Lloyd’s register integrated fatigue design assessment system

LLOYD'S REGISTER INTEGRATED FATIGUE DESIGN

ASSESSMENT SYSTEM

Abstract

This paper describes the multi-level fatigue design

assessment procedure developed by Lloyd's Register

to estimate the fatigue strength of ship structural

details. The method used to determine a representative fatigue strength capability for ship

structural details, and the supporting fatigue testing programme of large ship structure are reviewed. The traditional maximum lifetime load approach to

estimate the fatigue demand is discussed, and the spectral analysis procedure adopted in the subject

procedure is described. Examples of distribution of fatigue damage around the hull envelope are given to illustrate the influence of the load components, and the wave environment.

1. INTRODUCTION

The objectives of Ship Classification are to develop

and implement Rules and Regulations, which in conjunction with proper care and operation on the part of the shipowner and operator, will provide an

appropriate standard of structural strength during the service life

of the ship. On the provision

of an

adequate level of maintenance associated with regular structural surveys and appropriate repair procedures for incidental structural damages, Classification Society Rules are intended to

safeguard the structural integrity of the hull structure

for a service

life period

of at

least 20 years.

Computational structural analysis methods combined

with service experience of large and complex ship structures have shown that success or failure in a

structural sense, is directly affected by the

performance of structural details. The failure of structural details is generally associated with two interrelated cumulative damage processes, namely,

fatigue and corrosion, Violette (1994). As early as

1913, Lloyd's Register expressed its concern on the fatigue mode of failure of ship structural details in a Ab. No. 021

Franck L. M. Violette, DUT., M.ENG.

Advanced Studies & Rule Development Group

Technical Planning & Development Department

Lloyd's Register of Shipping - Ship Division

71, Fenchurch Street

London, EC3M 4BS, United Kingdom

paper presented to the Royal Institution of Naval

Architects, Lloyd's Register (1913). Over the last

decade, fatigue failures of structural details and their potential consequences have attracted an increased attention amongst the shipping industry. In view of the higher percentages of high tensile steel being used, the application of sophisticated techniques to peorm structural and fabrication optimisation, and the

implementation of strict environmental regulations, the occurrence of fatigue cracking cannot be considered any longer as a mere fact of life. Since fatigue cracks can be possible points of initiatión for the failure of the cargo containment barrier, and/or significant structural failures, it is essential that fatigue performance be given a more detailed consideration at the design,

and construction stages as well as during the ship's operational life.

In 1989, Lloyd's Register (LR) initiated a long term

research and development programme with the aim to develop and implement an integrated fatigue

design assessment procedure for ship structural details. In June 1994, ShipRight FDA (Fatigue Design Assessment), a multi-level integrated fatigue design procedure has been released to world-wide LR plan approval offices and to shipyards. This paper reviews the main components of the research project,

and outlines the level 2 ShipRight FDA integrated

fatigue design assessment procedure. Some of the more salient aspects of the procedure are illustrated

with examples from the fatigue

analysis

of the

longitudinal members of a double hull tanker.

2. SHIP STRUCTURAL DETAIL FATIGUE DESIGN 2.1 GeneraI

Whilst for at least the past two decades, Classification

Societies have carried out fatigue calculations for

damaged or novel ship structural components, procedures or criteria for the design against fatigue

have not been explicitly or rationally considered in

1/17

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Classification Society's Rules until recently, except for

specific ship types such as LNG ships. However,

implicit fatigue requirements are incorporated for

many structural components through the use of

permissible stress levels determined from the application of simpUfied fatigue damage assessment techniques combined with service experience data. In fact, fatigue has been given due considerations at all stages of the Classification process through implicit requirements or procedures as follows

During the plan approval process, in addition to the use of permissible stress levels, the surveyor applies particular attention to the design of critical

structural details by using a set of experience

based detail design recommendations compiled

over the years. This s carried out in close co-operation with the ship designers in order to achieve the best solution in terms of both the

fatigue performance and fabrication;

During the construction stage, the field surveyor

assigns particular attention to the critical areas

which have been

identified during the plan

approval process, in order to ensure that satisfactory levels of workmanship, alignment and fit up are achieved;

And, finally, during the life of the ship, the scope

and extent of the periodical surveys give due

attention to the critical structural details, and the

detection of any onset of cracking. Should a fatigue crack be detected, remedial action is taken to prevent its reoccurrence, and the survey data is transferred to the LR damage database for analysis. Noteworthy defects or increase in

fatigue failure trends of structural components are communicated to plan approval surveyors, and

Rule development engineers, and an implicit fatigue criteria such as a permissible stress level, or a detail design recommendation may be

formulated for inclusion in the Rules.

2.2 Limitations

of

an Experience Based

Procedure

However, it should be recognised that with new generation ships such as double hull tankers, the

value of service experience may not be directly

applicable. Whilst many of

the new

oil tanker structural

details can be related

to existing hull

configurations, the increased proportion of high

tensile steels, the extent of structural optimisation, as well as changes in the loading patterns could change the significance of an experience based procedure,

Ferguson & Violette (1991). Significant changes in

terms of ship structural design and construction, and

the introduction of strict environmental regulations have called for the development of rational direct

calculation fatigue design procedures where the following aspects which are of particular concern need to be addressed:

The extrapolation of ship structural concepts may not be directly applicable to new designs, and the service experience of double hull tanker structural

details, especially large double hull tankers, is limited;

Recent experience with high tensile steel used on second generation single skin VLCC structures

has shown that when a significant degree of structural optimisation has been carried out on

the primary web structures, the additional loads created on the secondary members may increase the risk of fatigue damage;

With double hull tankers, cargo leakage as a

result of fatigue cracking would result in serious difficulties to clean and/or ventilate the double hull spaces due to their cellular arrangement. Moreover, leakage of cargo oil, or inert gas into

the ballast spaces would place the ship in a

hazardous situation with potential risks of explosion, loss of life, and environmental disaster; The ship structural performance and life could be

severely influenced by the ballast tank coating

performance. The initiation of fatigue cracks in the ballast spaces may precede coating breakdown. If undetected, or if no remedial action has been taken, the localised coating breakdown in way of a detail stress concentration will create a severe localised corrosion cell, which in turn will

accelerate the fatigue crack propagation process, Violette (1994);

The quality of the workmanship, and the

construction tolerances may influence the fatigue

performance of ship structural details. Fatigue cracking as a result of poor detail design may be difficult to cure without significant and expensive structural modifications.

