FRACTURE MECHANICS:
SEVENTEENTH VOLUME
Seventeenth National Symposium
on Fracture Mechanics
sponsored by
ASTM Committee E-24
on Fracture Testing
Albany, New York, 7-9 August 1984

ASTM SPECIAL TECHNICAL PUBLICATION 905
J. H. Underwood, U.S. Army Armament
Research & Development Center, R. Chait, U.S.
Army Materials & Mechanics Research Center,
C. W. Smith, Virginia Polytechnic Institute &
State University, D. P. Wilhem, Northrop
Aircraft, W. A. Andrews, General Electric
Company, and J. C. Newman, NASA Langley
Research Center, editors
ASTM Publication Code Number (PCN)
04-905000-30


Library of Congress Cataloging-in-Publication

nata

National Symposium on Fracture Mechanics (17th:
1984: Albany, N.Y.)
Fracture mechanics.
(ASTM special technical publication; 90S)
"ASTM publication code number (PCN) 04-905000-30."

Includes bibliographies and index.
1. Fracture mechanics-Congresses.
I. Underwood,
John H. II. ASTM Committee E-24 on Fracture Testing.
III. Title. IV. Series.
TA409.N38 1984
620.1 ' 126
86-8000
ISBN 0-8031-0472-3

Copyright

© by AMERICAN SOCIETY FOR TESTING AND MATERIALS 1986
Library of Congress Catalog Card Number: 86-8000

NOTE
The Society is not responsible, as a body,
for the statements and opinions
advanced in this publication.

Printed in Baltimore,
July 1986

Md.


Dedication
This publication is dedicated to the following group of individuals
and their pioneering work in fracture testing:
.


William F. Brown, Jr .•
James E. Campbell
Roy H. Chirstensen
John Hodge
George R. Irwin
Joseph M. Krafft
William T. Lankford
John R. Low, Jr.
Richard A. Rawe
John E. Srawley
Henry J. Stremba
Charles F. Tiffany

.

Their important contributions were central to the ASTM Special
Committee on Fracture Testing of High Strength Sheet Materials.
forerunner of Committee £-24 on Fracture Testing.
As a tribute to the founders of ASTM Committee £-24 and to the
series of symposia which they helped to establish, the poem on the
following page was offered as a special presentation at the Albany
meeting.


Related
ASTM Publications
Fracture Mechanics: Sixteenth Symposium, STP 868 (1985),
04-868000-30


Fracture Mechanics: Fifteenth Symposium, STP 833 (1984),
04-833000-30

Fracture Mechanics: Fourteenth Symposium-Volume
Analysis, STP 791 (1983), 04-791001-30

I: Theory and

Fracture Mechanics: Fourteenth Symposium-Volume
Applications, STP 791 (1983), 04-791002-30

II: Testing and

Fracture Mechanics (Thirteenth Conference), STP 743 (1981),
04-743000-30

Fracture Mechanics (Twelfth Conference), STP 700 (1980),
04- 700000-30


A Note of Appreciation
to Reviewers

.

The quality of the papers that appear in this publication reflects not only
the obvious efforts of the authors but also the unheralded, though essential,
work of the reviewers. On behalf of ASTM we acknowledge with appreciation
their dedication to high professional standards and their sacrifice of time and
effort.


ASTM Committee on Publications


Contents

,
1

Introduction
ApPLICATIONS
An Application of Fracture Mechanics to a Ship Controllable Pitch
Propeller Crank Ring-p.
D. HILTON, R. A. MAYVILLE;AND
D. C. PEIRCE

S

A New Wide Plate Arrest Test (SCA Test) on Weld Joints of Steels...
for Low Temperature Application-K.
TANAKA,M. SATO,
T. ISHIKAWA,AND H. TAKASHIMA
.

22

Variable Flaw Shape Analysis for a Reactor Vessel under Pressurized
Thermal Shock Loading-c.
Y. YANGAND W. H. BAMFORD


41

Growth Behavior of Surface Cracks in the Circumferential
Solid and Hollow Cylinders-R.
G. FORMANAND
V. SHIVAKUMAR

S9

Fracture

Toughness of Ductile Iron and Cast Steel-w.
K. E. McKINNEY,AND'P. C. GERHARDT,JR.

