[Elsevier] Review of Fatigue Assessment Procedures for Welded Aluminium Structures

7/26/2019 [Elsevier] Review of Fatigue Assessment Procedures for Welded Aluminium Structures

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International Journal of Fatigue 25 (2003) 13591378

www.elsevier.com/locate/ijfatigue

Review of fatigue assessment procedures for weldedaluminium structures

S.J. Maddox

TWI Ltd, Granta Park, Great Abington, Cambridge CB1 6AL, UK

Received 5 February 2003; accepted 24 February 2003

Abstract

This paper presents a review of methods and corresponding Codes and Standards for the fatigue assessment of welded aluminiumalloy structures. Methods for the fatigue evaluation of welded aluminium structures are assessed from the viewpoints of originaldesign and estimation of the residual life of existing structures. Based partly on a literature search, but also reference to data usedin the formulation of recent fatigue design Standards, it goes on to review the information available for such assessments in designor guidance specifications in the light of relevant fatigue data. With regard to design specifications, particular attention is focussedon recent fatigue data obtained from structural components representative of actual structures. Recommendations are made forfuture research. 2003 Elsevier Ltd. All rights reserved.

Keywords:Aluminium alloys; Cumulative damage; Design codes; Experimental data; Fatigue design; Fatigue crack growth; Fitness for purpose;

Stress analysis; Structural fatigue tests; Variable amplitude fatigue; Welded joints

1. Introduction

There is growing interest in the structural use of alu-minium alloys, for such applications as automotive andrailway vehicles, bridges, offshore structure topsides andhigh-speed ships. In all cases, welding is the primary

joining method and fatigue is a major design criterion.However, as is well known, welded joints can exhibitpoor fatigue properties. Thus, clear design guidelines areneeded to ensure that fatigue failures are avoided inwelded aluminium alloy structures. Apart from basicdesign of new structures, there is also increasing interestin methods for assessing the remaining fatigue lives ofexisting structures.

Prompted by difficulties experienced in reaching aconsensus on fatigue design rules, extensive testing andanalysis of the fatigue performance of welded aluminiumalloys have been undertaken over the past 20 years. Ameasure of the research effort is the series of Inter-

Corresponding author: Tel.: +44-1223-897762; fax: +44-1223-

892588.

E-mail address:[email protected] (S.J. Maddox).

0142-1123/$ - see front matter 2003 Elsevier Ltd. All rights reserved.

doi:10.1016/S0142-1123(03)00063-X

national Aluminium Conferences (INALCO), which hasproduced seven volumes of papers since 1981. Thefatigue research work culminated in the production ofnewdesign specifications, notably BS 8118[1], Euroc-ode 9[2],the International Institute of Welding (IIW)[3]recommendations and specifications from the AluminumAssociation[4]in the USA and the Canadian StandardsAssociation [5]. In relation to ships, DNV issued sup-plementary guidance [6] to their Rules for the Classi-fication of High Speed and Light Craft based on theECCS recommendations [7], the forerunner to Euroc-ode 9.

Thus, it is an opportune time to review the variousfatigue design procedures for welded aluminium alloystructures. This paper summarises methods for designand remaining-life assessments of fatigue-loaded alu-minium alloy welded structures and compares and con-trasts the information contained in the various recentfatigue design specifications. This includes assessmentof the corresponding design curves in the light of fatiguetest results, chiefly obtained in recent research pro-grammes where a major aim was to reproduce the fatigueperformance of full-scale welded structures, particularlywith respect to the effect of high tensile residual stresses.


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1360 S.J. Maddox / International Journal of Fatigue 25 (2003) 13591378

Nomenclature

A constant in equation for SNcurve

a crack sizeC constant in fatigue crack growth law

K stress concentration factorKm stress magnification factor due to misalignmentm slope ofSNcurven exponent in fatigue crack growth law

ni number of applied load cycles at stress rangeSi (i = 1, 2, 3,)

N fatigue life in cyclesNi fatigue life at stress range Sip exponent in plate thickness correctionS stress rangeSeq equivalent constant amplitude stress rangeY correction factor in formula for stress intensity factor

K stress intensity factor range

Finally, areas still requiring further research are ident-

ified.

2. Fatigue assessment methods

2.1. Background

Broadly speaking, any fatigue assessment involves

comparison of theactionswhich the component or struc-

ture under consideration will be required to sustain dur-

ing its design life (e.g. fatigue loading history, resultingstresses and the number of times they occur, anyenvironmental influences, etc.) with its resistance tofatigue. Clearly, the resistance must be sufficient to with-stand the actions without failure occurring. The form and

source of the resistance data depend on the type of

assessment being performed.

A useful summary of the four main methods for

assessing the fatigue lives of welded joints is contained

in the new IIW fatigue design recommendations [3].

They are:

(a) SN curves for specific welded joints used in con-junction with nominal stresses.

(b) SNcurves for welds used in conjunction with hot-

spot stresses.

(c) SN curves for materials used in conjunction withlocal notch stresses.

(d) The fracture mechanics approach, whereby fatigue

crack growth data are used in conjunction with the

stress intensity factor to calculate the progress of a

knownflaw.

The first three are intended for application at the designstage and are described in detail in Section 2.2. The

fourth is not generally used for design but for assessing

known or assumed flaws. Thus, it would be applicablein an assessment of residual fatigue life, as described in

Section 2.3.

2.2. Design of a new structure

2.2.1. Method

Fatigue resistance data for design are usually

expressed in terms of SN curves, relating nominal

applied cyclic stress rangeSand the corresponding num-

ber of cycles Nneeded to cause failure. In the simplestsituation, the designer would ensure that the number ofapplied load fluctuations, n, in the design life thatresulted in stress range Sdid not exceed N. In the more

general case of a detail, which will experience a spec-

trum of applied loads, the cumulative damage due to

individual load cycles would need to be determined. The

usual method is to apply Miners rule, which assumesthat the fatigue damage due to ni cycles of stress Si is

directly proportional to ni/Ni. An important step in the

assessment is estimation of the stress history that willbe experienced by the detail under consideration. In gen-

eral, this involves identification of the loading history,conversion from loads to stresses (e.g. by finite elementanalysis (FEA) or strain gauge measurements) and,

finally, extraction of recognisable stress cycles from thestress spectrum (the process of cycle counting) to pro-vide input to Miners rule. The full procedure is illus-trated in Fig. 1.

2.2.2. SN curves used with nominal stresses

The SNcurves used in fatigue design depend on theprocedure being used. Referring to those mentioned earl-

ier, by far the most common approach is to use SN

curves obtained from fatigue tests on specimens contain-


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1361S.J. Maddox / International Journal of Fatigue 25 (2003) 13591378

Fig. 1. (a) Miners linear cumulative damage rule for estimating fatigue lives under variable amplitude loading; (b) analysis of fatigue loadingfor cumulative damage calculations.

ing the weld detail of interest (Fig. 2). Such SNcurvesappear in many codes and standards, including some that

apply to welded aluminium alloys. The design curve is

usually some statistical lower bound to published experi-

mental data, typically mean2 standard deviations of

logN. Since the SN curves refer to particular weld

Fig. 2. Examples of designSNcurves for welded joints (from IIW

recommendations for aluminium [3]).

details, there is no need for the user to attempt to quan-tify the local stress concentration effect of the weld detail

itself. Thus, the curves are used in conjunction with the

nominal stress range near the detail. In codes and stan-

dards, the curves are identified by arbitrary letters or,increasingly, by the fatigue strength at a particular life,

usually 2 106 cycles. The current status of fatiguedesign rules for welded aluminium alloys is discussed

in more detail later.