3. ShipRight FDA FATIGUE DESIGN ASSESSMENT

lt can be appreciated that the evaluation of fatigue performance of ship structural details is a complex

process, and that fatigue failures may have dramatic consequences. To attain and maintain a satisfactory

fatigue performance, a realistic fatigue procedure

should give due consideration to the following stages in the life of a structural detail

The conceptual design of ship structural details;

The analysis of the fatigue performance by a

direct calculation method giving due consideration to the cumulative nature of the fatigue damage process, and accounting for fabrication and

workmanship factors;

The verification of the structural details

workmanship, alignment and fit-up tolerances, through adequate survey procedures and defects acceptance criteria during the ship construction; The in-service survey procedures, and the

monitoring of the critical structural detail items

during the ship lifetime.

For this purpose, ShipRight FDA (1994) has been

developed as a multi-level fatigue design assessment procedure (Level 1,2 and 3) to address each of the requirements highlighted above. The procedure is a

total approach to the prevention of fatigue failures

encompassing the design, construction and in-service performance of ship structural details. lt s supported by the ShipRight CM (1994) Construction Monitoring

procedure. and the ShipRight HCM (1994)

Hull

Condition Monitoring procedure. The main features of

the multi-level FDA procedure are reviewed in the

following sections.

LEVEL 1: STRUCTURAL DETAIL DESIGN GUIDE

The primary purpose of the Structural Detail Design

Guide is to promote good detail design at an early

stage of the design process, and to provide guidance for improvement of detail design. The Guide has been compiled from the world-wide detail design and the in-service expertise of plan approval, newbuilding and field surveyors. Therefore, it

is based on a vast

experience based knowledge database considering

aspects such as design and analysis, construction tolerances and fabrication issues, and in-service performance. In addition, extensive analytical, and Finite Element Analyses (FEA) have been performed for each recommended structural details to confirm and optimise the structural configuration to maximise

its fatigue performance. A typical Structural Detail Design Guide datasheet is shown in Figure 1. At

present, the Structural Detail Design Guide addresses double hull tanker, and bulk carrier critical areas. The

development of the Guide is considered to be a continuous process with regular updates to reflect

trends in service experience, design and construction

practice as well as to incorporate results from the

ongoing FDA and FEA studies, and LR fatigue testing programme.

SHIP STRUCTURAL DETAIL FATIGUE STRENGTH

For design purposes, the S-N curve approach to the estimation of the fatigue strength capability is considered to be the most common and convenient approach. However, the assignment of a S-N curve to a ship structural detail will invariably involve a certain amount of engineering judgement. To ensure that the FDA procedure could be applied consistently, it was considered essential

to automate the S-N curve

selection process. This task is performed by the FDA S-N Curve Expert procedure.

5.1 NominaI and Reference S-N Curve

The traditional nominal stress S-N curve approach s expressed as a function of the nominal stress range

AS for a given typical structural detail as follows:

(1)

Traditional nominal stress S-N curves are widely

available from design codes such as BS5400 (1980), UK.Den (1983), Eurocode (1985), 11W proposal

(1982), etc. In general, these standard are applicable to a limited set of typical standard details. In view of the large variety of ship structural details in terms of

both geometry and loading, the application of standard detail S-N curves has shown to be a difficult process. Comparative fatigue studies correlated with service experience and/or full scale measurements

have shown that standard S-N curves may yield

fatigue lives significantly different from the recorded values.

A more explicit method to apply the standard S-N curves to ship structural details is to derive,, the geometrical Stress Concentration Factor (SCF) of the standard detail geometry, as well as the SCF of the ship structural detail, in

order to obtain a more

representative S-N curve. The S-N curve may

therefore be rewritten as follows ship

N

1\ K,d =

j

,fl' AS

K,d '

The disadvantages of this approach is that geometrical information on the standard S-N Curve structural detail is often limited,

and FEA with

compatible mesh size for both the standard structural

detail, and the ship structural detail is required to

determine the SCF's. Moreover, it should be borne in mind that standard S-N curves tend to represent the lower limit of fatigue capability of the standard detail i.e. geometrical and scantling configuration leading to the worst fatigue strength within the scope of application of the detail. For example, the transition between BS5400 (1980) Class F and F2 is dictated

only by the attachment length (150 mm), and the

edge distance (10 mm), and no reference is made, for example, for the influence of the plate thicknesses.

To enable a more representative estimation of the fatigue strength for ship structural detail within the

Ab. No. 021 3/17

Fig. I

Structural Detail Design Guide Typical Datasheet

Lloyd's Register - ShipRight Strategic Research & Development Page 5

r

LOCATION: Connection of side and longitudinal bulkhead longirudinais to transverse webs in double side tanks

EXAÍVWLE

No. 1:

Asymmetrical face/higher tensile steel side longTh.idinal to

transverse web flat-bar stifferters

GROUP

No. i

CRITICAL AREAS DETAll. DESIGN IMPROVEMENT

SHELL PLATING

______________

-BULKHEAD

x

NOTE

SOIT

SOFT//EEL TOE _{[1 SOFT} ÁMD OR SR4CKETS DETAIL è4AX1 _(M S)-A1A1ETRIC4L TOE h-MAX 15

L2.Od

U U

liLi

U 4cRmcALAREAS B,ac*et Á.th,nwm

o.J

75

4

X

0.5)

4

--MAX15

I

ThIckness Flat Sa,

Thickness 12Omm

-4-MAX15 Thickness I

=6/18

dj

d180-300 Areaof stiffener accordancewith Rule requirements

CONSIDERATION ITEMS FOR IMPROVEMENT

Critical Location: Asymmetrical face / higher tensile steel side longitudinal face bar connections at the heel and the toe of the web stiffeners. Connections between the base line and O.8D above the

base line.