Plane of

L. BRADLEY,

Effect of Loading Rate on Dynamic Fracture of Reaction Bonded
Silicon Nitride-B.
M. LIAW, A. S. KOBAYASHI,AND
A.F.EMERY
Resistance Curve Approach
Characterization-M.

to Composite Materials
M. RATWANIAND R. B. DEO

7S


9S

108

A Comparison of the Fracture Behavior of Thick Laminated
Composites Utilizing Compact Tension, Three-Point Bend,
and Center-Cracked Tension Specimens-c.
E. HARRIS
AND D. H. MORRIS

124

Residual Strength of Five Boron/Aluminum
Laminates with
Crack-Like Notches After Fatigue Loading-R.
A. SIMONDS

136


SUBCRITICALCRACKGROWTH
Hold- Time Effects in Elevated Temperature Fatigue Crack
Propagation-To
NICHOLASAND T. WEERASOORIYA

155

Interactive Effects of High and Low Frequency Loading on the
Fatigue Crack Growth of Inconel 7IS-A. PETROVICH,
W. BESSLER, AND W. ZIEGLER


169

Creep Crack Growth under Non-Steady-State

Conditions-A.

SAXENA 185

An Application of Stress Intensity Factor to Fatigue Strength .
Analysis of Welded Invar Sheet for Cryogenic Use-I.
SOYA,
H. TAKASHIMA,AND Y. TANAKA

2{)2

An Automated Photomicroscopic System for Monitoring
of Small Fatigue Cracks-I.
M. LARSEN

226

the Growth
.

An Experimental and Numerical Investigation of the Growth and
Coalescence of Multiple Fatigue Cracks at
NotcheS-A. F. GRANDT,JR., A. B. THAKKER,AND
D. E. TRITSCH
Near- Tip Crack Displacement Measurements During

High-Temperature
Fatigue-w.
N. SHARPE,JR., AND 1.

1.

239

LEE

Viscoplastic Fatigue in a Superalloy at Elevated
Temperatures-R.
WILSONAND A. PALAZOTTO

253

265

FRACTURETESTING
Fracture
1.

Testing with Arc Bend Specimens-I.
A. KAPP, AND M. D. WITHERELL

JIe Testing Using Arc-Tension
W. J. BILINSKY

Specimens-I.


H. UNDERWOOD,
279
A. KAPPAND

Investigation and Application of the One-Point-Bend
Test-I.
H. GIOVANOLA
Mode

297
Impact
307

n Fatigue

Crack Growth Specimen
Development-R.
1. BUZZARD,B. GROSS, AND 1. E. SRAWLEY

Discussion

329

345


A Compact Mode II Fracture
M. ARCAN

Specimen-L.


BANKS-SILLSAND
347

Influence of Partial Unloadings Range on the JI-R Curves of ASTM
AI06 and 3-Ni Steels-G. E. SUTTONAND M. G. VASSILAROS

364

\

Fracture Toughness Testing of Zircaloy-2 Pressure Tube Material
with Radial Hydrides Using Direct-Current Potential
Drop-p.
H. DAVIESAND C. P. STEARNS

379

Assessment of J·R Curves Obtained from Precracked
Specimens-I.
A. KAPP AND M. I. JOLLES

401

Charpy

DUCTILE FRACTURE
A Single Specimen Determination of Elastic-Plastic Fracture
Resistance by Ultrasonic Method-K.
HIRANO, H. KOBAYASHI,

AND H. NAKAZAWA
415
J-Resistance Curve Analysis for ASTM AI06 SteelS-Inch-Diameter
Pipe and Compact Specimens-M.
G. VASSILAROS,R. A. HAYS,
AND I. P. GUDAS
435
Influence of Crack Depth on Resistance Curves for Three-Point Bend
Specimens in HY130-o.
L. TOWERSAND S. I. GARWOOD

454

An Investigation of the I and dJ/da Concepts for Ductile Tearing
InstabilitY-M.
R. ETEMADAND C. E. TURNER