2.2.3. Hot-spot stress approach

The hot-spot stress method is an extension of the SN curve approach in that it makes use of SN curvesobtained from tests on actual welded joints. However,theSNcurve is based on the hot-spot stress range ratherthan the nominal. Nominal stress is easy to define insimple laboratory specimens. However, in real structures

the presence of gross structural discontinuities, non-uni-

form stress distributions and through-thickness stressgradients can be so complex that the nominal stress is

no longer obvious. Experimental (e.g. strain gauges) and

numerical (e.g. FEA) stress analysis methods are capable


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of providing detailed information about the stresses aris-

ing near welded joints. In such circumstances, the struc-

tural stress, which includes the effect of all sources of

stress concentration except the weld itself, can be used.

The hot-spot stress, which is discussed in more detail byNiemi [8], is the structural stress at the weld toe. It is

usually necessary to determine it by extrapolation fromthe stress distribution approaching the weld (Fig. 3).

However, parametric formulae exist for calculating the

hot-spot stress in some tubular joints[9],and more suchformulae are likely to be developed as the hot-spot stress

method becomes more widely used. A practical limi-

tation is that the hot-spot stress method is only suitable

for assessing weld details from the point of view of

potential failure from the weld toe.

Apart from tubular joints, there are no established SNcurves for use with the hot-spot stress. The SNcurvesfor use with the nominal stress are not generally suitable

because they include some influence of the stress con-centration effect of the welded joint. Thus, for example,

the SN curve for a fillet welded cover plate is belowthat for a simple fillet welded stiffener because of thegreater stress concentration effect of the cover plate. An

obvious candidate for a hot-spot stress SNcurve is thatfor transverse butt welds, since there is essentially nostress concentration effect due to the joint (provided it

is perfectly aligned), only the weld bead. Indeed this is

the case for tubular joints.

2.2.4. Notch stress approach

While the notch stress approach applies only to assess-

ments of potential failure from the weld toe or root, themethod attempts to include all sources of stress concen-tration, including the weld itself, in the stress used with

the design SN curve. Thus, in principle, a single SN

curve is sufficient for a given type of material. A practi-cal problem is that the local geometry of the toe or root

Fig. 3. Stress distribution approaching a welded joint and the defi-

nition of the hot-spot stress.

of a weld is highly variable and, at the design stage, not

known. In recognition of this problem, the weld

geometry is normally idealised as having a particular

shape and weld toe or root radius. The local stress is

then calculated by numerical analysis. Alternatively,parametric formulae are available for a range of joint

geometries[10].Until very recently[6,11],the notch stress method did

not appear in any fatigue design specifications. Indeed,one of its protagonists [10]only recommends it for car-rying out comparative studies of the fatigue performance

of different welded joint options. Furthermore, the

method has not been developed to any extent for alu-

minium alloys. Consequently, it is not considered further

in this report.

2.3. Remaining life of existing structures

Broadly three approaches can be envisaged for thefatigue assessment of existing structures which have

already experienced some service duty. The approach

used will depend on the circumstances, particularly

whether or not the structure was designed for fatigueloading, the time in service and what measures will be

taken to assess its current condition with respect to

potential fatigue damage already introduced during pre-

vious service.

Three assessment methods are described sub-

sequently. Examples of their application or reference to

their development may be found in Refs. [1214].

2.3.1. Fatigue design assessmentThis method follows the procedure outlined in Section

2.2.1 for original design. If the structure was designed

for fatigue loading, the same actions can be assumed,after any modifications to allow for changes such asreduced severity of the stress history from reinforcement

or a change in the operating conditions. Fatigue resist-

ance is still represented by the design SN curves. Ifrepairs are introduced, the design curves may still be

applicable, but a safety factor could be introduced if the

repair was of uncertain quality. Post-weld improvement

of repair welds, for example by toe grinding (to be dis-

cussed in more detail later), would justify a higher SNcurve, which may be included in a design specificationor obtained from appropriate published information.

Finally, assuming Miners rule (see Section 3.3.5 regard-ing validity of this assumption), it is used to calculate

the fatigue damage introduced before and after the time

of the assessment, on the basis that

n/Nbefore

n/Nafter

1 .

Then, the remaining life is given by:

Remaining life (e.g. in years) (1)


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life so far (years) 1n/N

before

n/Nafter

per year .

2.3.2. Fatigue design review

The aim of a design review would be to improve theaccuracy of the original design process to provide a bet-

ter estimate of the proportions of fatigue life used andremaining at the time of the assessment. A fatigue design

process involves many assumptions and hence potential

inaccuracies. When assessing an existing structure, there

may be scope for improving the accuracy of some of

those assumptions. For example, records of the service

operation or even measurements made on the structure

may enable a more precise definition of the stress his-tory. Knowledge of the actual structural arrangement and

weld details used, including their quality (e.g.alignment), coupled with appropriate stress analysis may

allow the more precise hot-spot stress approach to be

used instead of the normal SN curves. A furtherrefinement might be the characterisation of actionsand/or resistance data in statistical terms to enable

reliability methods to be used to assess the risk associa-ted with a particular estimate of remaining life. Some

progress has been made in such an approach in the con-

text of steel bridges [15].

2.3.3. Fracture mechanics approach

The third method specifically addresses circumstancesin which it has been found, or it must be assumed, thatflaws (e.g. fatigue cracks) have been introduced duringthe service life endured so far. Such flaws would bethose detected or measured by non-destructive testing(NDT), or assumed flaws corresponding to the limit ofdetection of the NDT methods used.

A fracture mechanics assessment [16] utilises the

same actions as those determined for design calculations.

However, fatigue resistance is represented by fatigue

crack growth rate data for the material under consider-

ation, expressed in terms of the fracture mechanics stress

intensity factor parameter K. K is a function of

applied stress range (S), and crack size (a), such that:

K YSa (2)

where Yis a function of geometry and loading. The use

ofKensures that the relationship between crack growth

rate and Kcan be regarded as a law applicable to any

geometry of the same material. The crack growth law

approximates to a linear relationship (usually referred toas the Paris law):

da

dN C(K)n (3)

with deviations as K approaches a threshold value

below which crack growth is insignificant (Ko) and asKmax approaches the critical value for fracture, as illus-

trated inFig. 4.In practice, the gradual transitions from

the Paris law may be modelled more accurately bydefining several linear relationships. For a flaw size ai

and a critical fatigue crack size of af, the remainingfatigue life N under stress range S is obtained by inte-

grating Eq. (3):

af

ai

da

(YSa)n CN (4)

For variable amplitude loading the integration will be

performed for each individual cycle or block of equal

stress cycles, to give:

a1

ai

da

(YS1a)n

a2

a1

da

(YS2a)n % etc. (5)

CN

3. Fatigue design data

3.1. Design specifications

As noted earlier, there is a wide choice of fatigue

design specifications for welded aluminium alloys. Themain ones, in chronological order, are as follows:

BS 8118:1991. Structural use of aluminiumPart 1Code of practice for design, BSI, London 1991.

Fig. 4. Fracture mechanics fatigue crack growth relationship.


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ECCS. European recommendations for aluminium alloy

structures, fatigue design, European Convention for

Constructional Steelwork, Document No. 68, 1992.

Canadian Standards Association CAN/CSA-S157-M92.

Strength design in aluminium, 1993.The Aluminum Association. Specifications for alu-

minium structures, Washington, DC, 1994.DNV. Class note: fatigue assessment of aluminium

structures, Technical Report No. LIB-J-000010, 1995.

International Institute of Welding. Fatigue design ofwelded joints and components, Abington Publishing,

1996.

Eurocode 9. Design of aluminium structures. Part 2.