Fatigue Mechanism: Soft toe and heel detail or symmetrical soft toe brackets to reduce peak stresses under fatigue loading from dynamic seaway loads and ship motions.

Building Tolerances: Ensure alignment of the web stiffener, the back bracket and the web of the side longitudinal.

Welding Requirements: Fillet welding having mínimum weld factor of 0.44 (Web stiffeners to face bars of side longitudinals. Back brackets to face bars of side longitudinals).

A wrap around weld, free of undercut or notches, around the plate thickness.

FIGURE

i

DETAIL DESIGN GUIDELINES FOR DOUBLE HULLTANIR STRUCTURAL DETAILS

framework of a design procedure, it was considered

_{Ka = Kg(R,tid7.la,d,c,th)Kje)K,i(ö)K,(e)}

₍₅₎

that:

The nominal stress should remain the reference stress used for FDA purposes, as it is a convenient stress measure for engineers involved in structural design assessments;

And, a large library of S-N curves based on

actual ship structural detail geometries should be made available.

To satisfy these requirements, the hot spot stress

approach was reformatted as follows

Thus,

=

N = IÇ(Ka AX)"

N=

_nl, ShI

Based on FEA of standard structural details, and the

available S-N curves data, a reference S-N curve

defined by K1, has been evaluated. The reference

S-N curve represent the fatigue strength

of the

welded material including the fillet weld geometrical stress concentration. The geometrical SCF of the ship structural detail then becomes the driving parameter to estimate the fatigue strength, and its derivation is reviewed in the following Section. For reference

purposes, the S-N curves derived by the FDA S-N

Curve Expert procedure are assigned a value which represent the stress range magnitude of the mean S-N curve at stress cycles.

5.2 Geometrical Stress Concentration Factor Based on systematic

linear elastic FEA of ship

structural details, parametric formulations

of the

geometrical SCF's have been

derived. The FE

modelling standard adopted to determine the SCF's is

based on a FE mesh of t x

t

in way of stress

concentration areas, where t is the thickness of the plate containing the crack initiation site. The SCF has been determined using the centroidal stress normal to the expected crack plane for the element adjacent to the theoretical intersection edge. Taking into account that the weld factor for the connections under

investigation is of the order of 0.44t, the SCF stress

gradient pick up point is at a minimum distance of

0.06t from the weld toe. A typical FE mesh for a soft toe - soft heel web stiffener connectìon to the flange of a longitudinal member is shown in Figure 2.

The parametric SCF formulation is defined as a combination of influence functions defined as follows for the end connection geometry shown in Figure 3.

(4)

A-A

ti

A-A

nCrltical Location

th

(3)

Fig. 3. Stress Concentration Factor Parameters.

The axial loading SCF has been chosen as the

reference SCF for the computation of the structural

detail fatigue strength capability by the FDA S-N Curve Expert. To account for the influence of the

SCF's for modes of loading other than axial, loading mode bias factors are introduced in the computation of the total nominal stress applied to the detail. The loading mode bias factor for loading componen.t i is defined as follows

K

KB,

=-Ka

(6)

To illustrate the SCF parametric formulation, Figure 4

shows the variation of Ka

for a soft toe flat bar

stiffener connection with

d = loo

mm,

t1/t2 =1.07th =I5mm,Kw(0)K,,,(S)Ke(e)=l.

Ab. No. 021 5/17

100 150 200 250 300 350

SoftToe Radius -mm

Fig. 4. Ka versus Soft Toe Radius

ASh. K

ç ''

h, K 1.350 1.325 1.300 1.275 1.250

Figure 5. illustrates the variation flat bar stiffener connection

R=I5Omm, th=l5mm,

Kw(0)K,,,(Et)Ke(e) = i

K 1.350 1.300 1.250 1.200 1.150 1.100 1.00

Fig. 5. K versus t1 ¡t. Ratio

lt can be appreciated that the FDA S-N Curve Expert

approach permits to determine more realistic S-N

curves for ship structural details, since the S-N curve

is a direct function of the SCF, and thus the detail

geometrical parameters.

5.3 Fatigue Testing of Ship Structure Models

To increase the confidence level in fatigue strength predictions and to gain a better understanding of the

fatigue crack initiation and propagation process in ship structural details, Lloyd's Register initiated a

programme of fatigue testing of large ship structure models in 1992. The Krylov Shipbuilding Research

Institute in Russia was commissioned to carry out

these tests. There are many benefits associated with performing large scale fatigue tests of realistic ship structures which may be summarised as follows

Shipbuilding workmanship standards are used to fabricate the models;

Realistic ship loads are applied to the models resulting in a more realistic stress field , stress concentration levels, and cyclic stress patterns in way of the critical locations;

The kinematics

of crack

propagation at the potential crack initiation sites provide valuable information with regard to the relative severity of each potential crack initiation sites, their propagation rates, and the potential consequences when applied to real ship structures;

The assessment of the redundancy level of the

structural system, and the load redistribution

of K for a soft toe

with

il = 200

mm,

= 300 mm,

and

mechanisms which may affect the crack propagation rates at certain crack initiation sites.