485

Computation of Stable Crack Growth Using the
J-Integral-I.
E. CARIFO, I. L. SWEDLOW,AND C.-W. CHO

503

Evaluation of Environmentally Assisted Cracking of a High Strength
Steel Using Elastic-Plastic Fracture Mechanics
Techniques-E.
M. HACKETT,P. I. MORAN,AND I. P. GUDAS


512

Plastic Energy Dissipation as a Parameter to Characterize
Growth-T.
I. WATSONAND M. I. JOLLES

542

Crack

ANALYSISAND MECHANISMS
Stress Intensity Factors for a Circular Ring with Uniform Array of
Radial Cracks of Unequal Depth-s.
L. PU

559

Weight Functions of Radial Cracks Emanating
in a Plate-G.
T. SHAAND C.-T. YANG

573

from a Circular Hole


An Empirical Surface Crack Solution for Fatigue Propagation
Analysis of Notched Components-I.
F. YAU


601

Extension of Surface Cracks During Cyclic Loading-H.
S. MULLER, D. MUNZ, AND I. NEUMANN

625

M. MULLER,

Comparison of Predicted versus Experimental Stress for Initiation
Crack Growth in Specimens Containing Surface
Cracks-w.
G. REUTER

of
644

Comparison of Ductile Crack Growth Resistance of Austenitic,
Niobium-Stabilized
Austenitic, and Austeno-Ferritic Stainless
Steels-p.
BALLADONAND I. HERITIER

661

Minimum Time Criterion for Crack Instability in Structural
Materials-H.
HOMMA,D. A. SHOCKEY,AND S. HADA
Discussion


683
694

Dynamic Crack Propagation and Branching under Biaxiai
Loading-A.
SHUKLAAND s. ANAND

697

Assessing the Dominant Mechanism for Size Effects on CTOD
Values in the Ductile-to-Brittle Transition
Region-T.
L. ANDERSONAND S. WILLIAMS

715

Dynamic l-R Curve Testing of a High Strength Steel Using the Key
Curve and Multispecimen Techniques-J.
A. JOYCE
AND E. M. HACKETT
Discussion

741
773

Boundary Layer Effects in Cracked Bodies: An Engineering
Assessment-c.
W. SMITH, I. S. EPSTEIN, AND
O. OLAOSEBIKAN


775

Stress Intensity Factors for Circumferential Surface Cracks in Pipes
and Rods under Tension and Bending Loads-I.
s. RAJU
AND I. C. NEWMAN

789

Summary

809

Author Index

817

SubjectIndex

819


Introduction

,
This volume and the Seventeenth National ASTM Symposium on Fracture
Mechanics on which it is based are part of a continuing series. These symposia have become clearly the most prestigious in the field of fracture. As such,
they are the focus and forum for quality work in all areas of the field, and this
is the important purpose of the symposium and volume.
If the field can be divided into testing and analysis, the former has been

and continues to be the more emphasized in this symposium series. This is
l!Ppropriate, considering the sponsor, ASTM Committee E-24 on Fracture
Testing. Nevertheless, analysis is a required part of any test, and much of the
work reported here is primarily analysis.
At least four general topics or categories of work frequently occur in the
papers: ductile fracture, test method development, surface cracks an.d crack
shape effects, and high temperature and loading rate effects. The prevalence
of these four categories attests to the basic practical nature of the field of
fracture and of those who work in it. Each of these categories defines an area
of important current concern in the design and use of load-carrying components and structures. It is the hope and belief of all those involved that this
symposium and volume have contributed to these and other important areas
in the field of fracture.
The National Symposium on Fracture Mechanics is often the occasion at
which ASTM awards are presented to recognize the achievements of current
investigators. At the Seventeenth Symposium two awards were presented.
The ASTM Committee E-24 Irwin Medal was presented by Dr. Irwin to Mr.
John G. Merkle, Martin Marietta Energy Systems, for his outstanding work
in the field of fracture mechanics. The ASTM Award of Merit and honorary
title of Fellow were given to Mr. David P. Wilhem, Northrup Corporation, for
his distinguished service and leadership in Committee E-24. Dr. J. Gilbert
Kaufman, Arco Metals, past chairman of E-24, made the presentation to Mr.
Wilhem ..
We take this opportunity to thank two groups who deserve a significant
share of credit for this symposium. The first is the combined support staff of
all of us listed below. The administrative and clerical work of this whole group
was essential to the task and is greatly appreciated. The second group is made
up of those behind-the-scenes people whose work is nonetheless critical.
In particular, we thank Professor Ray Eisenstadt of Union College for his
help in administering the symposium, Mr. Jim Gallivan of the Army Materials and Mechanics Research Center for financial support, the late Dr. Fred
1