Structures susceptible to fatigue, ENV, 1999-2: 1998,

CEN, Brussels, 1998.

Apart from the DNV document, these all provide a

selection of design SN curves expressed in terms of

nominal stress ranges. In the DNV note, attention is con-

fined to the use of the hot-spot stress range, a methodthat is also referred to in the IIW recommendations and

Eurocode 9, but specific design data are not provided.There are significant differences between the SNcurvesin the rules and how they are used, and hence the differ-

ent specifications will lead to different estimates offatigue life. In order to provide a basis for judging their

applicability to welded aluminium structures, key fea-

tures are compared and where possible assessed in the

light of relevant published data.

3.2. Historical developments

In order to appreciate why so many fatigue designspecifications have been produced in recent years, it isuseful to review the developments over the past 20 years,

mainly in Europe, which have influenced them.In 1979, it was decided that the British Standard

design specification for aluminium, CP 118, which wasthe most comprehensive standard for aluminium at the

time and used throughout the world, should be revised

as a limit state Standard. This followed the publication

of new steel bridge design rules, BS 5400, on the samebasis. With regard to fatigue, as a starting point the

possibility was examined that the new rules for steels

could be simply factored in accordance with the differ-

ence in Youngs modulus between steel andaluminiumto provide the aluminium fatigue rules [17]. This

approach stemmed largely from the good correlationbetween fatigue crack growth data for the two materials

on the basis ofK/Eand the assumption that the fatigue

lives of welded joints are dominated by fatigue crack

growth[18].Thus, the fatigue design stresses for welded

aluminium alloys would be obtained simply by dividingthose for steel by 3. A review of published data tended

to support this approach and was adopted for the Draft

for Public Comment of the Standard that would replace

BS CP 118, BS 8118, in 1985. However, meanwhile

some industries had drawn attention to the fact that the

new rules were considerably more conservative than

those in BS CP 118 in the high-cycle regime, while

others felt that the steel/3 approach was too simplisticand effectively penalised aluminium alloys as compared

with steel.The initial review of fatigue data for welded joints in

aluminium alloys had drawn attention to the wide scatter

in published data and the fact that most data referred tosmall-scale specimens of variable, unspecified quality.Another important characteristic of small-scale speci-

mens, particularly those incorporating transverse welds,

is that they will contain much lower tensile residual

stresses than would be expected to be present in a real

structure. It was felt that more realistic fatigue data rel-evant to actual welded structures were required. These

arguments influence the newly formed ECCS Committeecharged with the task of drafting a European Standard.

Consequently, they placed particular reliance on data

obtained from realistic structural specimens, mainly

beams. A large database was available from one source

(Alusuisse) and this was made available to the Commit-

tee. In addition, a number of new European projects pro-

vided additional data that were taken into consideration.To some extent, the same results were used to review

the BS 8118 Draft for Public Comment and the fatigue

rules were revised slightly as a result.

The resulting ECCS[7] and BS 8118[1] fatigue rules

were finallyconsidered together as the basis of the newEurocode 9[2]in the early 1990s. Even more large-scale

specimen data were available by then and so the finalform of Eurocode 9 is different from both BS 8118 and

the ECCS specification.Other significant developments were the drafting of

fatigue design rules in the USA[4]and Canada[5],bothof which are knownto have been influenced by the Euro-pean activities [19].

3.3. Summary of design rules

3.3.1. Design SN curves

All the fatigue design specifications for welded alu-minium alloys present a series ofSNcurves for parti-cular weld details, with a classification scheme linkinga description of the welded joint with the appropriate

design curve. Examples of the SNcurves provided areshown in Fig. 5. The classification usually depends onthe joint type, geometry, loading direction and mode of

fatigue failure, as illustrated for one group of joint types

in Eurocode 9 inFig. 2.Rather less comprehensive guid-

ance is given by the Aluminum Association [4], whichonly refers to joint types and loading direction. The SNcurves are derived from linear regression analysis of

logSversus logNfatigue data to establish mean curves


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Fig. 5. Examples of fatigue designSN curves for aluminium alloys

in recent specifications: (x) IIW [3]; (b) Aluminum Association [4];

(c) Eurocode 9[7].

and statistical lower bound, usually mean 2 standarddeviations of logN. The SNcurves have the form:

SmN A (6)

where A and m are constants. The curves are assumed

to extend up to stress levels corresponding to the static

design limit for the material, and down to a fatigue

endurance limit. A constant amplitude fatigue endurance

limit is introduced at an endurance of 5 106 cycles in

all the aluminium specifications. Two basic approacheshave been used to define the design SNcurves:

(a) An arbitrary grid of SN curves, usually equally

spaced on loglog scales, is defined and the curveclosest to the selected lower bound to experimentaldata of a particular detail is allocated to that detail.

The IIW recommendations [3] are based on thisapproach with the SN curves defined in terms ofthe stress range in N/mm2 at 2 106 cycles (see

Fig. 5(a)).

(b) The design SN curves are derived directly fromexperimental data. In some cases, the fatigue lives

of different details are so similar that the experi-mental data can be combined to produce a single

class for all of them. The resulting SN curves,which are not usually equally spaced, may be

described in terms of the fatigue strength at 2 106 cycles, and possibly the slope m of the SNcurve as in draft Eurocode 9[2],or by arbitrary let-

ters such as Class A, B, C, etc., in the Aluminum

Associations specification [4] (see Fig. 5(b) and(c)).

In most cases, the SN curves get progressivelysteeper as the fatigue strength of the detail decreases.

The steepest curve usually has a slope which is consist-

ent with crack growth data (i.e. m = n, typically 34),reflecting the fact that the lives of the low fatiguestrength details are dominated by crack growth [17,18].

Apart from the implied need to utilise mechanised

welding to achieve continuous welds without stop/starts,

no distinction is drawn between different welding pro-

cesses. The bulk of the test data upon which the design

curves are based have been obtained from arc welds.

However, as other processes became more viable forwelding aluminium, notably friction-stir welding [20]

(see later) which seems to offer advantages from the

fatigue viewpoint as well as production, it may become

necessary to introduce new process-related categories.

3.3.2. Effect of residual stress and mean stress

Welding introduces tensile residual stresses, which

modify the mean stress experienced by the welded joint

under fatigue loading. Long-range, or reaction, residual

stresses will also be introduced when welded sub-

assemblies are connected together, due to imperfect fit-up. It is generally assumed that tensile residual stresses

up to the proof strength of the material will be present

in a welded structure. As a result, its fatigue life will be

independent of mean stress and depend only on the

applied stress range, even if this is compressive [21].Consequently, all the fatigue design specifications arebased on the use of full stress range regardless of

whether it is tensile or compressive.


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3.3.3. Material

A common feature of all the specifications is that nodistinction is drawn between different aluminium alloys

when welded, unless they are exposed to a corrosive

environment. This reflects the fact that fatigue crackgrowth rates are not significantly different in different

alloys[22]and that fatigue crack growth dominates thefatigue lives of welded joints. Consequently, precise

details of alloys used to produce welded test specimens

discussed later are only given if they are significant.For unwelded material, some of the specifications

[2,3] provide higher design stresses for high strength

7000 series aluminium alloys as compared with all the

other alloy types.