Extensive post analysis using FEA can be

performed to confirm the strain / stress

measurements, to extend the results to the full scale structure, and to optimise the structural

configuration with respect to fatigue strength; And, finally, experimental data combined with FEA permits the calibration of the parametric

formulations of the SCF's associated with the

reference S-N curve which have been

implemented in the FDA S-N Curve Expert.

To date, the ongoing experimental research

programme has addressed the following ship

structural details

1/4 scale models

of VLCC hopper welded

connections to double side and double bottom;

1/4 scale models

of VLCC hopper

flanged connections to double side and double bottom;

1/3 scale models of VLCC longitudinal connecons

to web fiat bar stiffener with various end connections, and steel yield strength.

Photo 1-2 show the large scale ship structures models under testing conditiqns.

io

to 106 load

cycles are applied to the models which are

instrumented with uniaxial and rosettes strain gauges

to monitor the field and the hot spot stresses.

Acoustic monitoring has also been used to determine

the spatial location of the onset of cracking well

before the crack could be detected by visual inspection. The crack propagation rates have been recorded in order to monitor the crack growth in the structural components, and enable further analysis to be performed using fracture mechanics methods. 6. LONG TERM STRESS RANGE SPECTRUM

MODEL

Since the wave environment generates complex

loading patterns, it can be appreciated that the prediction of the long term stress spectrum for marine

structures remains a complex problem. Due to the

cumulative nature of the fatigue damage process, it is

essential that an adequate procedure be used to

predict the long term stress range distribution. Figure 6. illustrates the distribution of fatigue damage for

three typical long term stress range spectrum defined analytically as Weibull functions with shape factors of 0.8, 1.0, and 1.2. lt is shown that most of the fatigue damage is produced by the small to medium stress ranges i.e. the low to medium seastates, by virtue of their associated number of stress cycles ( iO5 to 108

Ab. No. 021 6/17

1.25 1.50 1 75

o o -c Q-N o o û-

N-Fig. 2 Typical Fine Mesh FE Model for Soft Toe Soft Heel Longitudinal End Connection

l.

LL

L I

T.1I I

LL L .L

L I

_[_ J 11.11

i L

L

-L -L_-L-L

.

L LLLLLLL.LL

..

-L_[:LLH

_JL

...

.:...

.

:J

L U

.-. U... :

LI

t i t I

Il

t

f

Li

1l(lM

:.c&: .. 1 1 t 1.1 LI_t I .

IllllL

j i_l I I Lt

'I

L

ro:r

rr

:i

tif.

..

(t11llIll

LL

_L

LL_1LU_

L .J

LL.1_l_.

ELL

-

..L.LLLLLLJJ LJ._

i

ILl l.i

ULLI.rJ.LLL.t .i.L

i i.._

..IL_[.LJJ.L.LHI!iH...

LLLLLLLLLL _L I

_L LL L

LLLL L

r

a

Long Term Weibull 0.80

- - - Long Term Weibull 1.20

Fatigue Damage - Weibull 1.00

Long Term Weibull 1.00 Fatigue Damage - Weibull 0.80

- - - Fatigue Damage - Weibull 1.20

Fig. 6. Typical Long Term Stress Spectrum and Fatigue Damage Distributions stress cycles). Therefore, in

order to achieve a

reliable prediction the structural detail fatigue life, the Ab. No. 021

mathematical model used to predict the long term stress distribution should give due attention to this

8/17

--/

--

I

1E+00 1E+01 1E+02 1E+03 1E+04 1E+05

1E06

1E+07 1E+08

Number of Stress Cycles

5.00E-03 4.00E-03 cl) o) - 3.00E-03 E

o

a) - 2.00E-03 o) u-1.00E-03 0.00E+00 225.00 200.00 175.00 150.00 125.00 C) 100.00 g _75.00 cl) 50.00 25.00 0.00

part of the spectrum. To determine the long term

stress spectrum, two approaches are available, namely, the maximum lifetime load approach, and the spectral approach. Both procedures are reviewed in the following sections, with emphasis on the spectral method of analysis used in the FDA procedure.

6.1 Maximum Lifetime Load Approach

The main advantage of the maximum lifetime load

approach is its inherent simplicity. Based on the readily available maximum lifetime loads parametric expressions used for strength assessment, this procedure is well suited for design purposes. Figure 7. illustrates the procedural steps involved in the prediction of the long term stress spectrum. However, since fatigue damage is proportional to the cube of the stress range, and the procedure emphasises on the 108 stress range which is remote from the area of interest with respect to fatigue, variations in the 108

stress range and/or the Weibull shape factor can

result in a significant variation in the resulting fatigue life as shown in Figure 8. In view of the sensitivity of the resulting fatigue life to the limited set of modelling variables, some of the modelling assumptions need to be reviewed:

For a given

structural detail, each individual applied maximum lifetime loads is calculated at a probability level of 108. Considerations to loads computed at a probability level of say, 10 , may be more appropriate provided the iO4 loads have not been extrapolated from the 108 base usingan

assumed long term distribution;

20.00 10.00 5.00

PiJ11

100 in

220 280

----'

0.80 400 ₄₆₀ Stress range - N/mm2

The partial load factors or correlation coefficients

used in the load effect combination model are

often based on the dominant load parameter. and represent a snap shot seastate where the

dominant load component is maximised. This combination may not be representative of the less severe seastates represented by the lower part of

the spectrum, where other headings, wave

heights and periods, or load components may

dominate;

The number of stress cycles for a service period

of 20 years

is

in general based on the 10

incident waves assumption used for the North

Atlantic longitudinal strength requirement. Where several load components are involved, the resulting number of stress cycles may differ from the number of incident waves;