2

INTRODUCTION

Schmeideshoff of the Army Research Office for his help in organizing the
symposium, and Professor Jerry Swedlow for his continuing support and
sound advice during the entire process.

John H. Underwood

David P. Wi/hem

Army Armament Research & Development
Center, Watervliet, New York; chairman
and co-editor.

Northrop Aircraft, Hawthorne, California; cochairman and co-editor.

Richard Chait
Army Materials and Mechanics Research
Center, Watertown, Massachusetts; cochairman and co-editor.

. C. William Smith
Virginia Polytechnic Institute & State University, Blacksburg, Virginia; co-chairman
and co-editor.

Wayne R. Andrews
.


General Electric Company, Schenectady, New
York; co-chairman and co-editor.

James C. Newman
NA~A .L.angley R~search Center, .Hampton,.
Vlrgmla; co-chaIrman and co-editor.


P. D. Hilton, 1 R. A. Mayville, 1 and D. C. Peirce1

\

An Application of Fracture
Mechanics to a Ship Controllable
Pitch Propeller Crank Ring

REFERENCE: Hilton, P. D., Mayville, R. A., and Peirce, D. C., "An Application of
Fracture Mechanics to a Ship Controllable Pitch Propeller Crank Ring," Fracture Mechanics: Seventeenth Volume. ASTM STP 905. J. H. Underwood, R~ Chait, C. W.
Smith, D. P. Wilhem, W. A. Andrews, and J. C. Newman, Eds., American Society for
Testing and Materials, Philadelphia. 1986. pp. 5-21.
ABSTRACT: A fracture mechanics analysis was conducted to establish the fracture
toughness of a controllable pitch propeller crank ring material required to prevent a fracture mode in which loss of a propeller blade occurs. Loss of the propeller was assumed to
be prevented if fracture instability could not occur before the fatigue crack grew to a size
beyond which crack growth would proceed radially through the flange of the crank ring
and not around the circumference. The fracture analysis was conducted by modeling the
cracked crank ring as a plate with a part-through crack in bending. Numerical solutions
for part-through cracks in bending were combined with results for large crack length-toplate width geometries for through cracks in bending to determine K, for the large crack
size of interest. Values of K, with plastically adjusted crack lengths were converted to
values of J, and crack driving force curves were generated. Estimates of the plastic collapse moment for the crank ring were made as an alternative method of determining fracture conditions. The results of the analysis are a minimum acceptable value of yield

strength and curves of yield strength versus minimum acceptable values of J'R and T m.t at
a crack extension of 1.27 mm as determined by a J-R curve test.
KEY WORDS: fracture mechanics, application. bending. ship component

Controllable pitch propellers are commonly found in current ship propulsion systems. They are used for both small vessels and large ships with power
as great as 40 000 hp. All controllable pitch propellers require some mechanism to rotate the propeller blades. In the study described in this paper, rotation is brought about by a crank ring to which the propeller blade is attached
by several bolts. An illustration of such a crank ring is shown in Fig. 1. Not
'Arthur D. Little, Inc., Acorn Park, Cambridge, MA 02140.
5


6

FRACTURE

MECHANICS:

