on which remedial actions are to be taken is the fatigue crack growth (FCG) behavior

under service conditions. The desired representation of FCG behavior is a fatigue crack lengthversus-time curve; figure 1 is an example. FCG behavior is best described by the fracture mechanics

approach, in which fatigue crack growth rates, da/dN, are correlated to the applied stress intensity

factor range, AK:

TIME, t

Figure 1. Schematic fatigue crack growth behavior with respect to time.
da— = C (aK)"'

dN

(1)

where C and m are material parameters determined from laboratory specimen tests, and aK is a function of the applied stress range, aS, the crack length, a, and a geometrical factor, F:

aK = AS /tt a F (2)

Integration of equation (1) gives crack length- versus-time (a-versus-t) results.

From equations (1) and (2), one must know the values of C, m, aS, a, and F to evaluate FCG

behavior. With reference to figure 2 and equations (1) and (2), the basic ingredients for fatigue

damage assessment are service loading history (reflected in aS), component K solution (reflected in

F), material da/dN data (reflected in C and m), and crack length, a, found by inspection of the

structure. Two problems associated with the service loading history are calculation of stress

range, and cycle counting in cases of irregular loading histories, such as sea loadings on offshore

structures. The determination and modeling of each piece of information relevant to evaluation of

FCG in an offshore structure is a major task by itself. Even with all the basic information available, crack growth calculation is not as straightforward as an integration of equation (1) because of

load-sequence interaction effects, which must be understood and taken into account.

Owing to the complexity involved in fatigue analysis and limited resources available, the major

portion of this research program has focused on development, improvement, and verification of models

for FCG analysis under random-amplitude loading. The specific output of this program is a procedure

to calculate the stress range and count cycles from which aS and N in equations (1) and (2) are determined. The research results are given in five technical papers that follow in this report.

Service

Loading

History

Stress Range

Calculation

& Cycle

Counting

Component

K-Soiution

Material

da/dN

Data

Cracl<

Length

Calculation

Fatigue

Damage

Assessment

Figure 2. Fatigue damage assessment diagram.
The additional detailed information needed for accurate FCG analysis of specific offshore

structural components, such as service loading histories, stress redistribution due to crack growth,

component K solutions, and material da/dN data, must be sought in the literature, especially the

results from the European research program on the fatigue behavior of welded tubular joints in offshore structures for the North Sea, and from the fabricators and operators- of individual structures.
AN AUTOMATED FATIGUE CRACK GROWTH RATE TEST SYSTEM"^

Yi-Wen Cheng

David T. Read

Fracture and Deformation Division

National Bureau of Standards

Boulder, Colorado

Abstract

An automated fatigue crack growth rate (FCGR) test system has been developed that can be used

for tests of constant-load-amplitude FCGR above 10~° m/cycle (ASTM E647-83) at normal (-10 Hz) or low (-0.1 Hz) cyclic frequencies and for tests of near-threshold and variable-load-amplitude FCGR.

The test system consists of a minicomputer, a programmable arbitrary waveform generator, a servohydraulic test frame, and a programmable digital oscilloscope. The crack length is measured

using the compliance technique; the FCGR and the stress intensity factor range are calculated and

plotted automatically during the test.

Key words: automated test system; compliance technique; fatigue crack growth rate; fatigue of

materials; near-threshold fatigue test; variable-load-amplitude fatigue test.

To be published in ASTM STP 877.
Introduction

Fatigue crack growth rate (FCGR) data are used for material characterization and for fracture

mechanics reliability analysis of structures subjected to cyclic loading. A standard test method

_ Q for measuring such data above 10 m/cycle under constant-amplitude loading has been developed and

published in the 1983 Annual Book of ASTM Standards under the designation ASTM E547-83.

With the increased interest in near-threshold FCGR [1 ,2] and FCGR under environmental

influences at low cyclic frequencies [3], the demand for FCGR measurements has increased. Obtaining

such data can be tedious and time-consuming. An automated FCGR test system, such as that described

in this paper, allows testing to proceed, data to be taken, and loads to be altered in the absence

of an operator. The automated test system minimizes testing time. Data scatter is reduced owing to higher precision in crack length measurement and better control in data point spacing [4]. Because

the testing is interactive and automatic in nature, the procedure is relatively easy to follow and

requires minimal operator training. Finally, this approach eliminates the subjective interpretation

and influence of the experimenter.

The FCGR Test Method

The sequence of the FCGR test is: first, obtain the raw data, namely, fatigue crack length, a, versus elapsed fatigue cycles, N; second, reduce this data to a plot of da/dN versus AK, where da/dN

is the FCGR in m/cycle and AK is the crack-tip stress-intensity factor range in MPa/m. Typical

outputs are presented in figure 1

.

The number of elapsed fatigue cycles can be obtained from counters (electronic or mechanical)

or conversion from time elapsed at the actual testing frequency. The methods of crack length

measurement are complicated and have been a subject of extensive study [5,6]. Although several

methods of crack length measurement have been developed, some require specialized equipment not commonly available in mechanical testing laboratories. The compliance technique, however, requires

only monitoring of the load cell and the clip gage outputs, which is routinely achieved in

mechanical testing. Compliance is defined as the specimen deflection per unit load, which is a function of crack length for a given material and specimen geometry. The load and deflection

signals (voltages) can be interfaced to a computer. Because of the simple instrumentation and the

need in our laboratory for environmental chambers for cryogenic-temperature and saltwater corro- sion-fatigue tests, the compliance technique was chosen to measure the crack length.

Equipment for the Automated FCGR Test System

A schematic of the automated FCGR test system is shown in figure 2. Figure 2 also shows the

sequence of operation and interaction between various components. The test system consists of a closed-loop servo-controlled hydraulic testing machine, a programmable digital oscilloscope, a programmable arbitrary waveform generator, and a minicomputer.

Numbers in brackets denote references listed at the end of each paper,
E

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3.5 _ Stainless Steel Casting

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o m. 2.5 _ /

z

T3

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2.0

( 00°°

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1.5

1 n 1 1

1 2x10^

N, cycles

10r?

io-«

10

oo

o

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s> cP

20 50

AK, MPaVm

Figure 1. Data outputs from the automated FCGR test.

The machine control unit, which is included in the hydraulic mechanical test machine, includes

a servo-control system, a feedback system, two dc conditioners, and a valve drive. A non- programmable function generator with an electronic pulse counter is usually built into the machine

control unit of a commercially available mechanical testing machine. Signal amplifiers and a load

cell are also included in the mechanical testing machine.

The programmable digital oscilloscope contains two 15-bit 100-kHz digitizers, and it serves as

an analog-to-digital (A/D) converter. In addition to its high speed A/D conversion rate, this

oscilloscope features the ability to instantaneously freeze and hold data in memory. These features

eliminate the problems encountered with slower A/D converters, such as interruptions during the test

[7] and low test frequencies [8]. For the near-threshold and the variable-load-amplitude FCGR

tests, load levels vary with time and a programmable function generator is needed. For these tests,

a programmable arbitrary waveform generator is used. At present, this programmable waveform

generator is not connected to a cycle-counting device, and the fatigue cycle counts are inferred

from the cyclic frequency and elapsed time, as given by the computer. For the constant-loadamplitude FCGR test, a built-in function generator is used.

Included in the minicomputer are a cathode-ray-tube (CRT) terminal, a line printer, a dual

The frequency used for cycle calculation is checked with a frequency meter; the typical error in

frequency is 50 ppm.
Clip Gage

Specimen

Amplifier Hydraulic

Actuator

Machine

Control Unit

i

Programmable

Digital

Oscilloscope

Programmable

Arbitrary

Waveform

Generator

Processor

CRT

Terminal

Line

Printer

Floppy

Disks

Digital

Plotter

Figure 2. Schematic of the automated FCGR test system.
floppy disk storage unit, and a digital plotter. The minicomputer uses the 1 6-bit word and has 128K

words of memory. The minicomputer also contains an internal clock that reads to 1/60 s. The

IEEE-488/GPIB is used for the interface between the computer and the programmable digital oscilloscope and between the computer and the programmable arbitrary waveform generator.

Applications

In the following discussion, we describe how this test system is used to conduct the FCGR

tests. Requirements on grips, fixtures, specimen design, and specimen preparation are detailed in

ASTM E647-83 and other proposed standards [3,9] and are not discussed in this paper.

Constant-Load-Amplitude FCGR Test

The test system described in the previous section can be programmed to run the constant-loadamplitude FCGR test. The operational details below implement the procedures set forth by ASTM

E647-83.

As shown in figure 3, the operator feeds the input parameters into the computer through the CRT

terminal. The input parameters include specimen identification, specimen dimensions. Young's

modulus, selected time interval for measuring crack length, minimum load level for compliance

measurement, load levels, and test frequency. The time interval for measuring crack length must be

kept small enough so that every increment of crack growth will not exceed the recommended values

prescribed in ASTM E647-83. The minimum load level for compliance measurement is used to eliminate

possible crack closure effects [10], which have a significant effect on the accuracy of crack length

measurement.

The precracked specimen is fatigue cycled under the prescribed loading conditions and cyclic

frequency. A typical frequency is 10 Hz. When the preselected time interval (typical value is 1 min) for crack length measurement is reached, the computer requests the load-versus-deflection data

from the programmable digital oscilloscope, which freezes the load-deflection data in the memory

instantaneously and correlates the data to a straight line using a linear least-squares fit. A linear correlation coefficient of 0.999 or better is usually obtained. From the resulting

compliance, the instantaneous crack length is computed by using the appropriate expression for the

compliance calibration of the specimen [11]. The precision of the crack length measurement is

typically within 0.04 mm.

The inferred crack length, which is obtained from the measured compliance, published compliance

calibrations, and published Young's modulus, usually does not agree exactly with the actual crack

length for a given material and specimen geometry. The exact reasons for the discrepancy between

the inferred and the actual crack lengths are not clear and have been discussed in references 12 and

13. Correction factors to the compliance calibrations have been used to obtain more accurate

physical crack length predictions.

Another way of correcting the mismatch between the inferred and the actual crack lengths is to adopt an "effective modulus" for the material. The effective modulus, E^^^ ^ ^g deduced from a known
Input Parameters

'

Cycling

•*

'

Time to Measure No

Crack Length?

No

, Yes

Crack Growth Increment

Large Enough?

' Yes

Calculate

N, AK, da/dN

' r

Print, Plot, Store Results

, '

Figure 3. Summary flow chart of the automated constant-load-amplitude FCGR

test.
crack length in a given specimen geometry and compliance calibration. Typically, Eg^j. is deduced

from one well-defined crack front that is visible on a posttest fracture surface. The crack front

at the end of fatigue precracking or at the final fatigue crack length is generally used. The

effective modulus approach, which is used in our labaratory, thus forces agreement between the

inferred and the actual crack lengths and compensates for any errors regardless of source [13],

The portion of the load-versus-deflection curve used for compliance calculation is from the

specified minimum load level to a value corresponding to 95 percent of the maximum load. The

typical minimum load level used for compliance calculation is the mean load (load signal midpoint).

Note, however, that the specific value of the minimum load level used for a given material, specimen

geometry, and load ratio must be larger than crack closure loads. The reason for excluding the

upper 5 percent of the load for calculation is that the clip gage tends to vibrate, and noise in the

clip-gage signal increases at the maximum load during the high-frequency test.

The increment of crack length (the difference between the current measured crack length and the

last recorded crack length) is checked against specified values, which are within the recommended

values of ASTM E647-83. A value of 0.5 mm is typically specified for a 25 .4-mm-thick standard

compact-type specimen. If the increment of crack growth is equal to or greater than the specified

value, the computer calculates N, AK, and da/dN; the digital plotter plots the data points (a,N) and

(da/dN,AK) on the a-versus-N and da/dN-versus-AK graphs, such as those shown in figure 1; the line

printer prints the value of calculated compliance, the linear least-squares correlation coefficient,

a, N, da/dN, and AK results. All the resulting data are stored on floppy disks for posttest

analyses.

During the test, the point-to-point data reduction technique is used to calculate AK and da/dN

from a and N. Usually the results are consistent with minimum scatter, such as those shown in

figure 1. If the results of AK versus da/dN scatter, the seven-point incremental polynomial method

is used to smooth the results after the test is completed.

The computer programs for posttest analyses include the following capabilities:

1 . Reducing a-versus-N data to AK-versus-da/dN by seven-point incremental polynomial method

2. Converting units

3. Plotting data in desired units

4. Plotting data in desired coordinate ranges

5 Plotting data for several different specimens on one graph (for comparison)

6. Calculating the material constants C and n in the Paris equation [14], da/dN = C(AK)", and

drawing the regression line through the data

All computer programs, including the data acquisition routines, were written in the Fortran IV language.