3.3.4. Effect of plate thickness

It is generally acknowledged that the fatigue strengths

of welded joints failing from the weldtoe can decrease

with increase in plate thickness [23]. This has led to

thickness effect penalties, applied to the fatigue strength

obtained from the SN curve, in many fatigue designrules for welded steel of the form (tref/t)

p, where t is thethickness, tref the reference thickness (usually around 25

mm) and the exponent p = 0.25. Recent work showed

that the thickness effect also depended on the overallproportions of the welded joint[24,25].These influencesare incorporated in the fatigue rules in Eurocode 9. A

further refinement in the IIW recommendations[3] is tomodify the thickness correction exponent p for different

weld details. Values range from 0.3 to 0.1, reflecting thefact that the thickness correction also depends on the

level of stress concentration introduced by the weldedjoint. In contrast, the Aluminum Association take theview that the database used to establish the design SNcurves covered the full range of thicknesses of alu-

minium alloy likely to be used in practice [19]. Hence,

there is no requirement to apply a thickness effect cor-

rection. This assumption may be reasonable for some

applications (e.g. automotive or railway vehicles where

plate thickness is unlikely to exceed 25 mm) but not for

large structures such as bridges or LNG tankers where

plate thickness may be 100 mm or more.

3.3.5. Cumulative damage

Miners rule is universally adopted as the method forpredicting fatigue lives under variable amplitude loading

using the constant amplitude design SNcurves. How-ever, the accuracy of the rule has been called into ques-tion in recent years as more and more fatigue tests

obtained under random loading conditions have pro-

duced failures in shorter lives than those predicted by

Miners rule[26,27].It is thought that part of the reasonfor this is that the crack closure conditions for a givenstress fluctuation are different under constant and vari-able amplitude loading, with the result that a stress range

may be more damaging in a variable amplitude sequence

than it was under constant amplitude loading[27].How-

ever, a second problem concerns the damaging effect of

stress ranges below the constant amplitude fatigue limit.

Some specifications take account of such stresses byassuming that the SNcurve extends below the constantamplitude fatigue limit at a shallower slope. For an S

N curve of the form SmN = A, the extrapolated curvewould be of the form Sm + 2N= A, (seeFig. 5(c)). How-

ever, on the basis of fatigue tests on large-scale welded

beams (to be discussed later), the Aluminum Association[4]take the view that the SNcurve should be extrapo-lated indefinitely below the constant amplitude fatiguelimit without a slope change. The extent to which these

modifications to the SN curve are successful will beconsidered later in the light of new experimental data.

3.3.6. Hot-spot stress approach

Only the DNV note[6] gives specific guidance on the

use of the hot-spot stress fatigue design procedure. Thatguidance is related to four SN curves from the ECCSrecommendations [7], one for unwelded material, two

for welded connections and the fourth for welds exposed

to a corrosive environment (presumably seawater). The

corresponding ECCS design curves are as follows:

DNV class Material ECCS SNcurve

I Unwelded Unwelded high

strength 7020

alloy

II Welded Flush ground buttwelds

III Welded As-weldedtransverse butt

welds with good

profileIV Welded, in Not included:

corrosive 25% reduction in

environment design stress from

Class III curve

Thus, for as-welded joints it is assumed that the butt

weld design SN curve is applicable to both butt and

fillet welds if the hot-spot stress is used. The reason forthe choice of the SN curve for the highest strength ofunwelded aluminium alloy (lower curves were provided

for other aluminium alloys) is not known. The corre-

spondingSNcurves in Eurocode 9 are lower than those

in the ECCS rules.

The curves are used in conjunction with specifiedstress concentration factors, K, by which stress ranges

obtained from the specified design curve are divided, fora variety of typical welded connections used in ships.


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The range is not totally comprehensive and, as will be

seen later, no guidance is offered for some relevant

details.

3.3.7. Effect of environment

As already mentioned, the DNV note provides guid-

ance on the effect of immersion in a marine environmentby adopting an SNcurve which is 25% on stress below

the lowest curve for welded joints. That same curve

would be used for any detail, welded or unwelded. Thereduction in allowable stresses for as-welded joints of

25% is in-line with the results of an extensive series of

corrosion fatigue tests conducted in Norway in the early

1980s[28].

Eurocode 9 also provides guidance on the influenceof environment, industrial and marine. The basicapproach is to reduce the detail classification, by up tothree categories in the case of immersion in seawater,

and to reduce the fatigue endurance limit, as illustrated

in Fig. 6. The design penalty is most severe for 7000

series alloys, which are also susceptible to stress cor-

rosion cracking, while no reduction in design category

is required for 3000 series and aluminiummagnesium5000 series alloys, although the fatigue endurance limit

is still reduced. In general, the extent of the reductionin fatigue strength due to environment depends on the

detail and endurance since the SNcurves are not paral-

lel. It is not clear how the category-reduction approach

should be applied to the lower category details, for

which the required reduction would take them below the

design categories provided.

4. Comparison of design proposals and recent

fatigue data

4.1. Background

All the recent design SNcurves for welded joints in

aluminium alloys are claimed to have been derived from

Fig. 6. Eurocode 9[7] allowance for reduction in fatigue strength due to marine corrosion.

experimental data by linear regression analysis such that

they represent approximately 97.7% probability of sur-

vival [19,2931]. However, it is not always clear howthis has been achieved. It is evident that in most cases

some judgement has been applied and even someassumptions, like the slope of the design SN curve,

have been imposed. It is also claimed that special atten-tion was paid to the provision of data relevant to real

structures, particularly with respect to the influence oftensile residual stresses. Thus, wherever possible themain basis of the design curves has been experimental

results obtained from full-scale welded specimens, usu-

ally beams, or from specimens tested under high tensile

mean stress conditions to simulate the effect of high ten-

sile residual stresses. In view of this, the design SN

curves should be consistent with such data, includingdata generated since the curves were published. In order

to provide a basis for judging the validity of the pro-

posed design curves, relevant published data have been

assembled and they are presented in comparison with

some of the design curves. In this exercise, whenever

possible, attention has been focused on test results

obtained from specimens made from material 816 mmin thickness, since plate thickness is known to influencefatigue performance. Furthermore, only design curvesfrom the four most recent specifications, namely theAluminum Association specification, Eurocode 9, theIIW recommendations and the DNV design note, are

considered. The majority of the experimental data were

obtained under constant amplitude loading and presented

in terms of nominal stress range. Hence, these will be

used to assess the design SN curves intended for usewith nominal stress range. Many of the results for beamswere obtained from the compilation of data in Ref. [30],

which does not always give full details of the source

(which may have been an internal company report). Ref-

erence is made to the relevant series in that reference.

Limited data presented in terms of hot-spot stress

range are also available and these will be used to assess

the proposed hot-spot stress SNcurves. Finally, some


10/20


data have been obtained under variable amplitude load-

ing. These will be used to assess the validity of Minersrule and the methods proposed to modify the SNcurve

below the constant amplitude fatigue limit to take

account of the damaging effect of low stresses. The vari-able amplitude fatigue data will be considered in terms

of the equivalent constant amplitude stress range, calcu-lated on the basis that Miners rule is correct using theconstant amplitude SNdata obtained in the same inves-

tigation. This equivalent stress range is as follows:

Seq mSmi n ini (7)where ni cycles were applied at stress range Si before

failure, ni the total number of cycles to failure and m

is the slope of the SN curve. In those cases where a

significant number of applied stress cycles were belowthe constant amplitude fatigue limit, the SNcurve wasassumed to be extrapolated below this limit at a slope

ofm + 2, as proposed in most of the specifications.

4.2. Continuous longitudinal welds

The stress concentrations introduced by continuous

longitudinal butt and fillet welds, weld ripples or lumpsif stop/starts are present, are relatively minor. Their cor-

responding fatigue performance is relatively good. The

severity of a stop/start would be intensified if cratercracking occurred, which is certainly a possibility when

welding aluminium alloys, but these would normally berepaired if found.

In spite of the relatively good fatigue performance of

continuous longitudinally loaded welds, they are

important details, particularly in welded aluminium alloy

structures. Such details may be the governing ones in

well designed structures in which poorer transverse

welds have been avoided or located in low stress regions.