SEVENTEENTH

VOLUME

shown in the figure is the protrusion from the underside of the ripg to which a
mechanical "crank" is attached for the purpose of rotating the crank ring.
On installation, the crank ring is set over a central post which is attached to
the propeller hub. Next, a narrow bearing ring is threaded into the hub body
over the crank ring so that if the crank ring were lifted, its flange would contact the underside of the bearing ring. Finally, the propeller blade is bolted to
the crank ring. Thus, under the action of centrifugal and hydrodynamic loads
during operation of the propeller, significant pressure loads are transferred
between the underside of the bearing ring and the upper surface of the crank
ring flange. This in turn causes cyclic stresses in the fillet at the point where

the flange meets the main crank ring body (Fig. 1). These high stresses can
lead to cracking at the fillet [1], and there is a possibility that the crank ring
can fracture. In fact, one can imagine a scenario in which rapid fracture from
a fillet crack could proceed around the circumference of the crank ring and
lead to separation of the propeller blade from the hub body.
The objective of the investigation described in this paper was to establish
through analysis the material fracture toughness for a particular crank ring
such that, in the unlikely event that a fatigue crack does initiate, a fracture
mode leading to loss of the propeller would be avoided. Periodic inspection of
crank rings is generally not conducted, so that in this scenario some other
incident, such as excessive deformation, must oc'cur to make the failure detectable. There has been one reported failure incident in which a fillet fatigue
crack initially propagated around the circumference of the crank ring but
eventually propagated and broke through the flange. The severed piece of the
crank ring then prevented rotation on the next attempt at pitch control and
this led to the discovery of the fracture. It is not clear that fatigue cracks in all
crank rings will proceed in this manner and, in fact, results of our analysis,
presented below, show that there is a significant driving force for continued


HILTON ET AL ON PROPELLER CRANK RING

7

circumferential fatigue crack growth. Nevertheless, based on limited evidence, it has been assumed that the crack will propagate initially around the
circumference and then through the flange, provided the mode of fracture is
by fatigue and not by rapid brittle or ductile fracture. Furthermore, loss of
the propeller is assumed to be prevented i~fracture instability cannot occur
before the fatigue crack grows to a size beyond which crack growth by any
mode would proceed approximately radially through the flange and not
around the circumference.

The first problem in establishing the required crank ring toughness is to
choose a crack size and geometry from which fracture instability would proceed through the flange. Guaranteeing that fracture instability will not occur
prior to attaining this crack size is then equivalent to finding th~ conditionsmaterial toughness-required for instability to occur at this crack size; this
assumes that smaller crack sizes are less severe.•
Crank Ring and Crack Geometry

A cross section of the single crank ring geometry analyzed iii this investigation is shown in Fig. 2. A full-scale laboratory test was performed for this
crank ring resulting in a fatigue crack, the geometry of which was used in our
analyses and is shown in Fig. 3. The crack had several initiation sites located
in the fillet on the thrust side of the blade and at discovery extended about 85°
around the circumference. At its midpoint the crack was inclined approximately 45° to the vertical. The crack front extended part way into the flange
and to within about 8.9 mm of the crank ring bottom at the center of the


sumption of a rigid bearing ring. Two calculations are carried out: the first
prescribes a constant downward displacement on the outer circumference
(modeling the centrifugal load), while the second prescribes a set of displacements varying linearly in y (modeling the hydrodynamic load). These two solutions are superposed in such a way that the nodal reaction forces at the
outer edge of the flange balance the centrifugal load and hydrodynamic bending moment. These loads were obtained from the same example used in the
previous strain gage calculations.
The maximum load obtained by finite element~~nalysiswas approximately
2100 N/mm. This differs from the maximum load derived from the strain
gage readings by 17%. Both of the methods used to estimate the load distribution along the circumference of the crank ring are approximate, and it is
difficult to establish which of the estimates is more accurate. Since the finite
element analysis ensures that force and moment equilibrium are satisfied and
avoids the complications associated with load (or stress) concentrations, the
finite element analysis results are used to perform the fracture mechanics
analysis to be described later.