Near-Threshold FCGR Test

The computer programs used in the constant-load-amplitude FCGR test, with some modifications,

can be used for near-threshold FCGR tests. The major difference in test procedures is that the load

10
levels in the near-threshold FCGR test decrease according to the initial aK value (in the K- decreasing test technique). The load levels are calculated from the following equations [9]:

/ aK = aKq expEC (a - a^)] (1)

aP = BW ^^ AK/f,(a/W) for compact-type specimens (2)

aP = B AK/f2(a/W) for center-cracked-tension specimens (3)

Pmax = ^P/(l - R)= Pmin = Pmax « (^)

where P„_,^ = maximum load

'^min ~ minimum load

R = P / P min max

B = specimen thickness

W = specimen width

a = the current crack length

a = the crack length at the beginning of the test

fT(a/W) = [2 + (a/W)][0.886 + 4.64(a/W) - 13.32(a/W)^ + I4.72(a/W)3 - 5.6(a/W)^]/Cl - (a/W)''-^]

f2(a/W) = [(ira/W^) secdra/W) ]°-5

aK = current crack-tip stress-intensity factor range

aKq = crack-tip stress-intensity factor range at the beginning of the test

C = negative constant

A typical value of C is -0.08 mm" , which gives satisfactory results with no apparent anomalous

crack growth for AISl 300-series stainless steels.

The flow chart describing the automated near-threshold FCGR test is summarized in figure 4. After each crack length measurement, the crack length is compared with the last stored crack length

to ensure that a specified measurable small amount of crack growth has occurred. If this is not

done, then some unnecessary load level adjustments will take place because of scatter in the crack

length measurement. After the crack length has increased a certain amount (for example, 0.13 mm),

the new aK is calculated from equation (1) and the new crack length is stored. The load levels are then adjusted using equations (2), (3), and (4).

In high-frequency fatigue testing, which is desirable in the near-threshold FCGR test,

hydraulic lag might be a problem. As a result the specimen is not actually subjected to the load

range commanded by the computer (or waveform generator). The problem is usually corrected by using

proper signal conditioners and gain settings. However, overprogramming is sometimes necessary to overcome the persistent hydraulic lag. During the overprogramming process, which is done by a trial-and-error method, the computer monitors the values of P^av ^^^ Pmin through the programmable [lid A 111 X 11 digital oscilloscope and makes necessary changes to achieve the desired values of P„^^ and P„i„. lUaX UlX ll

If the measured crack-growth increment, which is the difference between the current measured

crack length and the last recorded crack length, is equal to or greater than specified values, which

11
Input Parameters

I

Calculate Load Levels

I

Cycling

#

"

Tinne to Measure No

Crack Length?

No

1 Yes

Crack Growth Incrennent Large Enough for New AK?

No

1 Yes

Calculate New AK

'

Adjust Load Levels

for New AK

1

'

Crack Growth Incrennent Large Enough for Storing Data etc.?

1 Yes

Calculate N, AK, da/dN

' '

Print, Plot, Store Results

Figure H. Summary flow chart of the automated neai—threshold FCGR test.

12
are within the recommended values [8], the values of N, AK, and da/dN are calculated and the results

are printed, plotted, and stored. A value of 0.5 mm is typically specified for a 25. '^-mm- thick

standard compact-type specimen. The previously mentioned computer programs for posttest analyses

are also applicable for analyzing the data obtained in the near-threshold FCGR test.

Variable-Load-Amplitude FCGR Test

The automated FCGR test system is also used in the variable-load-amplitude FCGR test. The

procedures used in this application are similar to those described in the previous two sections.

The computer reads the prerecorded load-time history from the floppy disks and controls the

hydraulic machine through the programmable waveform generator. At a preselected time interval, the

crack length is measured. A typical interval is 30 min for an average test frequency of 0.1 Hz. The desired output in the variable-load-amplitude test is time versus crack length. The results are printed, plotted, and stored for posttest analyses.

The present system has two limitations in the application of variable-load-amplitude FCGR

testing. One is that the waveform generator needs about 0.1 s for changing one commamd to another,

and this limits the average frequency to about 1 Hz. The needed 0.1 s for changing commands results

in a holding situation of about 0.1 s at the peak loads when higher test frequencies are used. In a corrosive or a high temperature environment, in which the hold time at peak load is important, this

might introduce anomalous fatigue crack growth. The other is limited storage capacity of the floppy

disks, which can only store a certain amount of load-time pairs. The present system uses soft

disks, which can store about 18,000 load-time pairs. If longer load-time histories are desired,

other means of storage, such as hard disks, must be used.

Summary

An automated FCGR test system has been developed that can be used for tests of constant-loadamplitude FCGR above 10~° m/cycle (ASTM E647-83), near-threshold FCGR, and variable-load-amplitude

FCGR. The test system saves considerable time in data acquisition and data reduction. The test

procedure is relatively easy to follow and enables technicians to produce data with less scatter

(with respect to the non-computer-aided technique), because higher precision in crack length

measurement and better control in data point spacing are obtained, while manual data interpretation

and data fitting are eliminated.

Acknowledgments

Mr. J. C. Moulder of NBS is acknowledged for helpful discussions on the interface between the

computer and the instruments. The work was supported by the Department of Interior, Minerals

Management Service, and the Department of Energy, Office of Fusion Energy.

References

[1] Fatigue Thresholds; Fundamentals and Engineering Applications , J. Backlund, A. F. Blour, and

C. J. Beevers, eds .

, Engineering Materials Advisory Services, Chameleon Press, London (1982).

13
[2] R. J. Bucci, "Development of a Proposed ASTM Standard Test Method for Near-Threshold Fatigue

Crack Growth Rate Measurement," in: Fatigue Crack Growth Measurement and Data Analysis , ASTM

STP 738, S. J. Hudak, Jr. and R. J. Bucci, eds ., American Society for Testing and Materials,

Philadelphia (1981), pp. 5-28.

[3] T. W. Crooker, F. D. Bogar , and G. R. Yoder, "Standard Method of Test for Constant-LoadAmplitude Fatigue Crack Growth Rates in Marine Environments," NRL Memorandum Report 459^, Naval

Research Laboratory, Washington, DC (August 1981).

[4] R. P. Wei, W. Wei, and G. A. Miller, "Effect of Measurement Precision and Data-Processing

Procedure on Variability in Fatigue Crack Growth-Rate Data," Journal of Testing and Evaluation , Vol. 7, No. 2 (March 1979), pp. 90-95.

[5] The Measurement of Crack Length and Shape During Fracture and Fatigue , C. J. Beevers, ed..

Engineering Materials Advisory Services, Chameleon Press, London (1980).

[6] Advances in Crack Length Measurement , C. J. Beevers, ed.. Engineering Materials Advisory

Services, Chameleon Press, London (1981).

[7] Y. W. Cheng, "A Computer-Interactive Fatigue Crack Growth Rate Test Procedure," in: Materials

Studies for Magnetic Fusion Energy Applications at Low Temperatures —VI , R. P. Reed and

N. J. Simon, eds., NBSIR 83-I690, National Bureau of Standards, Boulder, Colorado (1983), pp.

iJ1-51.

[8] J. J. Ruschau, "Fatigue Crack Growth Rate Data Acquisition System for Linear and Nonlinear

Fracture Mechanics Applications," Journal of Testing and Evaluation , Vol. 9, No. 6 (Nov. 1981),

pp. 317-323.

[9] "Proposed ASTM Test Method for Measurement of Fatigue Crack Growth Rates," in: Fatigue Crack

Growth Measurement and Data Analysis , ASTM STP 738, S. J. Hudak, Jr. and R. J. Bucci, eds.,

American Society for Testing and Materials, Philadelphia (1981), pp. 340-356.

[10] W. Elber, "The Significance of Fatigue Crack Closure," in: Damage Tolerance in Aircraft

Structures , ASTM STP 486, M. S. Rosenfeld, ed., American Society for Testing and Materials,

Philadelphia (1971), pp. 230-242.

[11] S. J. Hudak, Jr., A. Saxena, R. J. Bucci, and R. C. Malcolm, "Development of Standards of

Testing and Analyzing Fatigue Crack Growth Rate Data," AFML-TR-78-40, Air Force Materials

Laboratory, Wright-Patterson Air Force Base, Ohio (May 1978).

[12] T. Nicholas, N. E. Ashbaugh, and T. Weerasooriya, "On the Use of Compliance for Determining

Crack Length in the Inelastic Range," to be published in ASTM STP 833.

[13] R. L. Tobler and W. C. Carpenter, "A Numerical and Experimental Verification of Compliance

Functions for Compact Specimens," to be published in Engineering Fracture Mechanics .

[14] P. C. Paris and F. Erdogan, "A Critical Analysis of Crack Propagation Laws," Journal of Basic

Engineering, Trans. ASME , series D, Vol. 85, No. 3 (1963), pp. 528-534.

14
THE FATIGUE CRACK GROWTH OF A SHIP STEEL

IN SALTWATER UNDER SPECTRUM LOADING""

Yi-Wen Cheng

Fracture and Deformation Division

National Bureau of Standards

Boulder, Colorado

Abstract

Fatigue crack growth under spectrum loading intended to simulate sea loading of offshore

structures in the North Sea was studied using the fracture mechanics approach. A digital simulation

technique was used to generate samples of load-time histories from a power spectrum characteristic

of the North Sea environment. In the constant-load-amplitude tests, the effects of specimen

orientation and stress ratio on fatigue crack growth rates were negligible in the range 2 x 10"^ to 10~^ mm/cycle. Fatigue crack growth rates in a 3-5 percent NaCl solution were two to five times

higher than those observed in air in the stress intensity factor range 25 to 60 MPa/m. The average

fatigue crack growth rates under spectrum loading and under constant-amplitude loading were in

excellent agreement when fatigue crack growth rate was plotted as a function of the appropriately

defined equivalent-stress-intensity range. This procedure is equivalent to applying Miner's

summation rule in fatigue life calculations.

Key words: corrosion fatigue; fatigue crack growth; fracture mechanics; seawater environment;

spectrum loading; structural steel.

"" Published in the International Journal of Fatigue, Vol. 7 (April 1985), pp. 95-100.

15
Introduction

In recent years the petroleum Industry has built offshore drilling and production platforms in

deeper waters and more hostile climates. As the offshore platforms encounter more severe weather

and rougher sea-state conditions, fatigue becomes a more important factor in consideration of

structural integrity. In treating the fatigue problem, it is usual to separate the fatigue life

into two separate stages: (1) crack initiation and (2) crack growth. For welded structures, such as offshore platforms, crack initiation, during which microcracks form, grow, and coalesce to become a macrocrack, is less important than crack growth because fabrication imperfections are always

present. Most of the fatigue life is spent in the crack growth stage.

Analysis of fatigue crack growth under spectrum loading, which is usually irregular in nature,

is complicated because of load-sequence interaction effects. A cycle-by-cycle approach, taking into

account overload effects, has been used in the aerospace industry [1,2] . Other empirical

approaches, such as root mean square (RMS) [3] or root mean cube (RMC) [4], have also been

successfully used to correlate experimental results of spectrum loading of bridges with those of

constant-amplitude loading. Use of the latter approaches is empirical and implementation of the

former is time-consuming. A more efficient approach has been proposed [5,6], which will be

discussed later. This paper describes the work carried out at the National Bureau of Standards over

the past two years on the investigation of fatigue crack growth in ABS grade EH36 steel under

simulated offshore platform service conditions.

Load Spectrum

Service loads acting on offshore structures are random in nature. The main source of cyclic

loading derives from wave action, which excites a vibration at approximately the wave frequency.

The magnitude of the vibration depends mainly on wave height and direction, size of the component

and its location in a structure. Besides those due to wave action, additional vibrations are induced from structural responses to the wave action. The magnitude and frequency of the structural resonance depend on local structural characteristics. Thus, the precise definition of load-time

history is extremely complex and would be expected to vary between locations on the same structure.

Because of complexity in and lack of information on the precise load-time history experienced

by offshore structures, no standard load-time history exists for purposes of analysis and experiment. Numerous load-time histories, including Rayleigh peak distribution [7,8], Gaussian peak

distribution [7-9], Gassner blocked program [10], and others [11], have been used to evaluate

fatigue performance of weldments. The load spectrum selected for the present investigation was realistic for offshore structures in the North Sea environment [12], as shown in figure 1. The

principal loads in this spectrum, those with a frequency of about 0.1 Hz, are due to wave action.

The higher frequency (about 0.35 Hz) loads are due to structural resonance.

*Numbers in brackets denote references listed at the end of each paper,

16
O 1.0

>-"

z

m

Q 0.6

<

0.4

O

LU

q. 0.2

CO

QC

LUo

Q.

0.1 0.2 0.3

FREQUENCY, Hz

0.4

Figure 1 . Characteristic power spectrum for offshore structures in the North

Sea.