The Weibull shape factor is often mapped on the

dominant load parameter shape factor for the

North Atlantic wave environment, and it is

questionable whether this is representative of the

resultant load effect arising from several load

components. Furthermore, since the shape of the stress spectrum is directly affected by the wave environment, the stress response in each of the seastate, the ship loading conditions, and the ship seakeeping performance, ¡t is disputable whether this assumption is applicable for a range of ship types, and characteristics.

w C-_C

cn2

o

_DLL w

Fig. 8 Sensitivity of Fatigue Damage to Stress Range and Weibull Shape Factor

Ab. No. 021 9/17

Identification of Load Process Applied to Structural Detail

I

Computation of Maximum Lifetime Loads Magnitude

P=lo -

I

Computation of Structural Influence Coefficient

I

Computation of Maximum Lifetime Stress Magnitude

P=iO

-I

Identify Dominant Load Effect & Assign Partial Load

Effect Factors

I

Weibull Shape Factor based on Dominant Load Effect

I

% Lifetime Number of Incident Waves

I

Wave Induced Loads &

Motion Computation Structural Response Analysts Statistical Analysis Fatigue Strength Voyage Simulation Levei2 " Analytical / Parametric Formulation Level 3 - First Principle Ship Motions

& Load Software Level 2

- Analytical, FE Bean, Model SCF

Levei3

e

Global 3D FE Model Local Zoom FE Model

Seastate Wave Height - Wave Period

Ship Heading - Ship Speed

Structural Detail Fatigue Strength - S-N Curve

loo Al Fatigue Wave Environment Load Process I I-Pr@10

cl

V S1@ 10-e V L = LIC'? + ." Shape Parameter SpecFDA SoWer

Short Term Stress Response in Irregular Waves

Deterministic Fatigue Life Probability of Failure

Regular Waves Amplitude & Phase Angle

Hull Girder Loads, External Wave Pressur Internal Cargo Pressure

Local Stress Influence

Coefficients

SSC Wave Energy Spectrum

Service Profile Probability Matrix

Wave Height - Wave Period Ship Heading - Ship Speed

Loading Condition

Safe Life Acceptance Criteria N V Pn@10 ' C l IO L V

Loads - Load Effects Stress Combination Short Term Fatigue Damage Rat. Service Experience 4 Fatigue Monitoring

Uncertainty Stress Model

Uncertainty Workmanship

Uncertainty Fatigue Model

Maximum Load Approach Lifetime Stress Spectrum

I

Fig. 7 Maximum Lifetime Load Procedure

Fig. 9 Spectral Fatigue Analysis Procedure

Ab. No. 021 10/17

. Load Process n

.

..

.

L

Calculate Total Stress

I

Calculate Stress Range

lt can be appreciated that this procedure is subject to

a number of assumptions, and modelling

simplifications. To achieve reliable fatigue life

estimates, calibration using service experience data has been shown to be essential. Whilst this procedure

cannot be rejected

as a

design tool to obtain

efficiently an estimate of the fatigue strength, the

sensitivity of the fatigue process, and the dependence of the required calibration on the service experience

base make this procedure difficult to apply in a consistent manner. Furthermore, reduced confidence levels are introduced when the structural configuration or the loading patterns depart from the service experience base used for calibration.

6.2 FDA Spectral Analysis Procedure

To enhance the level of confidence in the

determination of the long term stress spectrum, and

address the issues of the maximum lifetime load

approach highlighted above, the FDA procedure uses

a first principles approach based on the spectral method of analysis. Two levels of analysis have been developed using the same theory, but different level of mathematical modelling. Level 2 is used for design purposes and uses simplified mathematical models. Leve! 3 uses sophisticated mathematical models, and is aimed at confirmation of the fatigue performance of novel structural details. The procedural steps of Level 2 and Level 3 are illustrated in Figure 9

The application of the level 2 procedure to longitudinal members of a double hull tanker can be summarised in the following steps

6.2.1 Computation of Wave Induced Loads

The wave induced primary load components considered are

External hydrodynamic wave pressure; Hull girder vertical wave bending moment; Hull girder horizontal wave bending moment; Water ballast/cargo inertia pressure.

The amplitude and phase angle of the above load

components are calculated in regular waves for the following ranges

of wave parameters

for each significant loading conditions as shown in Table 1. Table i

Ab. No. 021

FDA Level 2 computation of the wave induced loads is based on a parametric formulation of the ship six degrees of motions, and the global and local loads.

Using the basic concept of the single degree of

freedom vibration model, and the systematic analysis of the computation of the ship motions and loads of

over 250 representative hull forms, the Response

Amplitude Operator (RAO) for the motions and loads

has been decomposed into a series of influence

functions based on the work by KSRI (1992)

RAO(V, x,w)=

f (V,

,[a, ]

)f(V, x.

[a1 ]_) ( '

For example, for the vertical bending moment, the

RAO can be further decomposed to

isolate the

influence parameters as follows:

RAO(V,,w)= fa'80(L,B,T,Fn,Cw,CB,k,j(h)

fa(X)fa

(x) L)f (x)

(8)

The amplitude function

f

180(.) determines the maximum response amplitude which for the vertical

bending moment occurs at 180°. The amplitude function fa(X)describes the ratio the response to

the maximum response amplitude at 180° for a given

ship to wave angle as shown in

Figure 10. The

amplitude function

_{f (x)}

describes the position of the maximum response amplitude on the wave frequency axis as shown in Figure 11.

1.00 0.80 0.60

0.40

0.20

Parameter - Range Increment

7.

0-360°

20°

0.2-1.2rad/s

0.O4rad/s

V

0-Vs

25%Vs

Loading Conditions Fully loaded Ballast

0.00

o 30 60 90 120 150 180

Ship to Wave Heading - Degrees Fig. 10. Vertical Bending Moment fa(X)

The function

_{f, ()}

describes the RAO basic shape as a function of the wave frequency, response natural

frequency, the ship length, and the function f (x).