HI LTON ET AL ON PROPELLER


CRAN K RI NG

11

A finite element calculation was also performed to quantify the load redistribution which occurs in the presence of a crack. The mesh and loading were
exactly the same as for the uncracked case, except nodes on the inner flange
circumference were released to simulate a crack that extends 90° around the
circumference and 75% through the flange thickness. The center of the crack
was symmetric with respect to the y-axis.
The load distributions for the uncracked and cracked crank ring models;
as calculated by the finite element analysis, are shown in Fig. 5. There is a
substantial redistribution in load in the presence of a crack. In particular, the
load at the intersection of the y-axis with the flange periphery, which is the
location of maximum load in the un cracked crank ring, shows a reduction in
load to about 700 Nlmm from 2100 N/mm. Figure 5 also shows that the maximum load in the cracked crank ring occurs at about 50° from the y-axis.
This is in part due to the coarseness of the mesh used and the presence of the
crack tip at this point as well as the reduction in load with y which must occur
because of the bending nature of the problem. This large load near die crack
tip suggests that there may still be a significant driving force for circumferential crack growth ..



HILTON ET AL ON PROPELLER CRANK RING

13

This bending moment applies for one crack geometry. With continued
cracking-tearing-and
plastic deformation, the compliance of the cracked

flange will increase and the load will decrease. No attempt has been made to
quantify this decrease. Instead, the conservative assumption is made that the
cracked flange is subjected to a constant load.

Calculation of Crack-Driving Force
The fracture mechanics analysis for the cracked crank ring is carried out
using as a model a plate of finite width which contains a part-through crack
subjected to remote bending. This geometry and loading are illustrated in
Fig. 7. The model includes many of the important aspects of the cracked
crank ring configuration shown in Fig. 3: the fatigue crack does not completely penetrate the flange thickness and is shallower at its ends; benning
caused by the flange load appears to be the driving force for crack-growth;
and the crack is close enough to the outer flange edge to experience finite
width effects. Finite element analyses show that the remote bending model is
a good approximation even if the moment is produced by a vertical line load
applied close to the crack plane [3]. The bending moment used in the model
calculations is the moment per unit width across the cracked section resulting
from the flange load, as described in the previous section. The problem is
treated as quasi-static; dynamic effects on the crack driving force and fracture toughness are not included.
The finite element calculations for K[ by Newman and Raju [4] are used to
obtain the crack-driving force for the cracked crank ring. Newman and Raju
conducted analyses to determine K[ for a plate containing a part-through
crack in bending with the geometry shown in Fig. 7. Fracture in the cracked
flange under bending is considered to be most critical at or near the surface of
the flange at the ends of the crack. Therefore it is convenient to present


14

FRACTURE


MECHANICS:

SEVENTEENTH

VOLUME

results for K( at the surface in terms of the non dimensional parameter F defined by2

where O"b is the nominal outer fiber bending stress equal to 6m/ t2 and m is the
applied bending moment per unit width. The assumed crank ring crack is
characterized by the ratios ale = 0.16, alt = 0.82, and elW = 0.89 (Fig.
3). Newman and Raju's results closest to this case are for ale = 0.2, alt =
0.8 and c/W = 0.8, which give a value of F = 0.49. This value is certainly
low because F increases sharply as elW approaches unity and there is a considerable difference between elW = 0.8 and elW = 0.89.
Results by Boduroglu and Erdogan [5], who have recently published K(
solutions for plates of finite width containing through cracks and loaded in
remote bending, are used to quantify the effect of a greater crack length-towidth ratio for the plate containing a part-through crack. Results for the
through crack case are given for values of elW very close to unity. The approach in using these results is to assume that the effect of finite width for the
through crack geometry is the same as the finite width effect for the partthrough crack geometry. Table 1 lists the factors F for a number of elW
values for part-through cracks with ale = 0.2 and alt = 0.8 as obtained
from Ref 4 and for through cracks with elt = 4 [=(alt)/(ale)] as obtained
from Ref 5. The ratio of F-factors for the two cases is also listed.
The results in Table 1 indicate that for a/ e = 0.2 and a/ t = 0.8 the factor
Ffor the part-through crack geometry is approximately one half of the factor
for the through crack geometry. Therefore the approach taken to obtain a
value of F for the cracked crank ring geometry is to obtain a value of F from
the through crack analysis for a geometry close to the crank ring geometry
and to multiply it by 0.5. The value of F for a through crack geometry with
e/ W = 0.89 and eft = 5 is F = 1.6 [6] so that for the part-through crack
F = 0.5(1.6) = 0.8. This is the value used in the fracture mechanics analysis.