Simulation of Load-Time Histories

For purpose of experiment, the power spectral density function, S(a)), is not sufficient;

load-time history, X(t), has to be used. In this investigation, the following expression [IS.T^]

was used to reconstruct X(t) from S(a)):

X(t) = Z [2G(aj[^)Aa),^]°-5 cos(a),^t + $^)

k=1

(1)

where GCu), as shown in figure 1, is the one-sided power spectral density function in terms of

frequency, w [G(a)) = 2S(to) for co > 0] . $|^ is a random phase angle uniformly distributed between

and 2it; and m^ is the midpoint of Aa)|^ . The number of harmonic functions, J, is arbitrary; in this

investigation it was taken to be 50. Frequency is defined over the interval [0, w^j] with partitions

of length such that

J

= E Acji^ (2)

An X(t) with an undesired short period occurs if the minimum common divider for all the Ma^ is

large. This problem is avoided by using random intervals for aw|^ . In this investigation, AtO|^ was taken from a normal distribution with a mean equal to the average of A(jO|^ and a standard deviation

equal to one-tenth the average of acjw .

17
A computer program written in Fortran IV has been developed to simulate X(t) from equation (1).

Newton's method was then used to locate peaks and troughs with respect to time in the simulated

load-time history.

Two load spectra were used in this study. One contains only the wave-loading portion of the

power spectral density function with frequencies up to 0.2 Hz (case I), as shown in figure 1. The

other reproduces the whole curve (case II). Typical simulated load-time histories, X(t), from the

power spectral density function are shown in figure 2.

Values of the irregularity factor (number of mean crossings/number of peaks plus troughs)

calculated from the power spectra are 0.90 and 0.69 for case I and case II, respectively; they are 0.90 and 0.68, as determined from the simulated load-time histories. The excellent agreement

between the values obtained from the power spectra and the simulated load-time histories indicates

that use of equation (1) is satisfactory. Values of the clipping ratio are 3-84 and 3-91 for case I and case II, respectively. Clipping ratio is defined as the ratio of the maximum load amplitude,

which is the difference between the maximum peak and the mean load, to the root-mean-square value of

load amplitude.

Experimental Procedures

Test Material and Specimens

The test material was a 25 .4-mm-thick plate of ABS grade EH36 steel, a 350-MPa-yield-strength

C-Mn steel. The chemical composition is given in table 1. The steel was in the normalized

condition and had particularly uniform properties due to sulfide shape control.

Fatigue crack growth rate (FCGR) tests under constant-amplitude loading and spectrum loading were conducted using standard (25 .4-mm-thick) and modified [15] compact specimens. The modified

compact specimen was a lengthened and side-grooved (with a net thickness of 3.18 mm) version of the

standard compact specimen. The deep side grooves determine the plane of crack growth and provide a strip of material that undergoes large cyclic plasticity during fatigue. Specimens were in LT and

TL orientations.

Test Apparatus and Environment

Fatigue crack growth rate tests were conducted with a fully automated test system, which was described in a previous paper [16]. Briefly, the fully automated test system consists of a closed-loop, servo-controlled, hydraulic mechanical test machine, a programmable digital oscilloscope serving as an analog-to-digital converter, a programmable arbitrary waveform generator, and a minicomputer.

Tests were performed in laboratory air and in 3.5 percent NaCl solution with a free corroding

*This should not be confused with the root-mean-square value of load range which was used in

reference 3.

18
5 min

Figure 2. Samples of load-time histories: (a) case I, (b) case II

Table 1. Chemical composition of ABS grade EH36 steel,

Mn Si Cu Ni Cr Mo Fe

0.12 1.39 0.015 0.006 0.380 0.05 0.03 0.05 0.007 bal

,

condition (no cathodic protection). Crack lengths were measured by the compliance technique. The

crack-length measurement technique was accurate at least to 0.1 mm. In the saltwater tests, the

clip-gage used for displacement measurements was mounted on a scissors-like extension to avoid

immersion in the saltwater. The environmental chamber was a 19-J!.-capacity plastic container in

which the saltwater was continuously circulated at a rate of 26 £/min through a diatomaceous-earth

filter. The NaCl concentration, temperature, and pH value of the saltwater were monitored

periodically.

Loading Conditions

In the constant-load-amplitude tests, the stress ratio R (i.e., the ratio of minimum to maximum

stress) was kept constant at 0.1 or 0.5. Tests in air were conducted at 10 Hz and tests in 3.5

percent NaCl solution were conducted at 0.1 Hz. A sinusoidal load-time history was used.

In the spectrum-loading tests, the simulated load-time histories were recorded on floppy disks,

which were read by a minicomputer. The minicomputer then sent the signals to the hydraulic

mechanical test machine through a programmable arbitrary waveform generator. The loads were periodically monitored with an oscilloscope to ensure that the input values to the hydraulic

mechanical test machine and the output values from the load cell agreed. No modifications, such as truncation, were made to the simulated load-time histories with the exception of the levels of mean

loads. The mean loads were increased so that the minimum loads were slightly-above zero because the

19
apparatus was limited to tension-tension loading. The stress ratio, therefore, varied from about

(usually for large load ranges) to about 1 (usually for very small load ranges).

Because of the limited capacity of the floppy disk, the total recorded lengths of load-time

histories were 18.0 h for case I and 9.3 h for case II. The recorded lengths corresponded to return

periods of 15,773 and 11,890 mean-load crossings for case I and case II, respectively. The wave shape was triangular. It has been shown [17] that there are no differences in FCGR between tests

conducted with sinusoidal and triangular waveforms. Both tests in air and in saltwater were conducted at ambient temperature.

Experimental Results and Discussion

Constant-Load-Amplitude Tests

Fatigue crack growth rates were calculated using the linear-elastic fracture mechanics

approach; the experimental results are shown in figures 3 through 6. As shown in figure 3, specimen

orientation, TL versus LT , had little influence on FCGR in air and in saltwater. The FCGRs in air

and in saltwater are compared in figures 4 and 5. For stress intensity factor range, aK, between 30

and MO MPa/m, the FCGRs in saltwater were up to five times higher than those in air. A summary of

all results in figure 6 indicated that stress ratio had little influence on FCGR in air. Below

4 X 10~^ mm/cycle the FCGRs in air and in saltwater were about the same.

The minimal influence of stress ratio and specimen orientation observed is consistent with that

of other investigators [18] in the FCGR range 2 x 10~^ to 10~^ mm/cycle. The effects of stress

ratio and specimen orientation are expected to be more pronounced at higher and lower FCGRs [18].

Note that in each of the FCGR curves the high values were obtained with the modified compact

specimen. These data follow the same trend line as the data obtained with the standard compact

specimen. Thus, it appears that the linear-elastic fracture mechanics approach can be applied to fatigue crack growth in conditions of contained large cyclic plasticity.

Spectrum-Loading Tests

Fatigue crack growth rates under spectrum loading were analyzed using the equivalent-stressrange approach [5,6], which is described in the following. For simplicity the Paris equation,

da/dN = C(aK)", is used for discussion. Here, da/dN is crack growth increment per load cycle, aK is

stress intensity factor range, and C and n are constants. aK is defined as

AK = h (TTa)0-5 Y (3)

where h = stress range

a = crack length

Y = geometry factor

If da/dN € crack length, a, and there are no load-sequence interaction effects, then

20
10-2 10

AK, ksiy^

FT

10

-3

A D TL Orientation

_ X A o LT Orientation

In 3.5%

NaCI Soln

o

o

E

E

CO D

10-V

10

-5

50

"~i—

r

ABS EH36 Steel

R = 0.1

10

o

o

>0< ^

^In Air

#

S

n

a

a

T~l

I I

I I l__L

AK, MPayifn

50

10

-4

10

o

>

c 5 —

CO

10

-6

100

Figure 3- Fatigue crack growth rates in EH36 steel: effect of specimen

orientation. Different symbols represent results obtained from

different specimens.

21
10-2 10

AK, ksi^yirf

50

lO-^k

O

o

CO 10~^h

10

-5

-1 •

1 1 1 1 1

1 1 "" ABS EH36 Steel

- R = 0.1

-

- 3.5% *#

NaCI Soln ^^

-

- *% A

^

-

- xx

-

D XX ^

-

- Xxo - / - ^

- — » - - 0°

~ o

- - o

_ o

-

o

1 1 1 10

AK, MPayfrT

50

-10-4

10

o

>

o

c

-5 —

CO o

10

-6

100

Figure ^. Fatigue crack growth rates in EH36 steel: saltwater versus air at stress ratio equal to 0.1 . Different symbols represent results

obtained from different specimens.

22
AK, ksi^yin

10

10

o

o—

.

E

E

CO D

10

r2 10 50

< -1 1 1 1 1

1 1 1 — ABS EH36 Steel

-

- R = 0.5

-

-3 —

D

3.5% D cP

NaCI Soln :&

-

— -

1-4 -

A jf

^ oSf - o°Alr

A O

O

aV'

o

cP

-

- c?^ ^

1-6 1 1

1 1 1 1 1

10'^-

-10-4

O

-10,-5

CO D

-10-6

10

AK, MPaym

50 100

Figure 5. Fatigue crack growth rates in EH36 steel: saltwater versus air at stress ratio equal to 0.5. Different symbols represent results

obtained from different specimens.

23
O

>

o

AK, ksiv^

10

10"^-

10

r2 10 50 1 -1 1 ABS EH36 Steel

1 1 1 1 1

1 1

- AB R = 0.5 - — XOAD R=0.1

1-3

-

In 3.5%

NaCI Soln ^r /

-

-

AO

-

A j

^ In Air

- ^ ^° - ^ ^^ -

1-4 - -

" ^ -

r5 1 1 1 III 1

10"^-

10

AK, MPayfrT

50

-10-4

-10

>

o

c

-5 —

CO D

-10-6

100

Figure 6. Fatigue crack growth rates in EH36 steel: summary. Different

symbols represent results obtained from different specimens.

24
AaAa

1 = C hi" [(iTa)0-5 Y]"

2

= C h/ [(TTa)0-5 Y]" (5)

Aajj = C h,^" [(Tra)0-5 Y]" (6)

Summing equations (4) through (6) gives

(Aa^ + Aa2 + ... + Aajj) = C (h^" + h2" + ... + hjj")[ (Tra)°-5 y]" (7)

The left-hand side of equation (7) is the increment of crack growth in N successive cycles; the

average FCGR per cycle is then

da/dN = C [(h^" + h2" + ... + h^")/N][ (^a)°-5 y]"

= C [(h^)^/" (TTa)0-5 Y]"

= C [hgq (Tra)°-5 Y]"

(8)

(9:

Here N should be large in order for the equivalent-stress range, h , to be representative of a load

spectrum. The definitions of stress range and cycle used in this investigation are given in

figure 7. The value of n in 3.5 percent NaCl solution test is 5.5, which is derived from the

results of constant-load-amplitude test in the aK range of interest.

The results for FCGR under spectrum loading in 3.5 percent NaCl solution are given in figures 8 and 9 for case 1 and case II, respectively. Excellent agreement between spectrum and sinusoidal

loading is observed. This suggests that load-sequence* interaction effects are effectively

Mean

Figure 7. Definitions of stress range and cycle.

25
10-2

O

o

£

E

CO

10 -3

10"

10 -5

AK, ksifln

3 10 30

I I I I I l|

O Constant Amplitude Data

X Spectrum Loading Data

(power spectrum case I)

O

O

J \ I J_ 10

AK, MPafm

90

-10

I I I I 1 1

1

-4

10

-5

10,-6

100

o

>»o

(0

Figure 8. Fatigue crack growth rates in EH36 steel in saltwater: con- stant-amplitude loading versus spectrum loading (case I).

26
10r2

Q) 10"^

O

>»o

E

E

(0 o 10

-4

10

.-5

AK, ksiVTn

3 10 30

"1—I I I

I I l| 1 90

1—I

I I I I I

X Constant Amplitude data • O Spectrum Loading Data

(power spectrum case II)

i_LL J I I I I I I

10

,-4

O

o

10-^ z

(0 a

10

10

AK, MPa>^

100

Figure 9. Fatigue crack growth rates in EH36 steel in saltwater: constantamplitude loading versus spectrum loading (case II).

27
negligible. The lack of observed load-sequence interaction effects is probably due to low clipping

ratio of 3.84 (case I) and 3-91 (case II). The results also imply that under spectrum loading at a given da/dN, value of aK is smaller if RMS (n = 2) or RMC (n = 3) is used because h decreases with

decreasing n. This shifts the spectrum-loading results to the left of those of constant-amplitude

loading (figures 8 and 9), resulting in a higher FCGR in spectrum loading than in constant-amplitude

loading at a given value of aK. Conversely, a lower FCGR will result if either RMS or RMC is used

to predict FCGR in a region where n is larger than 3. such as in the present investiga1;ion.