L (x) is

the function describing the longitudinal distribution of the bending moment.

Expressions in a similar format have been derived for the horizontal, and torsional bending moments, the ship motion six degrees of freedom, and the external hydrodynamic wave pressure RAO's. Since the phase

angles do not lend themselves to a relatively simple description by analytical functions, the phase angle of

the RAOs has been

stored in a database for

computation purposes pending further analysis.

1.40 1,30 1.20 1.10 1.00 0 30 60 90 120 150 180

Ship to Wave Heading - Degrees

Fig. 11. Vertical Bending Moment f(x) 6.2.2 Structura! Influence Coefficients

For each load components considered above, the

structural influence coefficient C, ( i.e. direct stress normal to the most likely plane of crack propagation

at the weld toe connection of web stiffener to the longitudinal flange ) due to the application of a unit

load ¡ is determined as follows

.

Vertical wave bending moment

C1

= K8

(9)

.

Horizontal wave bending moment C2 = K13

zzz I

Hydrodynamic wave pressure

C = K8

12ZLF

f(x)

Water ballasticargo inertia pressure

C4 = K8 12ZLF

f(x)

C5=KB

6E1 C6

= K8

12 4 EI

C7=K8

=

S (w) dw

Assuming that the stress process is narrow banded,

(10)

_{the stress range distribution can be expressed} ¡n terms of a Rayleigh distribution

s

s2 4: 2 exp 8

-)

(20)

For the side shell, where the presence of the wave free surface creates a non-linear effect with a

truncation of the pressure load harmonics, a time

domain simulation procedure is performed to

calculate the short term stress statistics.

(12)

For the secondary load components arising from the deflection of the primary structure, the structural influence coefficient C due to the application of a unit deflection or rotation i is determined as follows

Deflection Side 1 Deflection Side 2 Rotation Side 1

4E1

Rotation Side 2

Ç = K1

/

(16)

For the secondary load component arising from the deflection of the primary structure, the deflections and rotations transfer functions are calculated from FEA

for a number of wave cases representative of the

array of regular wave conditions defined in Table 1.

6.2.3 Short Term Fatigue Damage

For a given ship to wave heading, wave frequency. ship speed and loading condition, the total stress can be expressed as follows

n

(17)

For the given stress check point location, ship loading

condition, ship speed, ship heading to waves, and

seastate expressed in terms of significant wave height H113, and mean zero crossing period T, the

short term stress

statistics are calculated. The

spectral function S (w) is calculated directly from

the wave spectral function S(w) (ISSC spectrum),

the transfer function H1 (w) of the ith load process,

and the complex conjugate of the H(co) of th jth load process as follows:

S(w)=S(w)C,C1H,(w)H(w)

(18)

The spectral moments required for calculation of the spectral bandwidth and zero crossing frequency are given as follows

Ab. No. 021 12/17

(19)

Using a closed form solution for the fatigue damage, the short term fatigue damage rate and associated stress cycle rate can be calculated. For a given stress

check point location, ship loading condition, ship speed, ship heading to waves, and seastate expressed ¡n terms of significant wave height H113, and mean zero crossing period T, the accumulated fatigue damage ¡s expressed as follows:

D=''

(21)

N(S,)

The accumulated fatigue damage in one seastate can be expressed as follows

D=

= Tu

and

K

+ i 1m,

Uo

=-:

The deterministic fatigue damage accumulated in a

given seastate can be obtained from the following

expressions TB " Q

D=

(25)

(22)

For a narrow banded process, the accumulated

fatigue damage in one seastate may be rewritten as follows:

B"

n

K _I)

JS"p(S) dS t(m,m1,S0,B) (23)

j..i(m,m1

,S0, B)(2J)"

"F

Since the stress process is not a strictly narrow

banded process, a rainflow correction factor X(m,c) is introduced to remove the conservatism due to the

narrow band assumption, Wirshing (1977). The expected number of stress cycles is obtained from the stress process zero crossing frequency as follows

(24)

m

+1 (26)

2

For each seastate, the short-term fatigue damage

accumulation rate, and stress cycle rate are computed to enable the computation of the long term fatigue damage.

6.2.4 Voyage Simulation

Since fatigue damage is a cumulative process, and the long term stress range distribution is a function of the long term wave environment, it is essential that due consideration

is given to the derivation of a

realistic wave environment. Using a concept similar to

100 Al longitudinal strength standard based on the

North Atlantic wave environment, the 100 Al fatigue wave environment standard has been formulated. lt is computed systematically for a combination of trading

routes for the ship type, and ship characteristics

subject to the FDA investigation. The trading routes are a direct function of the ship type, and they have

been determined from statistical analysis of

world-wide trading pattern. The Global Wave Statistics data,

BMT (1986), is used to determine a service profile

matrix giving the probabilities of occurrence of the seastates defined in terms of significant wave height, mean wave period, loading condition, ship to wave heading and service speed.

6.2.5 Computation of Long Term Fatigue Damage

The total lifetime accumulated fatigue damage Dt over

a specified service period Ts is computed from the

Q= X(n

).i(m,m ,B,S0)o(2-.h)"o'

probability matrix of occurrence of the short term

seatates as follows'

T,. B"Q

Dt=

(27)

K

where

Q, =

_P,P,PkP/,

_jkl and ¡s the

. .k ,1

stress level parameter for a given seastate i,j,k,l

6.2.6 Fatigue Acceptance Criteria

The fatigue life results are given in two formats as

follows

The conventional deterministic fatigue life format calculated with an S-N curve with an associated

probability of survival of 97,5%, and a fatigue

damage factor of 1,0 for 20 years;

The probabilistic format based on a simple log-norma? format for multiplicative limit state functions yields a probability of failure and a safety index for a given number of service years Wirshing (1987). 7. DOUBLE HULL VLCC FDA APPLICATION To illustrate typical distributions of fatigue damage,

FDA computations have been performed foF an

idealised double hull VLCC ship with uniform typical

end connection design i.e. flat bar web stiffener

200x12 around the envelope.