The potential for doing this is enhanced in the case of

aluminium alloys because of the enormous scope for

producing special extrusions, for example of sufficientrigidity in relevant parts to avoid the need for welded

stiffeners.

Continuous longitudinal welds are not explicitly

included in the DNV note. Of the other specifications,both the IIW and Eurocode 9 distinguish between welds

with and without stop/starts. The IIW recommendationsalso draw a distinction between butt and fillet welds. TheAluminum Association gives only one design category

for both butt andfillet welds with and without stop/starts.Recent fatigue data obtained from structural compo-

nents [30 (series C2, D1 and D2), 32,33] are confinedto I-section beams in which the test detail is either a

continuous butt weld in the web or, in the case of fabri-

cated beams, the web to flange weld. Data are available

for 5000, 6000 and 7000 series alloys in thicknesses

from 6 to 15 mm. They are shown together with relevant

design SN curves in Fig. 7(a) for welds without

stop/starts andFig. 7(b)for welds with stop/starts.

Referring first to the results for welds withoutstop/starts(Fig. 7(a)), it will be seen that they are most

consistent with the slope of the IIW design curves (i.e.m = 3), although only the Category 40 design curve is

safe for all the results. The Eurocode 9 and Aluminum

Association design curves appear to be too shallow. Thismay be a situation in which the slope of the design curve

has been imposed. Both the Eurocode 9 and Aluminum

Association specifications provide SN curves whichbecome gradually steeper as the fatigue strength of the

detail decreases and the adoption of a slope of m = 3

for such a high fatigue performance detail would intro-duce an anomaly into such a scheme. However, com-

pared with actual data, the result is that both the Euroc-

ode 9 and Aluminum Association design curves are

particularly conservative in the high stress/low fatigue

life regime.

The case for a shallow SNcurve is better for welds

containing stop/starts positions, as seen inFig. 7(b).This

Fig. 7. Comparison of fatigue test results obtained from (a) continu-

ous longitudinal butt and fillet welds without stop/starts and design

curves and (b) continuous longitudinal fillet welds containing

stop/starts and design curves.


11/20


situation arises mainly because of the wide scatter asso-

ciated with thefillet weld results. There are some fatiguedata below all the design curves, the IIW curve being

particularly non-conservative and the Eurocode 9 curve

being the most suitable. Further investigation of the rea-son for the low [29] results would be worthwhile. By

and large, it may be noted that the database does notprovide a strong indication that distinction should be

drawn between butt and fillet welds in design specifi-cations, but the results do support the distinction betweenwelds with and without stop/starts.

4.3. Transverse butt welds

This section is concerned with transverse butt welds

made from one or both sides, with the condition thatthey should be full penetration welds. A number of fac-

tors will influence the fatigue performance of transversebutt welds and some of them influence the design curves.In particular, a distinction may be drawn between welds

made from one or both sides and welds with different

profiles (expressed in terms of the weld toe angle).Further conditions might be that the weld should be

proved free from significant defects (i.e. those whichmight replace the weld toe as the site for crack initiationand lead to a lower fatigue life) by appropriate inspec-

tion, and that the effect of misalignment as a source of

secondary bending stress should be taken into consider-

ation when calculating the stress experienced by the

weld.

A reasonable database from structural specimens con-

taining transverse butt welds is available, mainly fromI-section beams [30 (series B7, B8, B10, B11), 33,34].These include specimens fabricated or extruded from

5000, 6000 and 7000 series alloys in thicknesses ranging

from 8 to 15 mm. In some cases, the weld toe angle is

reported. The data are shown in comparison with rel-

evant design curves in Fig. 8.

There is some indication of an influence of weld toeangle in that the highest results were obtained from

welds with a toe angle not exceeding 30, while the low-

est were from welds with angles up to 60. However,some good profile welds also gave lives near the lowerbound and overall the results do not indicate a strong

correlation between weld angle and fatigue strength.

Similarly, the results do not provide support for dis-

tinguishing between one- and two-sided full-pen-

etration welds.The DNV note distinguishes between one- and two-

sided welds and welds with different profiles. Defaultstress concentration factor values (by which stresses

obtained from the DNV III curve are divided) of K =

1.7 and 1.3, respectively, are given for weld toe anglesup to 50. The IIW recommendation distinguishes

between one- and two-sided welds and different weld

profiles, Eurocode 9 only distinguishes between one- and

two-sided welds, but the Aluminum Association pro-

vides only one design curve for any transverse butt weld.

Referring to Fig. 8, it will be seen that the Aluminum

Association Category C curve and the DNV curve for

welds made from both sides are very similar and providereasonable lower bounds to the data. The Eurocode 9

curves are unaccountably low while the IIW curvesappear to be too steep and, apart from FAT 28, too high.

However, it is interesting to note that regression analysis

of all the experimental data together results in a meanSN curve with slope of m = 2.95, very similar to the

assumed slope ofm = 3 in the IIW recommendations.

4.4. Transverse butt welds made on permanent

backing

One technique for ensuring full penetration for butt

welds from one side only is to use a permanent backing

bar or, in the case of aluminium alloys, backing lip

included in the extrusion. For joints in steel, the fatiguestrength of the resulting joint is lower than that obtained

from butt welds made from both sides, due to the severestress concentration introducedat the weld root between

the main plate and backing bar [21].Since this is a geo-

metric effect, it would be expected that the same would

be found from aluminium alloys. To some extent this is

the case, but the database is surprisingly limited in view

of the potential for extruding aluminium sections

incorporating backing lips. In fact, only one reference to

tests on structural components [35] could be found. In

view of this, data obtained from specimens are also con-sidered. The data found in the literature search are given

in Fig. 9 together with the appropriate design SN

curves. These refer to plate specimens in 6005 and 7020

alloys [30 (series B4)], extruded bridge deck panels in

6005 alloy and specimens extracted from such panels

[35].In fact, these specimens were reported to be sever-ely misaligned (angular distortion) with the result that

secondary bending occurred at the weld. The corre-

sponding stress magnification factor Km was estimatedby the authors and the results are presented in terms of

Km nominal stress range in Fig. 9.

It will be seen that most of the data lie above the DNV

and Aluminum Association design curves, which are

shallower than the IIW and Eurocode 9 curves. The datatend to follow the slope of the shallower curves, but with

such a limited database confined to a very limited rangeof relatively low endurances this may be a misleadingimpression. Certainly, in the light of experience of joints

in steel, the slopes of the IIW and Eurocode design

curves seem to be more appropriate, but further experi-

mental data are needed to confirm this.

4.5. Transverse cruciform joints

Fatigue data are available for I-section beams incorpo-

rating cruciform joints [30 (series F1), 33] in 15 mm


12/20


Fig. 8. Comparison of fatigue test results obtained from transverse butt welded 5000, 6000 and 7000 series aluminium alloy beams and design

curves.

Fig. 9. Comparison of fatigue test results obtained from transverse

butt welds made on permanent backing and design curves.

thick5083 alloy and 8 mm thick 6005 aluminium alloy

[34].In all cases, fillet and full penetration joints, failurewas by fatigue cracking from the weld toe. Also, some

data have been obtained from a model of a structural

connection used in an aluminium alloy ship [36]. Thiswas essentially the joint between the hull, T-section

longitudinal stiffeners and a transverse bulkhead (see

Fig. 10). The specimens were made in 6.4 mm thick

5086 H116 alloy. In all cases, the fillet weld sizes weresufficient to avoid failure in the weld throat in preferenceto failure from the weld toe.