Miner's rule [19] states that a component (or specimen) will fail if

I (fi/F^f) i 1 (10)

where f^ is the number of fatigue cycles applied at stress range aS^^ and Fj_^ is the number of

fatigue cycles to failure at stress range aS^ . This rule implies that there are no load-sequence

interaction effects. Miner's rule, as originally stated, applied to fatigue failure rather than

fatigue crack growth. Terms such as "Miner's rule of fatigue crack growth" are often used to mean

fatigue crack growth with no load-sequence interaction effects. Such statements represent a generalization of the original Miner's rule. The data of this study, along with others [5], support

such a generalization for a clipping ratio less than k and constant mean stress, which might be

stated as follows: Load-sequence interactions are small, or they tend to cancel, such that overall

effect on fatigue life is small. For such a rule to be applicable to random or quasi-random

load-time histories, a definition of a cycle is needed. In this study, the load amplitude of one cycle has been defined as the maximum load difference among three successive mean crossings

(figure 7).

The value of h can be obtained in a closed-form expression from the power spectrum if the

fcjq loading is a narrow-band random process [5,20]. However, no closed-form solutions are available for

wide-band random processes.

Conclusions

The following conclusions were drawn from this investigation:

1

.

The digital simulation technique is adequate to generate samples of load-time histories from a given power spectrum.

2. In constant-load-amplitude tests, the influence of specimen orientation and stress ratio on fatigue crack growth rate were negligible in the fatigue crack growth rate range 2 x 10 -^ to 10~^ mm/cycle. Fatigue crack growth rates in a 3.5 percent NaCl solution were two to five

times higher than those observed in air in the stress intensity factor range 25 to 60 MPa/m.

3. The average fatigue crack growth rates under spectrum loading and under constant-amplitude

loading were in excellent agreement when fatigue crack growth rate was plotted as a function of

the appropriately defined equivalent-stress-intensity range. This procedure is equivalent to applying Miner's summation rule in fatigue life calculations.

28
Acknowledgments

Helpful discussions with Drs. H. I. McHenry and D. T. Read, and Professors S. Berge and

P. N. Li are appreciated. This work was supported by the Department of Interior, Minerals

Management Service.

References

[I] 0. E. Wheeler, "Spectrum Loading and Crack Growth," Journal of Basic Engineering , Trans. ASME

,

Vol. 94 (March 1972), pp. I8I-I86.

[2] J. Willenborg, R. M. Engle, and H. A. Wood, "A Crack Growth Retardation Model Using an Effective Stress Concept," AFFDL-TM-7I-FBR, Air Force Flight Dynamics Laboratory, Dayton, Ohio

(January 1971 )

[33 J. M. Barsom, "Fatigue Crack Growth under Variable-Amplitude Loading in Various Bridge Steels,"

in: Fatigue Crack Growth under Spectrum Loads , ASTM STP 595, American Society for Testing and

Materials, Philadelphia (1976), pp. 217-235.

[i)] P. Albrecht and K. Yamada , "Simulation of Service Fatigue Loads for Short-Span Highway

Bridges," in: Service Fatigue Loads Monitoring, Simulation, and Analysis , ASTM STP 671,

American Society for Testing and Materials, Philadelphia (1979), pp. 255-277.

[5] W. D. Dover, S. J. Holbrook, and R. D. Hibberd, "Fatigue Life Estimates for Tubular Welded T Joints Using Fracture Mechanics," in: Proceedings of European Offshore Steels Research Seminar , The Welding Institute, Cambridge, UK (November 27-29, 1978), pp. V/PD-1 - V/PD-11.

[6] F. A. McKee and J. W. Hancock, "Fatigue Crack Growth and Failure in Spectrum Loading," in:

Proceedings of European Offshore Steels Research Seminar , The Welding Institute, Cambridge, UK

(November 27-29, 1978), pp. V/PC-1 - V/PC-10.

[7] L. P. Pook, "Proposed Standard Load Histories for Fatigue Testing Relevant to Offshore

Structures," NEL Report No. 624, National Engineering Laboratory, Glasgow, UK (October 1976).

[8] M. H. J. M. Zwaans, P. A. M. Jonkers, and J. L. Overbeeke, "Random Load Tests on Plate

Specimens," Eindhoven University of Technology, The Netherlands (December 1980).

[9] P. J. Haagensen and V. Dagestad, "Corrosion Fatigue Crack Propagation in Structural Steel Under

Stationary Random Loading," SINTEF Report No. I8 A 78017, The Foundation of Scientific and

Industrial Research at the Norwegian Institute of Technology, Trondhein, Norway (October 2,

1978).

[10] H. P. Lieurade, J. P. Gerald, and C. J. Putot, "Fatigue Life Prediction of Tubular Joints," in: Proceedings of the Offshore Technology Conference , OTC Paper No. 3699, Houston, Texas (May

1980).

[II] R. M. Olivier, M. Greif, W. Oberparleiter , and W. Schutz, "Corrosion Fatigue Behavior of

Offshore Steel Structures under Variable Amplitude Loading," in: Proceedings of International

Conference on Steel in Marine Structures , Paper 7.1, Paris, France (October 5-8, 1981).

[12] R. M. Kenley, "Measurement of Fatigue Performance of Forties Bravo," in: Proceedings of the

Offshore Technology Conference , OTC Paper No, 4402, Houston, Texas (May 1982).

[13] J. N. Yang, "Simulation of Random Envelop Processes," Journal of Sound and Vibration , Vol. 21, No. 1 (1972), pp. 73-85.

[14] P. H. Wirsching and A. M. Shehata, "Fatigue under Wide Band Random Stresses Using the Rainflow

Method," Journal of Engineering Materials and Technology , Trans. ASME, Vol. 99, No. 3 (July

1977), pp. 205-211

[15] H. I. McHenry and G. R. Irwin, "A Plastic-Strip Specimen for Fatigue Crack Propagation Studies

in Low Yield Strength Alloys," Journal of Materials , JMLSA, Vol. 7, No. 4 (December 1972),

pp. 455-459.

29
[16] Y. W. Cheng and D. T. Read, "An Automated Fatigue Crack Growth Rate Test System," to be

published In ASTM STP 377.

[17] J. M. Barsom, "Effect of Cyclic Stress Form on Corrosion Fatigue Crack Propagation below Kj

in a High Yield Strength Steel," in: Corrosion Fatigue: Chemistry, Mechanics, and Micros - tructure . National Association for Corrosion Engineers, NACE-2, Houston, Texas (1972),

pp. 42iJ-435.

[18] R. 0. Ritchie, "Influence of Microstructure on Near-Threshold Fatigue-Crack Propagation in

Ultra-high Strength Steel," Metal Science , Vol. II (1977), pp. 368-381.

[19] M. A. Miner, "Cumulative Damage In Fatigue," Journal of Applied Mechanics , Trans. ASME , Vol. 12 (September 1945), pp. A159-A164.

[20] J. N. Yang, "Statistics of Random Loading Relevant to Fatigue," Journal of the Engineering

Mechnaics Division , Proceedings of the American Society of Civil Engineers, Vol. 100, No. EM3

(June 1974), pp. 469-475.

30
ESTIMATION OF IRREGULARITY FACTOR FROM A POWER SPECTRUM

Pei-Ning Li*

Yi-Wen Cheng

Fracture and Deformation Division

National Bureau of Standards

Boulder, Colorado 80303

Abstract

This paper presents several simplified methods of evaluating the irregularity factor of a power

spectrum. The irregularity factor can be computed either from Integration of the power spectrum or from the characteristic bandwidth and the center frequency of the power spectrum. The characteristic bandwidths and the center frequencies of power spectra with irregular shapes are defined in this

paper. Estimated errors associated with the simplified methods in the cases of practical interest

are given.

Key words: bandwidth; irregularity factor; power spectrum; random loading fatigue.

Guest worker, on leave from East-China Institute of Chemical Technology, Shanghai, China.

31
Introduction

In many structural applications, such as offshore platforms in the North Sea environment,

fatigue under random loading is a major problem. Load-time histories under random loading are usu- ally difficult to predict and can only be treated in a statistical manner. If the random loading is

a stationary Gaussian process, as is commonly assumed, then there exists a power spectrum, G(f),

which possesses all the statistical properties of the original load-time history [1] . Therefor(

the power spectrum is conveniently used to represent the random load-time history.

Several important parameters in the random-loading fatigue analysis can be derived from the

power spectrum. These parameters include root-mean-square (RMS) value of the load amplitude, average rises and falls, and the irregularity factor, a [2,3]- The RMS value equals the square root of

the area under the power spectrum-versus-frequency curve. The average rises and falls, which are related to the RMS value and the irregularity factor of the power spectrum, have been analytically

and numerically studied [^]. The irregularity factor has been used as a parameter to normalize the

fatigue damage caused by narrow-band and broad-band loadings [5]. The irregularity factor of a power spectrum is usually computed from the integration of the power spectrum, which can become

tedious and time-consuming if the shape of G(f) is irregular. This paper presents simplified techniques for estimating a from a power spectrum.

Irregularity Factor

The irregularity factor, a, is defined as the ratio of the number of positive-slope zero crossings, Nq, to the number of peaks per unit time in a load-time history, F^:

No

Ct = (1)

The exact value of N and F^ cari be evaluated from G(f) as follows:

-(x)'

(^)-

(2a)

Fq = ( — I ^-^ (2b)

where M , Mo, and Mj^ are the zeroth, second, and fourth moments of G(f) about the origin (zero frequency) and are defined as:

Mq = /„ G(f) df (3a)

Numbers in brackets denote references listed at the end of each paper.

32
M2 = /r f^ G(f) df (3b)

Mij = /, f^ G(f) df (3c)

where f is frequency. Thus,

a = (Mq Mi,) 0.5

(4)

The Irregularity factor, a, not only describes the irregularity of the random load-time history

but also is a measure of the bandwidth of G(f). As a approaches unity, the distribution of the

loading peaks approximates to the Rayleigh distribution [1], and the shape of G(f) is sharply peaked

at the center frequency or far away from the origin. This is called narrow-band power spectrum. A single-frequency sine-wave loading can be described as a Dirac-Delta function power spectrum; it has

a = 1 . The value of a decreases with increasing width of the power spectrum.

Evaluation of Irregularity Factor from Power Spectra

Direct Integration of Power Spectral Density Function

The value of a can be evaluated from equation m) by integrating equations (3a), (3b), and

(3c). The integrations can become tedious and time-consuming if the shape of G(f) is irregular.

One simplified way of evaluating a is to break G(f) into n simpler geometries, such as those shown

in figure 1 (n = 5), and then to evaluate the moments according to the following equation :

o

>-"

CO

z

LJJ Q

_l

<

COI-o

UJQ.

CO

CC

LUo

jy Y^

v

\

/

t \\

N / \>^

*1 h fa ^4 FREQUENCY. Hz

Figure 1 . Power spectrum: dashed line represents the original spectrum, and

solid lines are a simplified diagram.

33
^J

= (j+1)(j+2) i=l

I (Gi - G,,,)

f.(j + 2)_ (j + 2) M+ 1 M

fi.l - h

0, 2, 4 (5)

where G^^ ' s and f^'s are power spectral densities and frequencies respectively, as shown in figure 1. (The derivation of equation (5) is given in the appendix.) For example, from equation (5) and the

smplified diagram shown in figure 2 (n = 9), the value of a was calculated to be 0.699. Using integration technique of equation (4), the value of a was 0.697. The error of using equation (5) was

0.14 percent.

Estimation from Characteristic Width and Center Frequency of the Power Spectrum

The irregularity factor can also be estimated from the characteristic width and the center

frequency of a power spectrum. For the case of rectangular power spectra, a can be obtained from

the following expression [6].

5 (9 + 68^ + b'*)

9 (5 + lOB^ + b'*)

0.5

(6)

where B = W/2f , W is the width, and f^ is the center frequency of the rectangle (power spectrum).

Here, B is the geometric dimensionless bandwidth of the power spectrum. Different power spectra

with same values of B have the same values of a, regardless of their shapes and positions with re- spect to the origin (zero frequency)

both are determined in a straightforward manner,

The definitions of W and f are obvious in this case, and

For irregular spectra, the determination of W and

GC

lU

o

Q.

0.1 0.2 0.4

FREQUENCY, Hz

0.3

Figure 2. Example of computing the irregularity factor from a simplified

diagram of a power spectrum.

34
f can be difficult; it is discussed in the following sections.

From geometric analysis of rectangular power spectra, it -was noted that the values of B were bounded by and 1. Accordingly, a was within 1 and O.T'JS.