7.1 Influence of Loading Components

Figures 12,13 illustrate the fatigue damage distribution due to the hydrodynamic wave pressure

alone in head seas, in a fully loaded and ballast

condition respectively for a given seastate (H113,T). lt

can be seen that the maximum fatigue damage

occurs below the waterline. In the fully loaded

condition, the fatigue damage in way of the outer

bottom is comparatively low due to both the

depthwise variation of the longitudinal scantlings, and the exponential decay of the hydrodynamic pressure. In ballast condition, the fatigue damage on the lower part of the side shell tends to be more uniform.

N

/

Fig. 12. Fatigue Damage Factor Fd Distribution

-Hydrodynamic wave pressure only - Full

Load - Head Seas - Envelope Fd= 6.0

Ab. No. 021 13/17

= nr

K

J-''I'll

Fig. 13. Fatigue Damage Factor Ed Distribution -Hydrodynamic wave pressure only - Ballast - Head Seas - Envelope Fd10.0

Figure 14. illustrates the fatigue damage distribution in beam seas. lt is shown that on the weather side of

the ship, the fatigue damage is significantly larger than on the lee side. The influence of the rolling

motion is also noticeable with the maximum fatigue damage zone extending over the waterline.

I

Ï

I

!

Fig. 14. Fatigue Damage Factor Fd Distribution

-Hydrodynamic wave pressure only - Fully

Loaded - Beam Seas - Envelope Fd=20.0 Figure 15. illustrates a typical distribution of fatigue damage due to the wave induced hull girder vertical bending moment.

Fig. 15 Fatigue Damage Factor Fd Distribution

Vertical Bending Moment- Fully Loaded Head Seas - Envelope Fd13.0

Figure 16 illustrates a typical distribution of fatigue damage due to

the wave induced

hull girder horizontal bending moment . The difference in fatigue damage on the side shell s due to the variation of S-N curves due to the difference in the SCF arising from the ratio of the flange thickness to flat bar thickness. Ab. No. 021

!I. ï

Fig. 16 Fatigue Damage Factor Ed Distribution Horizontal Bending Moment Fully Loaded -Beam Seas - Envelope Fd=0.15

7.2 100 Al Fatigue Wave Environment

Figure 17 shows the probability distribution of the

wave direction for the subject ship based on the 100

Al fatigue wave environment for large crude

oil

tankers. lt is shown that these distributions differ for

the fully loaded and ballast voyage. Due to a slight

dominance of the wave direction, the fatigue darage

is maximised on one side of the ship in the fully loaded voyage, and the other side in the ballast

voyage. 240

20'

180

.--Fully Loaded

--- Ballast

a Fully Loaded & Ballast

Fig. 17 Probability Distribution of Wave Direction

Figures 18,19 illustrate the distribution of fat!gue damage at midship and at a forward frame next to the

forward bulkhead of tank No. 1 respectively. lt is

shown that the fatigue damage in way of the midship section is dominated by the vertical bending moment at the deck and bottom, whisit the side shell is subject

to higher fatigue loading in way of the fully loaded

waterline. For the forward section, due to the larger amplitude of wave pressure as a result of the relative motion of the ship, and the reduction in the vertical bending moment amplitude, the side shell and outer

14/17 280 260

3o414

340 320

P4

40 60 111111 HhllilIllhli 11h11 Ill

bottom, especiaHy in way of the fully loaded waterline area is more prone to fatigue damage.

-n.

/=

Fig. 18 Fatigue Damage Factor Ed Distribution -Midship Section -Envelope Fd=1.0

Fig. 19 Fatigue Damage Factor Fd Distribution -Forward Section -Envelope Fd=1 .0

7.3 Alaska to Gulf of Mexico Wave Environment

o 340 320 300 280 260

iI

100 240 220 180 -.-- Fully Loaded

-- Ballast

e-- Fully Loaded & Ballast

Fig. 20 Probability Distribution of Wave Direction

Figure 20 shows the probability distribution of the

wave direction for the subject ship based the Alaska to Gulf of Mexico trading pattern . The distributions for the fully loaded and ballast voyage exhibit some degree of symmetry about the quartering seas axis.

Due to the dominance of the quartering wave

direction, t is expected that the fatigue damage will be higher due to the combined action of the vertical and horizontal bending moment, as well as the lateral motions inducing higher hydrodynamic wave pressures.

I

llIllIllIIlllIllll!I. II IlillIllhl lii

/

Fig. 21 Fatigue Damage Factor Ed Distribution

-Midship Section -Envelope Fd=1.0

Fig. 22 Fatigue Damage Factor Fd Distribution -Forward Section - Envelope Fd1 .0 Figure 21,22 illustrate the distribution of fatigue damage at midship and at a forward frame next to the

forward bulkhead of tank No. 1 respectively. lt is

shown that due to

the increased probability of

occurrence quartering seas, the fatigue damage due to the hydrodynamic pressure in way of the side shell

and the outer bottom is increased. lt should be

pointed out that this wave environment only applies to

a small percentage of the total

trading patterns

encountered by this

ship type, and

is given to highlight the significance of the wave environment to fatigue damage.