All the results are plotted in Fig. 11 in comparison

with the appropriate design SNcurves. The detail is not

explicitly covered in the Aluminum Association specifi-cation but it has been assumed that the design curve for

transverse fillet welded stiffeners is appropriate. TheEurocode 9 design curve depends on plate thickness and

joint proportions and the curve shown is applicable to

the sizes of specimens used to generate the data

presented. As will be seen, the data are widely scattered.

However, both the IIW and Eurocode 9 curves appear

to be representative of the slope of the data and close to

the lower bound. The DNV and Aluminum Association

curves are similar, but both seem to be too shallow.

4.6. Transverse fillet welded attachments and

stiffeners

Transverse non-load carrying fillet welded attach-ments and stiffeners are very common in actual struc-

tures. Like transverse butt welds, small-scale specimensare unlikely to contain high tensile residual stresses andhence be representative of real structures from this view-

point. Therefore, fatigue test results obtained from struc-

tural specimens are particularly valuable.

A reasonable database now exists, as shown in Fig.

12. Most of the results were obtained from beams with

full or partial depth web stiffeners in 1115 mm thick5000, 6000 and 7000 series alloys [30 (series E1), 37].

In all cases, fatigue failure was from the weld toe in the

flange. In addition, a few results were obtained from I-section beams in12 mm thick 6061-T6[34] or 15 mm

thick 7020 alloys[30 (series E8)]with simple transverse

attachments on the tension flange.Fig. 12also includes the relevant design curves from

the four specifications being considered. As will be seen,the data strongly support the slopes of the Eurocode 9and Aluminum Association SN curves, between m =

3.2 and 3.6, and indeed those design curves are close to

the lower bound to the data. The DNV curve appears to

be too shallow, with a result that it is unduly conserva-

tive in the short life regime. The IIW curve, on the otherhand, seems to be too steep and over-conservative in the

long life regime. However, introducing a set of results

[38] obtained from small-scale specimens makes the


13/20


Fig. 10. Structural detail representing intersection of hull, longitudinal stiffener and transverse bulkhead fatigue tested by Beach et al.[36].

Fig. 11. Comparison of fatigue test results obtained from structural

specimens incorporating cruciform joints and design curves.

slope of the IIW curve seem more reasonable. Thesespecimens were tested with the maximum stress heldconstant at a value close to proof strength, to simulate

the presence of high tensile residual stresses. As seen,

they are rather similar to the results for beams with sim-

ple attachments. Clearly, more data for the long life

regime are needed to clarify the slope issue.

4.7. Longitudinal non-load carrying fillet welded

attachments

Specimens incorporating longitudinal non-load carry-

ing fillet welded attachments offer the advantage that

Fig. 12. Comparison of fatigue test results obtained from beams with

transverse fillet-welded attachments or web stiffeners, or plates with

transverse fillet-welded attachments, and design curves.

high tensile residual stresses exist even in small-scale

specimens[18].The detail itself is not particularly com-

mon in real structures, except perhaps as gussets tostiffen corners. Fatigue failure occurs by crack growth

from the weld toe at the end of the attachment, the

fatigue life being similar whether or not the weld is con-

tinued around the ends of the attachment. However, the

fatigue life varies with attachment size, the stress con-centration effect at the end of the attachment increasing

with attachment length.

Fatigue data are available for beams, extruded and


14/20


fabricated, with attachments fillet welded to the tensionflange [30 (series E6, E7), 34,39]. These have beenobtained using specimens made from 5000, 6000 and

7000 [34] series alloys in thicknesses between 11 and

15 mm. Some of these were tested under variable ampli-tude loading[39].The loading spectrum used was based

on strain measurements on a railway freight wagon.Finally, partly to extend the range of endurances, but

also because they include further fatigue test results

obtained under variable amplitude loading, recent resultsobtained from 10 mm thick 7019 alloy specimens can

be considered[38].

All these data are plotted in Fig. 13, together with

the appropriate design SN curves, including the curves

extrapolated below the constant amplitude fatigue limit

for use when performing cumulative damage calcu-lations. As will be seen, the SNcurves are rather similar

above the constant amplitude fatigue limit, with slopes

that are consistent with the experimental data. The

design curves also lie close to the lower bound to the

data. Thus, any of the design curves could be supported

by the database.

Comparing the data in Fig. 13 with those in Fig. 7,

which are representative of beams without stiffeners or

other attachments, provides a striking illustration of thedetrimental effect of attachments on fatigue perform-

ance. In practice, they should be avoided in highly

fatigue-loaded areas.

The variable amplitude fatigue test results for both

beams and specimens are consistent with the respective

constant amplitude data, supporting the validity of

Miners rule for the spectra used. Furthermore, the beamdata extend well below the constant amplitude fatiguelimit and hence provide a useful check on the validity

of the extrapolated SNcurve. Again, the data could be

used to support any of the proposals, but the fact that

they appear to be consistent with the same SN curve

as the constant amplitude data provides support for the

Aluminum Associations approach for extrapolating the

Fig. 13. Comparison of fatigue test results obtained from beams and

plates with longitudinal fillet-welded attachments and design curves.

SN curve without any slope change. More results in

the high-cycle regime are needed to check this further,

particularly for other loading spectra.

It will be noted that the results obtained from small-

scale specimens are entirely consistent with thoseobtained from beams, confirming that the two types of

specimen incorporated similar high tensile residualstresses. This adds confidence to the use of small-scalespecimens for investigating the fatigue performance of

this particular detail.

4.8. Beams with cover plates

Large cover plates welded to beam flanges representextremely high stress concentrations and consequently

result in the lowest fatigue performance for welds failingfrom the weld toe. Consequently, beams with cover

plates have been widely investigated for providing

design data. This has resulted in a large database for

beams in aluminium alloys, chiefly from fabricatedbeams in 1015 mm thick 5000, 6000 and 7000 seriesalloys [30 (series F3), 37]. A large number of results

have also been obtained from smaller beams [40,41], in

4 mm thick 6261-T6 aluminiumalloy. All these results

are plotted together in Fig. 14, along with the relevantdesign SN curves from the four specifications con-sidered.

It will be seen that the IIW and Eurocode 9 design

curves are very similar and consistent with the database

in terms of slope and position. The Aluminum Associ-

ation curve is slightly lower but of similar slope, while

the DNV curve appears to be too shallow with a resultthat it is excessively conservative in the short liferegime.

This detail is one in which the thickness effect would

be expected to apply, and indeed is evident from a com-

parison of the results obtained from 4 mm thick speci-

mens with the remainder. This thickness effect is incor-

porated in both the IIW recommendations and Eurocode

Fig. 14. Comparison of fatigue test results obtained from beams with

cover plates and design curves.


15/20


9 and higher design curves would be used for the 4 mm

thick specimens.

4.9. Fatigue data expressed in terms of the hot-spot

stress range

Although guidance exists for the determination ofthehot-spot stress, notably the IIW recommendations [8],

there is still the need for corresponding SNcurves. Pre-

liminary proposals for weldable aluminium alloys havebeen made by Partanen and Niemi [42]on the basis of

fatigue data generated at their university over a period

of years. The data were obtained from a variety of butt-

and fillet-welded specimens in 5000 and 6000 seriesalloys, including a model of the structural connection

between the deck, longitudinal stiffeners and a transversebulkhead in a naval ship. The range of plate thicknesses

was 36 mm. In all cases, hot-spot stresses were determ-ined from FEA or strain gauges using the procedures in

the IIW recommendations. The results(Fig. 15)were in

reasonable agreement and the authors proposed that the

IIW FAT40 design curve for transverse butt welds,

which is included in this figure, was suitable as a hot-spot stress SN curve for both butt and fillet weldedjoints in plate thicknesses up to 6 mm.