Determination of characteristic width . In the case of symmetric, single-peak power spectra,

the center frequency is at the center of the frequency range. However, the determination of the

characteristic width, W, is not obvious and the following empirical equation has been used to evaluate it:

/ Vec \

W = w I I

0.5

(7)

where W is the arithmetic average width of the power spectrum, A is the area of a rectangle

enveloping the power spectrum, and A^^ is the area of the power spectrum. For example, the value of

W is half of its base width for an isosceles triangular power spectrum and the ratio of A^g^/A_g is

Results for several symmetric, single-peak power spectra with various shapes, including isosceles triangles, rectangles, isosceles trapezoids and pagodas, are plotted in figure 3 in the form

of a versus B. All values of a in figure 3 and other figures in this paper were calculated from

equation (4), if not otherwise specified. The value of a obtained from equation (6) coincides with

those of the rectangles in figure 3- It can be seen that a-versus-B curves for four different

shapes of power spectra were in reasonable agreement, demonstrating that estimates of W from equation (7) combined with equation (6), provide good estimates of a. The errors in a at B = 1 were

less than ±5 percent.

o

lU

cc

cc 0.2 0.4 0.6 0.8 1.0

GEOMETRIC DIMENSIONLESS WIDTH, B

Figure 3- Irregularity factor-versus-geometric dimensionless width for four

different symmetric, single-peak power spectra.

35
Determination of center frequency . The value of a of an arbitrary triangular power spectrum

with a fixed base (i.e., with a constant characteristic width) varies with the location of the peak

within the width, because the center frequency varies. Several candidates for the operational definition of center frequency, f^, including the frequency at the peak of the power spectrum, the

frequency at the center of gravity of the power spectrum, the frequency at the middle of the frequency range, and the frequency at the middle of the half-height width, were studied. The frequency

at the middle of the half-height width of the power spectrum gave the least scatter in the a-versusB curves of several asymmetrical single-peak power spectra. The a-versus-B curves of right triangular power spectra with the right angle to the left or to the right are compared with those of rec- tangular power spectra in figure 4. The figure shows that at a given value of B, the values of all

the triangles studied are slightly smaller than those of rectangles. The largest error at B = 1 was about 10 percent. This means that use of the frequency at the middle of half-height width of a triangular power spectrum slightly overestimates the value of a.

Behavior of double-peak power spectra . The irregularity factor of a double-peak power spectrum

was estimated from the following equations:

1 + A F'

[(1 + A)(1 + A F^)]°-5

(8)

A2

^ = A"^1 (9)

F = c2

'c1 (10)

where A^ , Ag, fd . and fj^2 af'© the areas and center frequencies of each peak, respectively, with A-i and f . being closer to the zero frequency (see table 1). As shown in table 1, the error of using

o

I-o

<

u.

>-

q:

<

O

LJJ ccoc

0.2 0.4 0.6 0.8 1.0

GEOMETRIC DIMENSIONLESS WIDTH, B

Figure M. Irregularity factor-versus-geometric dimensionless width for four

different asymmetric, single-peak power spectra.

3 6
equation (8) is within ±5 percent for a wide range of two-peak power spectra. Table 1 contains

results ranging A from 0.1 to 10, F from 2.333 to 7, G2/G^ (see table 1 for definition of G^ ) from

0.1 to 10, and a from 0.56 to 0.96.

/ Figure 5 presents results from equation (8) in graphic form. It shows that the larger the

value of F and the smaller the value of A are, the smaller the irregularity factor is. From

equations (3) and (4), one would expect that the higher frequency peak dominates the determination

of a because of the second and the fourth power of frequency in M2 and H^. In the case of ^2 ^ ^i

»

the lower frequency peak can be neglected in the determination of a. The smaller the value of F is,

the closer the two peaks are, and vice versa. Figure 5 shows that, at a given value of A, a de- creases with increasing F.

The two peaks in a double-peak power spectrum of practical interest usually connect to each

other at their bases. In this case, the double-peak power spectrum was divided into two parts.

Their center frequencies and areas were estimated; then the value of a was estimated from

equation (8). For example, the power spectrum characteristic of the North Sea environment, as shown

in figure 6, was divided at 0.3 Hz, 0.25 Hz, or 0.2 Hz and represented by two triangles, aABC and

aDEF, aABC and aD'EF, or aABC" and aD"EF. All triangles had the same areas as the original curves

that they represented. The values of F and A obtained were 2.70 and 0.15T*. 2.71 and 0.2278, or 2.7^ and 0.3165, respectively. The estimated value of a using equation (8) was 0.6520, O.66I6, or 0.6785. The errors were -6.5, -5.08, and -2.65 percent, respectively.

Summary

Several simplified methods for evaluating the irregularity factor of a power spectrum have been

derived. The irregularity factor was computed either from integration of the power spectrum or from

the characteristic bandwidth and the center frequency of the power spectrum.

For idealized shapes of power spectra, such as rectangles and isosceles triangles, the characteristic bandwidth and the center frequency were readily obtained. For irregular shapes, the power

spectra were represented by simplified geometries from which the characteristic bandwidths and the

center frequencies were estimated.

For all the cases studied in this paper, the largest errors introduced by using the simplified

methods was about 10 percent, but the majority were within 5 percent. Therefore, use of the simplified methods is recommended where approximation is allowable and as a check on more exact methods.

Acknowledgments

This work was supported by the Department of Interior, Minerals Management Service.

37
1.0

o

I-o

<

LL 0.6

< 0.4 - 3

CD

LU

CO

CC 0.2

10

-

-£^a-L__J_-^^^_ ~-

-^--y

^ - ^^s^^y

- ^9./ -

-

1 1

1

1/A (for A< 1)

1 5

A(for A >1)

10

Figure 5. Irregularity factor as a function of the ratio of area and

frequency of each peak for double-peak power spectra.

o

Q.

0.1 0.2 0.4

FREQUENCY, Hz

0.3

Figure 6. Example of calculating the irregularity factor of a double-peak

power spectrum of practical interest by dividing the two connected

peaks into two separate triangles.

38
Table 1 . Estimated errors in irregularity factor of double-peak power spectra

determined by simplified method.

-^ 0^:02 fl = f2 ' h f4 F A Value of a | Shape of Spectra by Eq(i|) by Eq(8) $error

G(

G2

f) 1:1

1:2

12

12

12

12

2

3

3

k

2

3

3

J4

3

5

3

5

52.333

7

3

52.333

73

1

1

1

1

2

22

2

0.72il9

0.8138

0.7162

0.7762

0.8160

0.8733

0.8l6i<

0.8532

0.731*8 0.8232

0.7213

0.7808

0^325

0.88H0

0.82117 0.8592

1 .50 1.16

-0.71

0.60

2.02

1.23

1.02

0.70

_ t

f, fs fg U

'

2:1 12

12

2

3

3

3

5

52.333

7

3

0.5

0.5

0.5

0.5

0.6192

0.7559

0.6000

0.6939

0.6225

0.761*1 0.6007

0.6971

0.53

1.08

0.12

0.1*6

5:1' 12

12

2

3

3

3

5

52.333

7

3

0.2

0.2

0.2

0.2

0.M969

0.7161

0.i<559

0.6150

0.1*880

0.721*1*

0.1*1*91* 0.6163

-1.79

1.16

-1.1*3 0.21

1:5 • 12

: 1

• 2

2

3

• 3

3

H

5

52.333

7

3

55

5

5

0.8989

0.9312

0.9057

0.92^49

0.9200

0.91*35 0.9165

0.9320

2.35

1.32

1.19

0.77

10:1 : 1

- 2

: 1

• 2

: 2

3

: 3

3

5

52.333

7

3

0.1 0.1

0.1

0.1

0.1*381 0.7280

0.3756

0.5996

0.H188

0.7396

0.3623

0.6298

-1*.1*7 1.59

-3.51* 5.00

1:10 . 12

! 1

• 2

: 2

3

: 3

J4

3

5

52.333

7

3

10

10

10

10

0. 93142 0.9568

0.9559

0.9572

0.9693

0.9551* 0.9635

2.1*6 1.31

1.27

0.80

5:1 : 3

: 3 5.5

: 86.6

3

3 0.1 0.6H00

0.63U

0.6727

0.6005

5.11

-11.89

39
Table 1. (cont.)

Shape of Spectra

G(f)

G2

U h h U

0^:02

1:3

1:2

1:1

2:1

3:1

: 1 : 1 :

: 3 : 3 :

: 2 : 2 :

: 1 1 :

: 3 3 :

: 2 2 :

: 1 1 :

: 3 3 :

: 2 2 :

: 1 2 :

: 2 3 :

: 1 1 :

: 3 3 :

: 2 2 :

: 1 2 :

: 2 3 :

: 1 1 :

: 3 3 :

: 2 2 :

: 1 2 :

: 2 3 :

321.666

3

21.666

3

21.666

52.333

321.666

52.333

321.666

52.333

0.5

0.5

0.5

0.5

0.5

0.3

0.3

0.3

0.3

0.3

Value of a by Eq(J4) Iby Eq(8)

0.8687

0.9106

0.9370

0.83^1

0.8872

0.921i4

0.7612

0. 8^112 0.892J4 0.7167

0.8186

0.684

0.8006

0.872

0.6208

0.7600

0.6455

0.7863

0.8680

0.5593

0.7275

Terror"

0.8963

0.9285

0.9495

0.8592

0.9045

0.9335

0.7810

0.8575

0.9050

0.7348

0.8232

0.697

0.8165

0.885

0.6225

0.7640

0.6547

0.8030

0.8828

0.5587

0.7389

3.18

1.97

1.33

3.01

1.95

1.31

2.60

1 .94

1.41

2.53

1.31

1.90

2.00

1 .49

0.27

0.52

1.43

2.10

1.71

-0.11 1.57

40
References

[1] J. S. Bendat , Principles and Application of Random Noise Theory , John Wiley & Sons, New York

(1958).

[2] S. H. Smith, "Fatigue Crack Growth under Axial Narrow and Broad Band Random Loading," in:

Acoustical Fatigue in Aerospace Structure , Syracuse, New York (1964), pp. 331-36O.

[3] F. Beer, R. Wagner, L. Bahar , and R. Ravera, "On the Statistical Distribution of Rises and

Falls in a Stochastic Process," Lehigh Institute of Research Progress Report, Lehigh

University, Bethlehem, Pennsylvania (1951). (Fracture Mechanics Research for the Boeing

Airplane Company).

[ij] J. R. Rice, "Theoretical Prediction of Some Statistical Characteristics of Random Loadings

Relevant to Fatigue and Fracture," Ph.D. Thesis, Lehigh University, Bethlehem, Pennsylvania

(1964).

[5] P. H. Wirsching and M. C. Light, "Probability Based Fatigue Design Criteria for Ocean Structures," API-PRAC Project No. 15, Final Report, American Petroleum Institute, Dallas, Texas

(1979).

[6] L. P. Pook, "Proposed Standard Load Histories for Fatigue Testing Relevant to Offshore Structures," NEL Report No. 624, National Engineering Laboratory, Glasgow, UK (October 1976).

Appendix

Straight lines can be used to approximate the original curves of any kind of power spectra.

For example, figure 1 shows a single-peak power spectrum (dashed line), approximated by four

straight lines (solid line). The important consideration in choosing the straight lines is that the area under the straight line should be the same (or close to) that of the original curve it re- presents. Taking G^ and Gc equal to 0, the four straight lines in figure 1 are expressed by the

following linear equations:

f - f^ G(f) = (Gp'2 - 0) for the first segment

f2- fi

f - f2

G(f) = (Go - Gp) + Gp for the second segment

f3- f2

f - U

G(f) = (G|^ - Go) + G^ for the third segment

f^ - f3

f - f5

G(f) = (0 - Gi^) for the fourth segment

f5 - r^

After integration of equations (3a), (3b), and (3c) and some manipulation, one finds,

41
•• [

2 2 2 2 2 2

(0 - G2) + (G2 - G^) + (Gg - G^)-

2 L ^2 ^1 ^3 ^2 ^4 ^3

2 2

+ (Gi^ - 0) (AD

f5 - f4 J

14 4 4 4 4 4 fp - f^ fo - f2 f^ - fo

M2 =

I (0 - G2) + (G2 - G3) + (G3 - G^)

12 I ^2 " ^1 ^3 " ^2 ^4 ~ ^3

+ (Gj^ - 0) I (A2)

^5

1 r ^2 "" ^1 ^3 " ^2 ^4 ^3

M^ = (0 - G2) + (G2 - Go) + (Go - G^)

30 L f2 - f^ fj - f2 fi4 - f3

fc: - fl, J

+ (Gn - 0) I (A3)

Equations (A1), (A2), and (A3) can be written in a general form:

1 n f{if ^- fP'^^

M. = I (Gi - G^^^)

-^ , j = 0, 2, 4 (A4)

(j + l)(j+2) i = 1 f^^i - f^

which can be used for single-peak as well as double-peak power spectra.