HILTON ET AL ON PROPELLER CRANK RING

15

Using the dimensions and loading for the cracked crank ring, m = 88.9
N-m/mm, t = 49.3 mm, and e = 246 mm, the value of KI at the free surface,
without correction for plasticity, is equal to 158 MPa.Jm. One notes immediately that, according to this analysis, a very tough material is needed to avoid
loss of the propeller in the presence of the assumed fatigue crack.
The fracture mechanics analysis of the crank ring will account for a certain amount of stable crack growth, so it is necessary to quantify the dependence of F on crack length. This is done by employing the dependence of F on
crack length for the through crack and multiplying by one half. The variation
of F with e for the crank ring crack dimensions, e/t = 5, Wit = [(elt)!
(elW)] = 5.64, is approximately'[6]
Ft

=

2.73e - 24.9

Multiplying this expression by one half provides the relation to be used in the
fracture mechanics assessment of the cracked crank ring:
.
Fpt

=

1.37 e - 12.4

(1)


Newman and Raju's results can also be used to estimate the stress intensity
factor at the bottom of the crank ring crack. For ale = 0.2 and alt = 0.8 the
stress intensity factor at the bottom of the crack is approximately one half of
the value at the surface for 0.2 < elW < 0.8; data for elW > 0.8 were not
given. Finite element analyses of an elliptical part-through crack in bending,
which model the close proximity of the vertical flange loads to the crack plane
and the short moment arm in comparison to the plate width and surface crack
length, also indicate that KI at the deepest point of the crack is about 50% of
KI at the surface [3]. This would appear to contradict the possibility of a
fatigue crack in the crank ring growing to the shape shown in Fig. 3. Without
attempting to explain this apparent contradiction, it is noted that crack propagation from the top part of the crack is the fracture that would lead to loss of
the propeller and is therefore of greatest interest in this analysis.
The large KI value calculated earlier for the crank ring crack shows that
tough materials must be used to avoid loss of the propeller according to this
methodology. This implies that the material will be at or near its upper shelf
behavior, that it can experience stable tearing, and that elastic-plastic fracture mechanics techniques are necessary to quantify its resistance to fracture
(at least to characterize the material toughnes~with small specimens). Consequently, crack driving force curves are calculated in terms of h since the
material's fracture resistance is expected to be expressed in terms of JI-R
curves.
The crack driving force curve is the relation between JI and crack length or
crack extension and is estimated from values of KI through the relation



HILTON ET AL ON PROPELLER CRANK RING

17

strength [7]. The idealized cross-sectional geometry shown in Fig. 9 was used

to perform this calculation. The neutral axis for this section is essentially at
the lower crack front, and the limit moment per unit width (2 W = 556 mm)
is given by
mlim

=

O.17ao

(4)

where the units for mlim and ao are N-m/mm and MPa. Therefore, for ao =
690 MPa, mlim = 117 N-m/mm, which is greater than the applied moment
assumed for the crank ring in this analysis: 88.9 N-m/mm. The actual collapse moment would be larger because the crack is probably smaller than the
idealization shown in Fig. 9, the material will harden, and the cracked geometry induces some constraint to plastic deformation.
Strength and Toughness Requirements for the Crank Ring

It is now possible to determine the strength and toughness of the crack ring
material required to prevent failure from occurring under tha assumptions of
this investigation. In the section on loads, it was determined that the effective
bending moment per unit width on the cracked section shown in Fig. 3 is
approximately equal to 88.9 N-m/mm.
A lower bound estimate of yield strength necessary to prevent collapse from
occurring can be calculated from Eq 4. In this case:


18

FRACTURE


MECHANICS:

SEVENTEENTH

VOLUME

ture and loading rates characteristic of the crank ring. This can be accomplished by requiring the minimum upper shelf temperature, say, as determined by Charpy tests, to be below the operating temperature.
In the first approach to specifying required toughness, no crack extension
by tearing is permitted in the engineering sense; that is,