This reference to thickness is important because the

thickness effect discussed earlier will still exist even if

fatigue strength is expressed in terms of the hot-spot

stress. This is evident from another set of results for a

wider range of thicknesses, 324 mm, also expressed interms of the hot-spot stress range [25]. The test speci-

mens were all 6061-T6 aluminium alloy plates withtransverse fillet welded attachments. The results areshown in Fig. 16. Since the lowest SN curve, for 24

mm thick specimens, happens to correspond exactly with

the FAT40 design curve, it is tempting to conclude that

these data support Partanen and Niemis proposal. How-ever, the results clearly show a thickness effect that justi-

fies different hot-spot stress SN curves for different

Fig. 15. Fatigue data presented by Niemi and Partanen[44]as a basis

for the hot-spot stress SNcurve for thickness up to 6 mm.

Fig. 16. Fatigue test results obtained from 6061-T6 beams withfillet-

welded attachments expressed in terms of the hot-spot stress range

which illustrate a thickness effect [25].

thicknesses. A much larger database is needed to estab-

lish such SNcurves.

Other variables that might need to be considered

further are the influence of the joint type and thethrough-thickness stress gradient. Referring to the data

from Partanen and Niemi, there is a tendency for the

higher results to be for butt welds and the lower ones

for fillet welds, suggesting that there may be a need fordifferent hot-spot stress SNcurves for the two types of

joint. A notable exception is the single lap joint that gave

the highest results. This probably reflects the influenceof bending stress gradient, which would have been parti-

cularly high in these joints due to their inherent mis-

alignment. The effect, which is particularly significant

in thin sections, is to increase fatigue resistance. Thus,this is another thickness effect that needs to be con-sidered when selecting data from which hot-spot stress

design curves could be deduced. In general, data for high

stress gradients should be excluded unless the hot-spot

stress curve will only be used for similar conditions.

Finally, inorder to assess the DNV hot-spot stressS

Ncurves[6] mentioned earlier, they are included inFig.

15.As will be seen, both curves look reasonable. How-

ever, it should be mentioned that many of the results in

Fig. 16, virtually all those for 24 mm thick specimens,lie below curve III, suggesting that the thickness effect

correction should be introduced at a lower value than

that currently specified (i.e. 25 mm) in the DNV note.

4.10. Effect of marine environment

Marine corrosion fatigue (full immersion and a saline

atmosphere) of weldedaluminium alloys was studied in

some depth in Norway [28]in the early 1980s. Fatigue

tests were performed on transverse butt- and fillet-welded joints in 812 mm thick 5052, 5083, 6351 and7004 aluminium alloys. The tests on specimens in saline

atmosphere or immersed in seawater were carried out at

the low frequency of 1 Hz to allow time for the corrosion


16/20


reaction to take place. The tests were conducted in bend-

ing which meant that relatively high fatigue lives were

obtained, as compared with those expected for axial

loading. Based on comparison of the fatigue perform-

ance in air and seawater, the results were consistent withearlier studies by Sanders and McDowell [43]on 5000

series alloys. However, the effect of the environmentvaried with alloy type, 5052 and 7004 alloys being more

susceptible to environment than the others. In general,

immersion in seawater produced fatigue lives approxi-mately one-third of those obtained in air, corresponding

to a 25% reduction in fatigue strength. A saline atmos-

phere was generally less harmful, but not always. It pro-

duced a similar reduction in fatigue life to full immersion

in seawater in the 5052 alloy, while it produced an order

of magnitude reduction in fatigue life in the case of buttwelds in 7004 alloy. The effect was less severe in filletwelds. The influences of environment and alloy typeseen in this study are reflected in Eurocode 9 (seeFig. 6).

Fatigue crack growth studies in Russian AlMg5 and

AlZnMg alloys immersed in 3% sodium chloride sol-

ution[44]showed rather similar effects of environment

for both alloy types. Crack growth rate was increased,

but only significantly, by up to seven times, at relativelyhigh crack growth rates, with little effect of environmentnear the threshold. The threshold itself was effectively

independent of environment. These results suggest that

the effect of environment on SNdata referred to earlier

may have been largely associated with crack initiation,

which might also explain why butt welded 7004 alloy

was more susceptible to environment than fillet welds.

5. Friction-stir welding

All the data presented so far were obtained from arc

welded specimens. A new welding process that offers

considerably better fatigue performance is friction-stir.

Friction-stir welding (FSW) was invented at TWI, and

the first patent application was filed in December 1991.The process is an entirely new method of making con-

tinuous welds in several configurations using a solid-state process. The concept of FSW is illustrated in

Fig. 17(a). This shows a rotating tool that consists of a

shoulder and a pin. The former is pressed against the

surface of the materials being welded, while the pin is

forced between the two components by a downward

force. The rotation of the tool under this force generatesfrictional heat which softens the work-piece, and the

movement of the rotating tool along the joint line causes

softened material to flow from the region ahead of thetool to the region behind, consolidating to form a solid

phase weld. The process uses no filler, and for mostmaterials a shielding gas is not required. As the process

does not melt the materials being joined, materials such

as series 2000 and 7000 aluminium alloy, which are

Fig. 17. Friction-stir welding: (a) FSW process; (b) FSW joint in alu-

minium sheet.

often difficult to weld by fusion processes due to solidi-fication problems, are readily weldable. Experience hasshown that as the process is fully mechanised, high lev-

els of consistency can be obtained in weld quality. Inaluminium, it is possible to make full penetration singlepass butt welds in thicknesses of less than 1 mm to over

50 mm.

The process is used commercially by an ever-growing

list of companies in the aerospace, shipbuilding, railway

and automotive sectors. Almost all of the current com-

mercial usage involves aluminium alloys, although some

copper and magnesium alloys are also being welded. The

joining of other materials, including titanium alloys,

steel and nickel alloys, is under development.FSW of aluminium alloys produces joints of high

quality with static mechanical properties that equal, or

generally exceed, those of competing processes, butwith

lower scatter. An example of a weld is shown in Fig.

17(b).In view of the favourable profile, it is not surpris-ing to find that, under similar conditions, the fatigueproperties of friction-stir welds in aluminium alloys

compare very favourably with those for welds made by

MIG, the normal alternative. There are several examples

in the literature, all relating to 6000 alloys (since 2000

and 7000 alloys cannot be welded easily by the MIGprocess), although they are mainly confined to tests onrelatively thin specimens[4547].A typical example, for5 mm thick 6082 alloy tested under the relatively severe


17/20


loading condition of R = 0.5[45] is shown in Fig. 18.

In the same investigation FSW joints in 6005 alloy gave

fatigue test results close to those obtained from the

unwelded material, which is deliberately low to allow

for possible weld root flaws.In principle, it is possible to make any weld design

which does not require the addition of a filler materialby FSW, although most experience to-date relates to

butt- and lap-joints. A particular benefit is that full pen-etration butt welds are readily achieved in joints madefrom one side only, whereas such joints are highly vul-

nerableto root flaws in welds made by other processes.Fig. 18 includes the current Eurocode 9 design curves

for transverse butt welds made from one or two sides,

for comparison with the test data. It will be clear that

friction-stir welds made from one side can achieve con-siderably better fatigue lives than those indicated by the

design curve. Root flaws can still arise in FSWs, butthey can be relatively large before they affect the fatigue

performance of the joint [48]. Unfortunately, FSW can-

not be used to make fillet welds, as no filler is used.Therefore comparative fatigue data for fillet welds donot exist.