42
FATIGUE CRACK GROWTH IN AREAS OF STRESS CONCENTRATION

PLASTICITY AND SMALL-CRACK EFFECTS

Pei-Ning Li Yi-Wen Cheng

Fracture and Deformation Division

National Bureau of Standards

Boulder, Colorado 80303

Abstract

Hole-in-plate specimens, made of ABS EH36 steel, were tested in air and in 3.5 percent NaCl

solution to measure fatigue crack growth rates in areas of stress concentration at the edges of

circular holes. The linear-elastic fracture mechanics analysis of fatigue crack growth appears to be adequate in yielded regions that are caused by monotonic loading but which have a linear stressstrain relation under cyclic loading. Small-crack behavior, in which cracks grow at higher rates

than predictions from the long cracks, was observed when the crack length was less than 1.5 mm in

air and less than 3 mm in 3.5 percent NaCl solution. The problem of small-crack behavior was accounted for by adding an intrinsic crack length to the physical crack length, as suggested by

El Haddad.

Key words: C-Mn steel; environmental effects; fatigue crack growth; small cracks; yielding effects,

Guest worker, on leave from East-China Institute of Chemical Technology, Shanghai, China.

43
Introduction

The fatigue life of a structural component is determined by the sum of the applied load cycles

required to initiate a crack and to propagate the crack from subcritical to critical size. Because

welded structures, such as offshore structures, usually contain weld defects at areas of stress

concentrations, the fatigue life depends mainly on the time required for crack propagation: a fatigue fracture mechanics analysis, therefore, is appropriate. The initial weld defects are small

and will propagate in the plastically deformed regions near stress concentrators.

Fatigue crack propagation in plastically deformed regions of structural stress concentrators,

such as at weld toes and notch roots, has been studied by many investigators [1-5] . They observed

that the linear-elastic fracture mechanics (LEFM) method was inadequate for predicting the fatigue

crack growth rate (FCGR) in this area. Some parameters have been proposed for correlation with

FCGR. Solomon [1] suggested that the plastic strain range, Ae , could be used to predict the crack

propagation rate, but this worked only for large plastic deformations. El Haddad et al. [2-i»] used

the strain intensity factor range instead of the stress intensity factor range as the driving force

for fatigue crack propagation. Dowling [5] proposed that for several kinds of specimens the J- integral was an adequate parameter for correlation with FCGR. These proposed parameters and

subsequent correlations with FCGR are empirical or semiempir ical in nature. Their ranges and

conditions of applicability need to be defined.

Another problem associated with the fatigue crack propagation in plastically deformed regions of

structural stress concentrators is the unexpected rapid growth of small cracks. Fatigue crack

growth rates of small cracks are higher than those predicted by the results from long cracks. The use of existing long-crack results for defect-tolerance fatigue-life calculations in components,

where the growth of a small crack represents a large portion of the fatigue life, leads to non- conservative life predictions. To account for the higher crack growth of small cracks.

El Haddad et al. [2] introduced the notion of an intrinsic crack length, a^, which is added to the

physical crack length. The value of a^ is constant for a given material condition and environment.

The term (a + a ) is viewed as an effective crack length and" the effective stress intensity factor

range is

AK = AS [it (a + a^)!'^-'^ Fq (^^

where AK is the stress intensity factor range, AS is the nominal stress range, and F is a geometrical factor. The value of a can be evaluated from the limiting condition of a smooth specimen,

where the physical crack length, a, approaches zero. When F is unity and AK becomes the threshold

stress intensity factor range, AK^j^, then AS approaches the fatigue limit of the material, AOg.

Therefore, from equation (1

(-)(^) (2)

Numbers in brackets denote references listed at the end of each paper,

44
In this paper, we report FCGRs in areas of stress concentrations and of small cracks, using the

hole-in-plate specimens in air and in 3.5 percent NaCl solution (saltwater).

Experimental Procedures

Test Material

The test material was a 1 2.7-mm-thick plate of ABS grade EH36 steel, a 350-MPa-yield-3trength

C-Mn steel. The chemical composition of the steel was given in reference 6. The steel was in the

normalized condition and had particularly uniform properties owing to sulfide shape control.

Tensile, fracture [7], and fatigue crack growth [6] properties of the steel have been studied

extensively (reference 6 is in this report). The tensile and fracture properties at ambient

temperature are listed in table 1

Specimen Preparation

The test specimens were 1 2.7-mm-thick hole-in-plate tensile panels. The test matrix and

specimen dimensions are given in table 2. The specimen configuration is shown in figure 1. A circular hole was drilled at the center of the plate and a fatigue crack was initiated at the edge

of the hole so that fatigue crack growth behavior in areas of stress concentration could be studied.

Pin-loading holes at the ends of the specimen were reinforced by welding on a doubler plate.

Except for specimen 1, which did not have notches, sharp notches about 0.5 to 1 mm in length

were machined from the edge(s) of the hole normal to the loading direction with a slitting saw. The

sharp notches were used as crack starters to facilitate fatigue precracking.

Table 1 . Tensile and fracture properties of ABS grade EH36 steel at ambient temperature [7].

Upper Yield Point: 331 MPa

Lower Yield Stress: 326 MPa

Ultimate Tensile Strength: 496 MPa

Fracture Stress: 1246 MPa

Elongation: 39.1 % Reduction in Area: 77.6 % Charpy V-Notch Absorbed Energy: -346 J Fracture Toughness*, CTOD: 0.481 mm Fracture Toughness""" , Jj^: 241 N mm

-1

3-point bend specimen with thickness = 25.4 mm.

Solutions for the stress intensity factor, K, of the specimen are available in reference 8. The

stress intensity factor is given as

45
Table 2. Test matrix and specimen dimensions,

Specimen Specimen Center Hole

No. Width, mm Diameter, mm Crack

Test Test

Environment Frequency, Hz

1 203 50,8 No Air

2 254 50.8 Asymmetric

(One Crack)

Air

3 254 50.8 Asymmetric

(One Crack)

Saltwater

4 254 50.8 Symmetric

(Two Cracks)

Air

0.1

•12.7 mm

^

i

Figure 1 . Schematic of hole-in-plate specimen.

46
'' K = S (it a)°-5 F(a/D) (1)

where the crack length, a, is measured from the edge of the hole, D is the hole diameter, S is the

remote tensile stress, and F(a/D) is a function of crack length and hole diameter.

Loading Conditions

Except for specimen 1 , specimens were cyclically loaded at ambient temperature with frequencies

of 3 Hz in air and 0.1 Hz in saltwater using load control with a 1 -MN-capacity servo-controlled

hydraulic testing machine. The stress ratio (i.e., the ratio of minimum to maximum stress) was kept

constant at 0.3-

Specimen 1 , which had no crack, was instrumented with eight electrical-resistance strain gages

extending from the edge of the hole to the edge of the test plate perpendicular to the loading

direction. The gages were spaced 2.5 mm near the edge and 25-^ mm away from the hole. The purpose

of testing specimen 1 was to study the strain distribution at the stress concentration under

loading-unloadlng-reloading sequence. The loading and unloading were controlled manually under

displacement control.

Test Environments and Crack-Length Measurements

Tests were conducted in air and in saltwater at ambient temperature. Crack-length measurements were made with a 30-power traveling microscope at various time intervals, depending upon crack

propagation rates. For tests in saltwater, a transparent plastic container was used to contain the

saltwater, which was continuously circulated at a rate of 26 )l/min through a diatomaceous-earth

filter. The NaCl concentration, temperature, and pH value of the saltwater were monitored periodically.

For tests in saltwater, crack lengths were measured with the traveling microscope through the

transparent container. The rust around the crack tips was scrubbed from the specimen surface with

sandpaper and cotton swabs before measurements.

Crack closure was monitored through the load-displacement curves with an x-y recorder or an oscilloscope. Displacements were measured at the crack mouth. To facilitate the displacement

measurements, razor blades, spot welded at the crack mouth and extending out of the specimen plane, were used for attachment of the clip-on gage.

Experimental Results and Discussion

Strain Survey

Specimen 1, which had no crack, was monotonically loaded from zero to 220.^4 MPa (nominal

gross-section stress) and then unloaded to zero nominal stress. The measured local strains along

the direction of the hole diameter normal to the loading direction are plotted in figure 2.

In the elastic range, that is, for nominal stress less than one-third the yield strength, the

47
strain distribution in the vicinity of a circular hole can be calculated accurately [9]. The edge

of the hole begins to deform plastically when the nominal stress is higher than one-third the yield

strength. If the extent of plastic deformation is small, the stress redistribution owing to the

plastic deformation is negligible, and the elastic solutions [9] are still adequate for regions that

do not yield. The results of the present Investigation indicate that elastic solutions are adequate

(the error is within 3 percent) for areas 2 mm or farther from the hole edge at a nominal stress of

155.1 MPa, which is much higher than one-third the yield strength (117 MPa)

.

The strain increases rapidly after the material has yielded. Strain distributions in the

elastic-plastic case can be estimated from Neuber's rule [10]. To obtain accurate results,

numerical methods, such as finite element analysis, must be performed.

After the specimen was unloaded to zero nominal stress, as shown in figure 2, residual strains

existed over a large region (>50 mm). The specimen was reloaded and strain increments, Ae , were recorded. A strain increment is the difference between the current measured strain and the residual

strain, Ae = ^current ~ ^residual* ^^ shown in figure 3, as nominal stress increased from 115.3 to 220. ^ MPa (about double) at a point 2 mm from the hole edge, Ae increased from 1.52 x 10~^ to 3.12 X 10 -^ (about double). The relation between nominal stress and local strain increment is linear

owing to strain hardening of the material in the plastically deformed region. Because of the linear

relation between stress and strain at stress levels above monotonic yield strength, LEFM analysis of

FCGR in this region should be adequate. However, the residual stress associated with the residual

strain might influence FCGR. The residual stress is in compression and effectively reduces the

mini mum -to -maximum stress ratio; this usually lowers FCGR, especially at higher and near -threshold

growth rate regions.

FCGR at Edges of a Yielded Hole in Air

Specimen 4, which had two symmetric cracks emanating from opposite edges of the hole, was fatigue tested with a maximum stress, S_„^, of 207 MPa and a minimum stress, S_,.„, of 52.1 MPa in

air. With a stress concentration factor of 3-1 4, the maximum and minimum local stresses at the edge

of the hole were 550 and 195 MPa, respectively. Fatigue crack growth rates were measured as the

cracks propagated from 0.5 to 9 mm, which was within the yielded region caused by the application of

a nominal stress of 207 MPa (figure 2).

The FCGR results, plotted in figure 4, show good agreement between the hole-in-plate specimen

and compact-type (CT) specimens [5] at AK higher than 30 MPa/m. Below 30 MPa/m, the FCGRs of the

hole-in-plate specimen are higher than those obtained from CT specimens, which is not unexpected.

Usually, the length of the existing crack is larger than 15 mm for a 25 .^-mm-thick CT specimen,

which was used in reference 6 and the crack behaves like a normal or long crack. In the hole-inplate specimen, the crack length was about 1.5 mm when AK was 30 MPa/m, and the crack behaves like a small crack. As mentioned previously, FCGRs of small cracks are higher than those predicted by

results from long cracks.

Using the approach of El Haddad et al. [2-4], we found the value of a^ in equation (1) to be

0.35 mm. Then the whole da/dN-versus-AK curve of the hole-in-plate specimen agreed well with the

results from CT specimens.

48
COo

<

CC

H

CO

10 20 30 40 50

DISTANCE FROM HOLE EDGE, mm

Figure 2. Measured strain distribution in the vicinity of a circular hole at various nominal stress levels.

COO

X

LU

LU

DC

O

<

DC

&5

10 20 30 40 50 60

DISTANCE FROM HOLE EDGE, mm

Figure 3. Measured strain distribution in the vicinity of a circular hole at various nominal stress levels after one loading-unloading sequence.

49
10

AK, ksi/In

20 50 90

10-2

10-3

o

>.o

E

E

CO 10-4

10,-510

I \ r T

• Specimen 4 in air,

R=0.3

CT specimen,

R=0.5

R=0.1

I J L

-J 10-4

o

o

>»o

10-5

CO o

10-6

20 50

AK, MPafrn

100

Figure iJ . Comparison of fatigue crack growth rates in plastically deformed

regions and compact-type specimens in air.

.50
Specimen 2, which had only one crack emanating from one edge of the hole, was fatigue tested

with a loading condition identical to that applied to specimen 4. The FCGR results are shown in

figure 5, which show reasonable agreement between the hole-in-plate specimen and CT specimens [6].

The results of this investigation show that the LEFM analysis of fatigue crack growth ip

adequate in yielding conditions under monotonic loading. This observation is consistent with one other study [6] in which a deeply-grooved CT specimen was tested. However, it should be cautioned

that the applicability of the LEFM analysis of fatigue crack growth in yielding conditions (under

monotonic loading) is probably limited to conditions where the local stress-strain relation under

cyclic loading remains linear.