The toughness data generated for the 4150H steel show that it is not in the
upper shelf at room temperature, which was the temperature for the full~scale
laboratory test. The Charpy energy at room temperature is quoted as ranging
from 8 to 15 J. The value of KI at fracture varied from 60 to 123 MPa.Jm. Two
of the tests provided valid KIc values, 60 and 82 MPa.Jm, in accordance-with
ASTM Test for Plane-Strain Fracture Toughness of Metallic Materials
(E 399); the other tests provided invalid values because either too much plasticity or crack growth occurred-it was not determined which-or the specimen dimensions did not satisfy the plane strain requirements. In any case, the
criterion proposed in the analysis of this paper, that the fracture mode be
ductile tearing, was violated.
The range of critical JI values converted from the Kc values is 16 to 68
kJ/m2, all of which are considerably lower than the required value of JI-R =
175 kJ/ m2• Therefore the methodology developed in this investigation predicts that the 4150H is not a suitable crank ring material. The fact that the
full-scale laboratory tested 4150H crack ring did not experience unstable
fracture shows that the analysis is conservative. The degree of conservatism
on the required JrR value in this case is greater than a factor of two.
This degree of conservatism arises because of the many assumptions made
in the analysis. Loads calculated using finite element analysis for an idealized
crack geometry are undoubtedly too high. Since JI is proportional to the load
squared, a decrease in load will cause a substantial decrease in required h


Summary and Conclusions

The objective of the investigation reported in this paper was to set material
toughness requirements to avoid loss of a ship propeller blade from a controllable pitch propeller crank ring that has a fatigue crack. Loss of the propeller
was assumed to be prevented if fracture instability could not occur before the
fatigue crack grew to a size beyond which crack growth would proceed radially through the crank ring flange and not around its circumference. Choice
of this crack size and geometry was based on a full-scale crank ring laboratory
test in which a large fatigue crack occurred.


Kiyoshi Tanaka, 1 Mitsuo Sato,1 Tadashi Ishikawa, 1 and
Hironori Takashima1

A New Wide Plate Arrest Test (SCA
Test) on Weld Joints of Steels for
Low Temperature Application
.

REFERENCE: Tanaka, K., Sato, M., Ishikawa, T., and Takashima; H., "A New Wide
Plate Anest Test (SCA Test) on Weld Joints of Steels for Low Temperature Application,"
Fracture Mechanics: Seventeenth Volume. ASTM STP 905. J. H. Underwood, R. Chait,
C. W. Smith, D. P. Wilhem, W. A. Andrews, andJ. C. Newman, Eds., American Society
for Testing and Materials, Philadelphia, 1986, pp. 22-40.
ABSTRACT: As a realistic and practical criterion for brittle fracture arrest in low temperature storage tanks, the present authors propose the short crack arrest (SCA) concept.
This aims at arresting brittle fracture before propagation to a catastrophe. In order to
investigate the capability of steel materials from the viewpoint of this concept, two types of
wide plate tests, the SCA tests on weld heat-affected zone (HAZ) and base metal, were
developed. These two test methods were designed to simulate the run-and-arrest phenomenon of brittle fracture in actual storage tanks.
The two types of tests, together with the compact crack arrest (CCA) test, were applied

to base metals and welded joints of a wide range of steels for low temperature application.
Efforts were made to clarify the effects of base metal chemical compositions and conditions for welding on the weld HAZ characteristics.
The results revealed that addition of about 3 % nickel to the base metal significantly
improved the weld HAZ arrest toughness. It was also shown that the SCA capability of
steels could be predicted from the Ka values obtained by the CCA test.
KEY WORDS: low temperature, fracture toughness, crack arrest, brittle fracture propagation, welded joint, weld heat-affected zone, storage tank

In Japan, in 1968, there was a catastrophic fracture accident in a spherical
tank with a diameter of 16.2 m during hydrotest [1]. The tank, made of an
800 MPa high tensile strength steel, collapsed after fast fracture in three
fourths of the circumference. The fracture surface consisted of shear fracture
in the base metal and brittle fracture along the heat-affected zone (HAZ) of a
'Senior Research Engineer, Research Engineer, Research Engineer, and Chief Research Engineer, respectively, R&D Laboratories II, Nippon Steel Corporation, Sagamihara, Japan.

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