6. Fatigue life improvement methods

Fatigue life improvement techniques play an

important part in achieving higher design stresses when

the fatigue lives of structures are restricted by the pres-

ence of low fatigue strength details like fillet-welded

attachments. They may also be needed to ensure that aweld repair of fatigue damage will survive longer thanthe original detail. Thus, they are relevant to both the

original design and life extension of existing structure.

There are two main principles behind the various

improvement techniques [21]:

(a) Reduction of the stress concentration due to weld

Fig. 18. Comparison offatigue data for 5 mm thick 6082 alloy butt

welded by MIG or FSW[45].

geometry, e.g. dressing by machining, grinding or

TIG remelting.

(b) Introduction of compressive residual stresses, e.g.

peening (hammer, needle, shot and brush), ultrasonic

impact treatment.

In general, both types of improvement technique areonly readily applicable to surface stress concentrations,

notably weld toes. Improvement techniques have been

widely studied in the context of welded steel, but less

so in aluminium alloys. However, the general principles

should be applicable to aluminium, as confirmed in arecent review by Hobbacher[49]. However, the reviewshowed that most published data referred to butt welds,

whereas in practicefillet welds present the greater poten-tial fatigue problem. It was concluded that a fatigue

strength improvement at 2 106 cycles of around 1.4 or

more was justified for all the joint types reviewed, lead-

ing to the recommendations summarised in the follow-ing table:

Structural detail Treatment Fatigue strength

method improvement

factor

Transverse butt Laser dressing 1.4welds TIG dressing

Brush peening

Shot blasting

Cruciform joint Hammerfillet welds peening

Longitudinal Grindingfillet-welded Hammerstiffener peening

Transverse fillet- Brush peeningwelded stiffener

A practical problem with aluminium that can arise in

the case of dressing techniques concerns porosity. Flush-

grinding of butt welds or TIG dressing of butt or filletwelds can result in previously embedded pores being

exposed.TIG dressing can actually cause surface-break-

ing pores[50].They then act as crack initiation sites andcan actually reduce the fatigue life of the weld detail.

Although some design codes refer to the use of

improvement techniques, none provides recommen-

dations on the improvement in fatigue life to be achi-

eved. This partly reflects uncertainty about the correctapplication of the techniques and their effect on the

fatigue performance of real structures. In relation to the

last point, it is known that the fatigue performance of


18/20


welds treated by the residual stress techniques are sig-

nificantly affected by mean stress, any benefit disap-pearing at stresses approaching yield[51].Thus the tech-

niques may not be suitable for large structures containing

high tensile residual stresses or there may be doubtsabout their benefits in situations where the mean stress

is not known. The IIW is currently addressing both theprovision of specifications for the application ofimprovement techniques and corresponding benefits interms of revised SNcurves[52], including preliminaryrecommendations for welded aluminium alloys.

7. Fracture mechanics assessment of fatigue

Fracture mechanics offers the ability to assess thefatigue performance of aluminium structures containing

known or assumed flaws. In the context of the originalstructure, these could be manufacturing flaws, when theassessment might be required if the flaw exceeds themanufacturing quality standard being worked to. In the

context of the assessment of existing structures, they are

likely to be cracks formed during previous service by

fatigue or other mechanisms.

Eurocode 9 contains an Appendix with guidance onthe use of fracture mechanics for assessing welded alu-

minium alloys. This includes recommended fatigue

crack growth laws based on a large database produced

by Alusuisse [22]. These are expressed as a series of

Paris laws, to model the data more accurately than a

single Paris law, as illustrated inFig. 19.The same data

are referred to in BS 7910 [16] and the correspondingIIW recommendations [53]. These also contain verydetailed guidance on the use of fracture mechanics for

assessing welded structures, including a comprehensive

Fig. 19. Example of multi-stage fatigue crack growth relationships

for aluminium alloys proposed by Jaccard[22].

set of stress intensity factor solutions for the types of

crack and welded joint geometry which are likely to be

encountered. One advantage of the procedure in the IIW

recommendations is that it is linked with the IIW fatigue

design curves for welded aluminium alloys to facilitatedirect comparison of the fatigue performance of flaws

and that of basic design details in the same structure.However, BS 7910, which relates the fracture mechanics

assessment to British Standard design SN curves, is

more up-to-date.The multi-stage Paris law crack growth data fromEur-

ocode 9 have also been presented as polynomials [54],

in order to deduce better estimates of the probability of

failure associated with upper-bound curves. However,

regardless of the method of presentation, little has been

done to confirm that the same complex, multi-stagecrack growth relationship is a general law, applicable to

any cracks in real structures.

8. Future research

A number of aspects of both the design specificationsand residual life assessment methods considered in this

review would be improved by further research. The fol-

lowing are suggested as being the most important:

(a) Provision of fatigue data for non-arc welding pro-

cesses, particularly friction-stir but also laser weld-

ing, and their incorporation in design specifications.

(b) Study of cumulative damage under realistic stressspectra, with particular emphasis on the high-cycleregime and the damaging effect of stresses below

the constant amplitude fatigue limit.

(c) Further fatigue tests and FEA of structural details to

establish hot-spot stressSNcurves and guidance on

the practical application of the approach.

(d) Identification of potential fatigue design improve-ments that could be achieved by better use of special

extrusions, and generation of appropriate fatigue

data. Data are also required for transverse butt weldsmade on the backing provided by an extruded lip.

(e) Establishment of specifications for applyingimprovement techniques (of particular relevance for

life extension) to welded aluminium and experi-

mental confirmation of their value under realisticloading conditions (e.g. mean stress, loadingspectrum).

(f) As far as the use of fracture mechanics for estimating

residual fatigue life is concerned, the information

incorporated in BS 7910 probably represents the cur-

rent state of the art. However, it does place particularemphasis on steel and experimental work to decide

on the choice of fatigue crack growth relationships

appropriate for aluminium alloys and experimental


19/20


validation of the fracture mechanics approach in gen-

eral would be useful.

9. Conclusions

Based on a review of published information on fatigueassessment procedures for welded aluminium structures

and supporting experimental data, the following con-

clusions can be drawn:

(a) Of the three fatigue design assessment procedures

described, that using the nominal stress SNcurves

is the most developed and standardised, but the hot-

spot stress approach will probably prove to be the

most valuable for structural design in future.(b) Several national and international fatigue design

specifications have been published in recent years.Eurocode 9 and the IIW recommendations are the

most comprehensive.

(c) Even so, most design data refer to arc welded joints

and there is a need for corresponding data for other

welding processes (e.g. friction-stir).

(d) There are significant differences between all the pro-posed design SN curves and the fatigue test data-base available for large-scale structural specimens,

mainly due to the choice ofSN curve slope. Thus,

some specifications are unduly conservative in thelow endurance regime and others in the high-cycle

regime. Eurocode 9 curves were generally the most

consistent with experimental data.

(e) There is a scarcity of data for weld details incorpor-ating special extrusion shapes, (e.g. built in backinglip) and little effort seems to have been made to

investigate improved fatigue performance of welded

aluminium structures by better use of extrusions.

(f) The fatigue performance of commonly welded alu-

minium alloys is not greatly influenced by immersionin a marine environment. The recommendations in

Eurocode 9 are consistent with relevant published

data.

(g) There is little information in the literature on remain-ing life assessment procedures. The most appropriate

are the use of design SN curves or, if allowance

must be made for damage sustained during previous

service, fracture mechanics. Comprehensive guid-

ance and recommended input data for the application

of fracture mechanics are contained in BS7910:1999.

Acknowledgements

This work described in this paper was funded partly

by the Australian Maritime Engineering CRC Ltd. and

partly by Industrial Members of TWI Ltd. The author is

also grateful to his colleague Dr P.L.Threadgill for his

assistance with the section on friction-stir welding.

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