FCGR at Edges of a Yielded Hole in Saltwater

Specimen 3. which had one crack emanating from one edge of the hole, was fatigue tested with a loading history identical to that applied to specimen 4 in saltwater. The FCGR results are shown in

figure 6. Reasonable agreement between results from CT specimens [6] and the present study is

observed, indicating that in saltwater as well as in air the LEFM analysis of fatigue crack growth

is adequate in yielding conditions under monotonic loading.

As shown in figure 6, the small-crack behavior occurs at a AK level of about 45 MPa/m, which

corresponds to a crack length of 3 mm, twice as long as that observed in air (1.5 mm). The reason

for this is not clear and further study is needed.

The FCGRs of hole-in-plate specimens are slightly lower than those of CT specimens because the K solutions [8] for a crack at the edge of a hole were larger than the experimental data [11,12].

Consequently, the FCGR is lower when a calculated K is used. In the case where there are two cracks

at the edge of a hole, results obtained with a calculated K agree with experimental results.

Summary and Conclusions

Even though significant plastic deformation exists in areas of stress concentration in structural components, the fatigue crack growth rates in those areas can be predicted very well using the

LEFM analysis, provided that the local stress-strain relation is linear under cyclic loading and the

crack is long enough that the small-crack behavior is absent. Small cracks grow at higher rates than those predicted from da/dN-versus-AK results for long cracks. The problem of small-crack

behavior can be accounted for by adding an intrinsic crack length to the physical crack length, as suggested by El Haddad. The intrinsic crack length appears to be dependent upon environmental

conditions for a given material. It is longer in saltwater (3 mm) than in air (1.5 mm)

Acknowledgments

This worked was supported by the Department of Interior, Minerals Management Service.

Drs. H. I. McHenry and D. T. Read provided helpful discussions.

51
10

10-2

o

o

^.

E

E

Z

•a

^^

CO

10"3|-

T3 10r4

10r5

10

AK, ksiVIn

20 50 —

I

1

1 r

90

1

\

\ 1 I

I l-l

• Specimen 2 in air,

R=0.3

CT specimen,

R=0.1

R=0.5

J \ \ I I I I IJ

- 10-4

-101-5

O

O

>»o

z

o

(0 T3

10"

20 50

AK, MPa>^

100

Figure 5. Comparison of fatigue crack growth rates in plastically deformed

regions and compact-type specimens in air.

52
AK, ksl/lrT

10

10

2-r20 50 90

1—I—I—r-TT

o

o

>

o

E

E

(0

10-3

10-4-

10-5

1 I I I I H

• Specimen 3 in saltwater,

R = 0.3

CT specimen

in saltwater

I I I I I I I

10 -4

O

>»o

-llO"^

as •o

J10"®

10 20 50 100

AK, MPa/m

Figure 6. Comparison of fatigue crack growth rates in plastically deformed

regions and compact-type specimens in saltwater.

53
References

[I] H. D. Solomon, "Low Cycle Fatigue Crack Propagation in 1018 Steel," Journal of Materials , Vol. 7, No. 3 (1972), pp. 299-306.

[2] M. H. El Haddad, K. N. Smith, and T. H. Topper, "Fatigue Crack Propagation of Short Cracks,"

Journal of Engineering Materials and Technology , Vol. 102 (1979), pp. 42-46.

[3] M. H. El Haddad, K. N. Smith, and T. H. Topper, "A Strain Based Intensity Factor Solution for

Short Fatigue Cracks from Notches," in: Fracture Mechanics , ASTM STP 677, American Society for

Testing and Materials, Philadelphia (1979), pp. 274-284.

[4] M. H. El Haddad, N. E. Dowling, T. H. Topper, and K. N. Smith, "J- Integral Applications for

Short Fatigue Cracks at Notches," International Journal of Fracture , Vol. 16, No*. 1 (1980),

pp. 15-30.

[5] N. E. Dowling, "Geometry Effects and the J- Integral Approach to Elastic-Plastic Fatigue Crack

Growth," in: Crack and Fracture , ASTM STP 601, American Society for Testing and Materials,

Philadelphia (1977), pp. 19-32.

[6] Y. W. Cheng, "The Fatigue Crack Growth of a Ship Steel in Saltwater under Spectrum Loading ," in this report, pp. 15-30.

[7] T. L. Anderson, "The Effect of Crack-Tip Region Constraint on Fracture in the Ductile-toBrittle Region," Ph.D. thesis, Colorado School of Mines, Golden, Colorado (1983).

[8] H. Tada, P. C. Paris, and G. R. Irwin, The Stress Analysis of Cracks Handbook , Del Research

Corp.. Hellerton, Pennsylvania (1973), pp. 19.2, 19.4, and 19.9.

[9] G. N. Savin, Stress Concentration Around Holes , translated from the Russian by E. Gros,

Pergamon Press, New York (1961), pp. 104-113.

[10] H. Neuber, "Theory of Stress Concentration for Shear-Strained Prismatical Bodies with Arbitrary

Nonlinear Stress-Strain Law," Journal of Applied Mechanics , Vol. 28 (1961), pp. 544-550.

[II] P. N. Li and Y. W. Cheng, "High/Low Stress Amplitude Effects on Fatigue Crack Growth of a Ship

Steel in Air and in Saltwater," in this report, pp. 55-65.

[12] J. S. Cargill, J. K. Malpani, and Y. W. Cheng, "Disk Residual Life Studies," AFML-TR-79-4173,

Air Force Materials Laboratory, Dayton, Ohio (1979).

54
HIGH/LOW STRESS AMPLITUDE EFFECTS ON FATIGUE CRACK GROWTH RATES

OF A SHIP STEEL IN AIR AND IN SALTWATER

Pei-Ning Li Yi-Wen Cheng

Fracture and Deformation Division

National Bureau of Standards

Boulder, Colorado 80303

Abstract

Hole-in-plate specimens, made of ABS grade EH36 steel, were tested in air and in 3.5 percent

NaCl solution (saltwater) to study the high/low stress amplitude effects on fatigue crack growth

rates (FCGRs) in an elastic stress field and at the edge of a yielded hole. Effects of tensile

overload prior to crack initiation on subsequent FCGR at the edge of a yielded hole were also

investigated. The results are summarized as follows:

1

.

The tensile overload retardation effects were similar in an elastic stress field and at the

edge of a yielded hole.

2. The tensile overload retardation effects were similar in air and in saltwater.

3. The Bowie analysis overestimated the stress intensity factors when compared with experimental

results.

4. The simple engineering approach gave accurate stress intensity factors when compared with

experimental results, except in areas close to the edge of a hole. In the latter case, the

simple engineering approach overestimated the stress intensity factors owing to overestimation

of crack length.

5. Tensile overload prior to crack initiation appeared to retard the subsequent FCGR at the edge

of a yielded hole. The retardation was explained by the presence of beneficial residual

stresses and crack closure.

Key words: C-Mn steel; environmental effects; fatigue crack growth; load-sequence interaction

effects; yielding effects.

Guest worker, on leave from East-China Institute of Chemical Technology, Shanghai, China.

55
Introduction

The fracture mechanics approach has been successfully applied to describe fatigue crack growth

under constant-amplitude loading. However, fatigue life predictions of structural components remain

complicated and imprecise owing to lack of methodologies that accurately account for the irregular

nature of service loading. Load-sequence interaction effects have been qualitatively recognized for

a long time, but because of numerous factors involved, detailed quantitative characterization has

not well been established. Significant effects observed by many investigators can be summarized as follows:

1. Tensile overloads cause retardation of fatigue crack growth. A sufficiently high tensile overload may stop the growth of a fatigue crack completely [1].

2. Preceding lower cyclic loads cause a slight acceleration of fatigue crack growth [2].

3. A compressive overload preceding a tensile overload reduces the retardation caused by the tensile overload [3].

k. A compressive overload following a tensile overload significantly reduces the retardation

effects caused by the tensile overload [3].

5. Compressive overloads slightly accelerate fatigue crack growth [3].

For sea loading on offshore structures, all the aforementioned load-sequence interaction

effects occur. In the case of offshore structures, one has to consider the possibility of additional interactions between load sequence and strong environmental effects due to saltwater. The

investigation reported in this paper pursued the following objectives: (1) to compare the high/low

stress amplitude effects on fatigue crack growth rates (FCGRs) in air and in 3.5 percent NaCl solution (saltwater), (2) to compare the high/low stress amplitude effects on FCGRs in an elastic

stress field and at the edge of a yielded hole, and (3) to study the effects of tensile overload prior

to crack initiation on subsequent FCGR.

Experimental Procedures

This investigation was a continuation of that reported in reference ^ (reference H is in this

report). The test material, specimen preparation, test environment and crack length measurement

technique of this investigation were identical to those described in reference 4. The test variables, which were different from reference 4, are described in the following.

The maximum and minimum loads were step decreased during the test, while the minimum-to-maximum

stress ratio was kept constant at 0.3, to study the high/low stress amplitude effects. The initial

maximum and minimum nominal stresses, S„^^ and S„.„, were 207 and 62.1 MPa. The stresses were step

Ilicl A III J.I1 decreased to 147.9 and 44.36 MPa, and then to 105.6 and 31-69 MPa.

In addition to the test matrix given in table 2 of reference 4, specimen 5 was used to study

the influence of tensile overload which was applied before crack initiation on the subsequent

fatigue crack propagation at stress concentrations. Specimen 5, which had no crack initially with a

Numbers in brackets denote references listed at the end of each paper,

56
hole diameter of 4M.45 mm, was monotonically loaded to a nominal stress of 242 MPa. After being

unloaded, two small notches of 0.5 mm in length were introduced with a jeweler's saw on opposite

edge of the hole normal to the loading direction. The specimen was fatigue cycled at stress levels

of S_„^ = 20? MPa and S_,,-_ = 62.1 MPa. After each crack had grown 3.2 mm, the specimen was

ula.X [lliri remachined to enlarge the hole diameter to 50.8 mm. The specimen was then fatigue cycled again at the same cyclic stress levels iS^^^ = 207 MPa and S^j^^^^ = 62.1 MPa).

Experimental Results and Discussion

Effects of High/Low Stress Amplitude on FCGR in Air

Specimens 2 and 4 (in reference to table 2 of reference 4) were tested to study the effects of

high/low stress amplitude on FCGR in air in the plastically deformed regions. The results, along

with the detailed loading histories, are given in figures 1 and 2 for specimens 2 and 4,

respectively. At the beginning of testing specimen 2, which had a single crack emanating from the

edge of the hole, the maximum nominal stress, S^^^^^^, and the minimum nominal stress, S . , were first

decreased from 207 and 62.1 MPa to 115 and 34.5 MPa, a load reduction ratio of 1.8. The overload

effect was so strong that there was no observed crack growth in 90.9 kilocycles. Subsequently, S_„^

and Sjjj^^ were increased to 147.9 and 44.36 MPa, respectively.

The values of AK, as shown in figure 1, were calculated using the Bowie analysis [5], the

simple engineering approach [6] (which considers the hole as a portion of the crack), and the ex- perimental method. In the experimental method, values of AK were inferred, using the da/dN-versusAK results of CT specimens [7], from the measured da/dN of the hole-in-plate specimen.

As shown in figure 1 , the trend of experimental AK without considering crack closure or retardation effects is lower than predictions of both the Bowie analysis and the simple engineering

approach. However, if crack closure is considered, the apparent stress intensity factor range

(AKg ) , plotted as crosses in figure 1, increases and agrees very well with predictions from the

simple engineering approach.

At crack lengths less than 5 mm, the simple engineering approach overestimates AK because it

overestimates the crack length by considering the hole as a portion of the crack. The Bowie analysis overestimates AK in the whole range studied. The present results are consistent with those of

another study [8] in which a single fatigue crack propagated from the bolt hole of a compressor

disk. The crack length at which the simple engineering solution crosses over the Bowie analysis

depends on the size of the hole; it increases with increasing hole diameter.

As expected, immediately after the load reduction the FCGRs were lower than the trend line,

indicating that the fatigue crack growth was retarded (figure 1). There was no observed crack growth

in 90.9 kilocycles after a load reduction ratio of 1.8.

The measured widths of retardation effect were in reasonable agreement with predictions by

Willenborg model [9]: 4.5 and 6.5 mm versus 4.9 mm. The Willenborg model assumes that, as long as

the instantaneous plastic zone is within the prior high-load-affected plastic zone, crack growth

retardation will be present. The retardation decays to zero when the instantaneous plastic zone

57
CO

<

60

50

40 -

30

20 -

10

207 MPal I I I 147.9 MPa loading history

105.6 MPa

62.1 MPa 44.36 "MPa 31.69 MPa

• •

••• • • •

•••'

Bowie solution x f a+R\

W /

Engineeri ng solution / ^ AK = ASV7r(R+a/2) x U^^j

• Experimental data (specimen 2)

+ Experimental data (corrected for closure)

10 20 30 40

CRACK LENGTH, mm

50

Figure 1 . Theoretical and experimental stress intensity factor ranges as a function of crack length: in air.