8. FATIGUE CONTROL PLAN AND CONSTRUCTION MONITORING

Since fatigue performance of ship structural details can be influenced by the workmanship standard, the fit-up and the alignment of the structural components,

t is essential that due attention is given during the

construction stages to ensure that the structure will be representative of the assumptions used in the fatigue

design assessment. To achieve this objective, a

Fatigue Control Plan ¡s developed at the plan

Ab. No. 021 15/17

approval stage, and the items to be considered are as follows

Identification of critical areas by the FEA Structural

Design Assessment procedure (SDA), and the

FDA procedure;

Marking of the critical areas and structural details on the ship plan;

Definition of the construction tolerances for the

critical structural details in terms of welding requirements, fit-up and misalignment tolerances. The erection sequence of the blocks is also to be specially considered iii order to minimise

misalignment, and locking of residual stresses

during assembly.

The fatigue control plan uses the fatigue life

results computed by the FDA procedure to

determine the level of inspection during construction.

During the newbuilding construction, the field

surveyor will draw particular attention to the critical areas highlighted by the Plan Approval Office, and the

Fatigue Control Plan. Enhanced levels of visual

inspection,

and NDEINDT may be

required at selected critical locations.

9. SUMMARY & CONCLUSIONS

Throughout the development of the ShipRight FDA

procedure, the objectives have been to provide a

flexible, reliable and total approach to assess the

problem of fatigue damage of ship structural details. The main features of the first principle Fatigue Design

Assessment procedure have been reviewed. The

procedure based on the structural spectral method of analysis has been implemented into a user friendly Window TM integrated software for design and

assessment purposes. The disadvantages associated with the application of the traditional maximum load approach to the determination of the long term stress

range spectrum which have prompted the

development of the subject procedure have been

summarised. Novel direct calculation features such as

the S-N Curve Expert, the computation of wave

induced loads by parametric expressions, the use of a voyage simulation procedure have been outlined, and typical application examples have been given to

illustrate the application of the procedure.

In summary, in

order to

attain and maintain a

satisfactory fatigue performance of ship structural

details, it is essential that a global approach giving

due consideration to the design, construction and ¡n

service stages of the life of structural details be applied. This has been achieved by

ensuring an adequate level of detail design is

performed by providing the Structural Detail

Design Guide;

ensuring that adequate structural design

concepts and sound analysis techniques are

used by providing a first principle fatigue design assessment tool supported by an ongoing research programme of theoretical and

experimental work;

ensuring that the workmanship, and construction standards are performed to a satisfactory level, and that the fatigue strength can be maintained by the provision of enhanced survey procedures, and the application of a hull condition monitoring system.

10. REFERENCES

British Maritime Technology, Global Wave Statistics,

1985.

BS 5400, Part 10, 1980. Code of Practice for Fatigue.

Steel, Concrete and Composite Bridges. British

Standard Institution.

ECCS - Technical Committee 6 - Fatigue Recommendations for Fatigue Design of Steel Structures, 1985

Ferguson J.M., Violette F.L.M, Some Effects on Ship Structural Design created by the Increased Application of Higher Tensile Steels, Proc. IMSDC 91, Kobe, Japan.

liS/lIW Doc. 700-82, Welding in the World, No 20.7/8,

1982

KSRI, Load Spectra for Ship Structure Fatigue

Evaluations, Restricted, 1992

Lloyd's Register, Paper presented at the Royal

Institution of Naval Architects, London, 1913

Lloyd's Register, ShipRight FDA - Structural Detail Design Guide, 1994

Lloyd's Register, ShipRight FDA - Fatigue Design Assessment - Procedures Manual, 1994

Lloyd's Register, ShipRight SDA - Structural Design Assessment - Procedures Manual, 1994

Lloyd's Register, ShipRight HCM - Hull Condition Monitoring - Procedures Manual, 1994

Lloyd's Register, ShipRight CM - Construction Monitoring - Procedures Manual, 1994

Lloyd's Register, Comparative Fatigue Damage Analysis of a 280,000 Dwt VLCC, September 1992 -Restricted

Offshore Installations : Guidance on Design and Construction. New Fatigue Design Guidance for Steel

Welded Joints in Offshore Structures. UK Den,

August 1983.

Violette ELM, The Effect of Corrosion on Structural Detail Design, Int. Conf. Marine Corrosion Prevention,

Royal Institution of Naval Architects, London, Oct.

1994

Wirshing P.H, Fatigue under wide band random

Stresses using the rainflow method, Journal of Engineering Materials and Technology, ASME, 1977 Wirshing P.H, Chen Y.N., Consideration of Probability - Based Fatigue Design for Marine Structures, Proc. Marine Structural reliability Symposium, Arlington, VA,

1987. 11. NOMENCLATURE

I

ZU.' zzz zYY s ¡

f(x)

K8, C, P,(t) H/i) n p(S)

a

so S, tS N(S) n(S) B p/mm I,SO,B) m K E T Ab. No. 021

SecQnd moment of area Section modulus about flange Hull girder section modulus about ZZ Hull girder section modulus about YY Stiffener spacing

Effective span

Bending moment shape function at

critical location x from span point

Bias stress concentration factor loadcase i

Structural influence coefficient of the ith load process P,(t)

Load process i

Stress process ¡ complex form

Total number of load processes I ship

motions influence parameters Stress range probability function

Standard deviation of the stress process

S-N curve stress range at stress cycles

Nominal stress,stress range

Number of allowable stress cycles at

stress range S,

Number of stress cycles with stress range S for the given seastate

Expected number of stress cycles in the given seastate

Modelling bias for the stress prediction model

Correction factor for multi-linear S-N curve

Slope of selected S-N curve Intercept of selected S-N curve Spectral bandwidth

Seastate duration

Pk =p(kijl)

17/17

U Mean zero crossing frequency

P, = p(i

jkl)

probability ith loading condition

PI =p(i

iki)

probability jth ship to wave heading

probability kth ship speed

p, = p(l ¿1k)

probability Ith seastate (H1I3,TZ)

SCF ship structural detail 50F standard detail

Knom Intercept nominal S-N curve

Kh, Intercept reference S-N curve

hv Hot spot stress range

m S-N curve inverse slope