58
is

a.

<

S i

62

1 i 1

07 MPa loading history

m 147.9 MPa

,JlMi' 11/

•1 ^"

44.36 PW»>a

60

50

•

40

/

>^ • _ /^ • .V

••

30

•

• 20

5» Newman solution

• Experimental data

(specimen 4)

10

n 1 1 1 10 20 30

CRACK LENGTH, mm

Figure 2. Theoretical and experimental stress intensity factor ranges as a function of crack length: in air.

59
passes beyond the prior high-load-affected plastic zone.

As indicated in figure 2 of reference ^, at a nominal stress of 207.5 MPa the yielded region

extended from the hole edge to about 19 mm away from the hole edge. Thus, at the first load reduction, the crack tip with a crack length of 9 mm was within the yielded region. At the second load

reduction, the crack tip with a crack length of 26 mm was outside the yielded region. The AK- versus-a trends (figure 1 ) of the first and the second load reductions are similar, indicating that

the high/low stress amplitude effects on FCGR are similar in the elastic and in the yielded regions.

However, note that LEFM analysis of FCGR in the yielded region is probably limited to conditions

where the stress-strain relation remains linear under cyclic loading. If it is cyclically yielded,

the analysis will probably not be valid.

The test results for specimen 4, which had two symmetrical cracks emanating from opposite edge

of the hole, are shown in figure 2. The available K solution is from an analysis by Newman [5]. As shown in figure 2, the experimental AK values are in excellent agreement with those of analytical

predictions except in the areas affected by the load reduction.

The test results of specimen 4 differed from those of specimen 2 in three ways: First, no apparent crack closure was observed. Second, the measured width of retardation effect (17 mm) was much larger in specimen ^ than in specimen 2. Third, the delayed retardation effect was more pronounced. The only difference between specimens 2 and 4 is that specimen 2 had a single crack

emanating from one edge of the hole and specimen 4 had two symmetrical cracks emanating from

opposite edges of the hole. To conform that these differences are geometry dependent, additional

tests are needed.

Effects of High/Low Stress Amplitude on FCGR in Saltwater

Specimen 3i which had one crack emanating from one edge of the hole (identical to specimen 2), was fatigue tested with a loading history identical to that applied to specimen 4 in saltwater. The

results obtained in saltwater, as shown in figure 3, are similar to those of specimen 2 in air

(figure 1): (1) The experimental AK values, without considering crack closure, were smaller than

those predicted by the Bowie analysis and by the simple engineering approach. (2) Overloading

effects were observed, and the width of the retardation effect was in good agreement with that predicted by the Willenborg model. This indicates that the high/low stress amplitude effects are practically the same in saltwater as in air.

Effects of Overload Prior to Crack Initiation on FCGR

Specimen 5, containing a 44 . 5-mm-diameter center hole without a notch, was preloaded to a remote stress of 242 MPa. After being unloaded, two small notches of 0.5 mm in length were introduced by a jeweler's saw on opposite edges of the hole. Then the specimen was fatigue cycled

at stress levels of S^^^ = 207 MPa and S^^^ = 52.1 MPa, identical to that applied to specimen 4. As the distance between the two crack tips increased beyond 50.8 mm, the specimen was remachined to enlarge the hole diameter to 50.8 mm, while the center of the hole was adjusted such that lengths of

the remaining small cracks were identical. Finally, the specimen was cycled -again at the same cyclic stress levels (S^^^^ = 207 and S^^^^ = 62.1 MPa).

60
10

AK, ksiVln

20 50

10-2 T

O

>»o

E

E

03 T3 lO"**

10-5

Specimen 5 In air, R=0.3

Crack Length:

from 0.43 to 4.4 mm

10 20 50

Ak, MPafm

90

T TTT

1—f~rii

- 10-*

I

' I I'll

o

o

o

- 10-5 .E

CO •a

10-6

100

Figure 3. Theoretical and experimental stress intensity factor ranges as a

. function of crack length: in saltwater.

61
The FCGRs plotted in figures 4 and 5 are much lower than those of specimen 4 (figure 2), which

had no prior overloading. The solid line in figure 4 was derived from Newman analysis [5], and the

dots were inferred from the experimental method. The different behaviors exhibited by specimens 4 and 5 can be explained as follows: Since the hole edge of specimen 5 was yielded during prior over- loading, beneficial compressive residual stresses were developed upon unloading. Moreover, when a crack is present in the prior yielded region, the crack surfaces will touch each other before the

specimen is unloaded to the minimum load. This means that, in addition to a compressive residual

stress field at the crack tip, crack closure occurs and the FCGR decreases. However, small-crack

behavior counterbalanced the retardation effect at crack lengths less than 0.5 mm (figure 5).

Summary and Conclusions

The retardation in fatigue crack growth rates owing to high/low stress amplitude was studied

with hole-in-plate specimens. The following observations were made:

1The retardation effects were similar in elastic stress fields and at the edges of a yielded

hole.

2. The retardation effects were similar in air and in saltwater.

3. The Bowie analysis overestimated the stress intensity factors, when compared with experimental

results.

4. The simple engineering approach gave accurate stress intensity factors, when compared with ex- perimental results, except in areas close to the edges of a circular hole. There the stress

intensity factors were overestimated because the crack length was overestimated.

5. Tensile overload prior to crack initiation appeared to retard the subsequent fatigue crack

growth rates in areas of stress concentration. The retardation was explained by the presence

of beneficial residual stresses and crack closure.

Acknowledgments

This work was supported by the Department of Interior, Minerals Management Service.

Drs. H. I. McHenry and D. T. Read provided helpful discussions.

62
0.

<

^Newman Solution • Specimen 5 in air R = 0.3

Cracl< Length:

from 0.43 to 4.4 mm

0.1 0.2 0.3

2a/D

Figure 4. Theoretical and experimental stress intensity factor ranges as a function of the crack length-to-hole diameter ratio: in air. a is

the crack length and D is the hole diameter.

63
^(0 o. <

30 —

20

10

S a

207 MPa

1 loading history

147.9 MPa

105.6 MPa

31.69 MPa

• • • • ^'^

/a + R\

Bowie solution X f ^ ^ ) •— Engineering solutions • Experimental data

(specimen No. 3)

In saltwater

10 20

CRACK LENGTH, mm

30

Figure 5. Comparison of fatigue crack growth rates at a preyielded hole with

those of CT specimens.

64
References

[1] CM. Hudson and H. F. Hardrath, "Investigation of the Effects of Variable-Amplitude Loadings

on Fatigue Crack Propagation Patterns," TND-I803, NASA (1963).

[2] E. F. J. von Euw, R. W. Hertzberg, and R. Roberts, "Delay Effects in Fatigue Crack

Propagation," in: Stress Anlaysis and Growth of Cracks , ASTM STP 513, American Society for

Testing and Materials, Philadelphia (1972), pp. 230-259.

[3] R. I. Stephens, D. K. Chen, and B. W. Hom, "Fatigue Crack Growth with Negative Stress Ratios

Following Single Overloads in 2024-T3 and 7075-T5 Aluminum Alloys," in: Fatigue Crack Growth

under Spectrum Loads , ASTM STP 595, American Society for Testing and Materials, Philadelphia

(1976), pp. 27-i<0.

[4] P. N. Li and Y. W. Cheng, "Fatigue Crack Growth in Areas of Stress Concentration — Plasticity

and Small-Crack Effects," in this report, pp. 43-5'^.

[5] H. Tada, P. C. Paris, and G. R. Irwin, The Stress Analysis of Cracks Handbook , Del Research

Corp., Hellerton, Pennsylvania (1973), PP. 19.2, 19.4, and 19.9.

[6] D. Broek, Elementary Engineering Fracture Mechanics , 3rd edition, Martinus Nijhoff Publishers,

The Hague, The Netherlands (1982).

[7] Y. W. Cheng, "The Fatigue Crack Growth of a Ship Steel in Saltwater under Spectrum Loading," in

this report, pp. 15-30.

[8] J. S. Cargill, J. K. Malpani, and Y. W. Cheng, "Disk Residual Life Study," AFML-TR-79-41 73, Air

Force Materials Laboratory, Dayton, Ohio (1979).

[9] J. Willenborg, R. M. Engle, and H. A. Wood, "A Crack Oowth Retardation Model Using an Effective Stress Concept," AFFDL-TM-71 -1-FBR, Air Force Flight Dynamics Laboratory, Dayton, Ohio

(1971).

65
NBS-n4A (REV. 2-6C) U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET (See instructions)

1. PUBLICATION OR

REPORT NO.

NBS TN-1088

2. Performing Organ. Report No. 3. Publication Date

August 1985

4. TITLE AND SUBTITLE

Fitness- for-Service Criteria for Assessing the Significance

of Fatigue Cracks in Offshore Structures

5. AUTHORCS)

Yi-Wen Cheng

6. PERFORMING ORGANIZATION (If jolnl or other than N&S, see instruct/onsj

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

7. Contract/Grant No.

8. Type of Report & Period Covered

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street. Cit>'. State, ZIP)

Department of the Interior

Minerals Management Service

Reston, Va 22091

10. SUPPLEMENTARY NOTES

I I Document describes a computer program; SF-185, FlPS Software Summary, is attaciied.

11. ABSTRACT (A lOO-word or less (actual summary of most significont informotion. If document includes a significant

bibliography or literature survey, mention it here)

Results of a research program to develop f itness-for-service criteria for

assessing the significance of fatigue cracks in offshore structures are presented

in five papers. Each paper describes the goals and approaches to a specific

task and details the results of the study.

12. KEY WORDS (S/x to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolons)

fatigue crack growth; f itness-for-service criteria; fracture mechanics;

offshore structures; random-loading fatigue

13. AVAILABILITY

r^ Unlimited

[ I For Official Distribution. Do Not Release to NTIS

f"X] Order From Superintendent of Documents, U.S. Government Printing Office, Washington, D.C.

20402.

I I Order From National Technical Information Service (NTIS), Springfield, VA. 22161

14. NO. OF

PRINTED PAGES

72

15. Price

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special publications appropriate to this grouping such as wall charts, pocket cards, and bibliographies.

Applied Mathematics Series—Mathematical tables, manuals, and studies of special interest to physicists,

engineers, chemists, biologists, mathematicians, computer programmers, and others engaged in scientific and

technical work.

National Standard Reference Data Series—Provides quantitative data on the physical and chemical properties of materials, compUed from the world's literature and critically evaluated. Developed under a worldwide pro- gram coordinated by NBS under the authority of the National Standard Data Act (Public Law 90-396).

NOTE: The Journal of Physical and Chemical Reference Data (JPCRD) is published quarterly for NBS by

the American Chemical Society (ACS) and the American Institute of Physics (AIP). Subscriptions, reprints,

and supplements are available from ACS, 1155 Sixteenth St., NW^ Washington, DC 20056.

Building Science Series—Disseminates technical information developed at the Bureau on building materials,

components, systems, and whole structures. The series presents research results, test methods, and perfor- mance criteria related to the structural and environmental functions and the durability and safety

characteristics of building elements and systems.

Technical Notes—Studies or reports which are complete in themselves but restrictive in their treatment of a

subject. Analogous to monographs but not so comprehensive in scope or definitive in treatment of the subject

area. Often serve as a vehicle for final reports of work performed at NBS under the sponsorship of other government agencies. Voluntar> Product Standards—Developed under procedures published by the Department of Commerce in Part 10, Title 15, of the Code of Federal Regulations. The standards establish nationally recognized re- quirements for products, and provide all concerned interests with a basis for common understanding of the

characteristics of the products. NBS administers this program as a supplement to the activities of the private

sector standardizing organizations.

Consumer Information Series—Practical information, based on NBS research and experience, covering areas of interest to the consumer. Easily understandable language and illustrations provide useful background

knowledge for shopping in today's technological marketplace.

Order the above NBS publications from: Superintendent of Documents, Government Printing Office,

Washington, DC 20402. Order the following NBS publications—FIPS and NBSIR 's—from the National Technical Information Service, Springfield, VA 22161.

Federal Information Processing Standards Publications (FIPS PLfB)—Publications in this series collectively constitute the Federal Information Processing Standards Register. The Register serves as the official source of

information in the Federal Government regarding standards issued by NBS pursuant to the Federal Property

and Administrative Services Act of 1949 as amended, Public Law 89-306 (79 Stat. 1127), and as implemented

by Executive Order 11717 (38 FR 12315, dated May 11, 1973) and Part 6 of Title 15 CFR (Code of Federal

Regulations).

NBS Interagency Reports (NBSIR)—A special series of interim or final reports on work performed by NBS

for outside sponsors (both government and non-government). In general, initial distribution is handled by the sponsor; public distribution is by the National Technical Information Service, Springfield, VA 22161, in paper copy or microfiche form.
U.S. Department of Commerce

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