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Marine Structures 15 (2002) 113
Comparative fatigue strength assessment of a
structural detail in a containership using various
approaches of classification societies
W. Frickea,
*, W. Cuib
, H. Kierkegaardc
, D. Kihld
, M. Kovale
,T. Mikkolaf, G. Parmentierg, M. Toyosadah, J.-H. Yooni
aTU Hamburg-Harburg, AB 3-06, L .aammersieth 90, D-22305 Hamburg, GermanybSchool for Naval Architecture & Ocean Engineering, Shanghai Jiao Tong University, Shanghai 200030,
ChinacOdense Steel Shipyard Ltd., P.O. Box 176, DK-5100 Odense C., Denmark
dNaval Surface Warfare Center, Carderock Division, West Bethesda, MD 20817, USAeKvaerner Krylov Maritime Ltd., 196158 Moskovskoye shossee, 44 St. Petersburg, Russia
fVTT Manufacturing Technology, P.O. Box 1705, 02044 VTT, FinlandgBureau Veritas, 17 Bis Plau des la Defense 2, Paris la Defense 92077 Cedex, France
hDepartment of Marine Engineering, Kyushu University, 6-10-1 Hakozaki, Higashi-ku, Fukuoka 812-8581,
JapaniSamsung Heavy Industries, 530 Jangyung-ri, Shinhyun-up, Koje-City, Kyungnam 656-710, South Korea
Received 18 October 2000; received in revised form 23 March 2001; accepted 13 May 2001
Abstract
A comparative study on fatigue strength assessment procedures used by the classification
societies has been performed by Committee III.2, Fatigue and Fracture, of the International
Ship and Offshore Structures Congress (ISSC2000). A pad detail on the longitudinal coaming
of a Panamax container vessel was selected as an example. This detail was chosen because ofthe well-defined loading due to hull girder bending. Large differences in predicted fatigue lives
were found, ranging from 1.8 to 20.7 years. The spreading of results is attributable to
assumptions regarding loads, local stress determination and SN curve. For comparison, a
direct calculation of loads using the spectral method was performed. Also this calculation
showed a relatively short fatigue life of 5.3 years, although the structural detail is considered
not to be prone to fatigue failures. r 2001 Elsevier Science Ltd. All rights reserved.
Keywords: Fatigue strength assessment; Welded detail; Classification society; Container ship; Fatigue life
*Corresponding author. Tel.: 49-40-428-32-3148; fax: 49-40-428-32-3337.
E-mail address: [email protected] (W. Fricke).
0951-8339/02/$ - see front matter r 2001 Elsevier Science Ltd. All rights reserved.
PII: S 0 9 5 1 - 8 3 3 9 ( 0 1 ) 0 0 0 1 6 - 8
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1. Introduction
During the past decades, a great number of fatigue failures occurred in ship
structures, particularly in areas made of higher tensile steel. In the early 1990s,following the occurrence of cracks in side shell longitudinals of very large crude
carriers (VLCCs) after only a few years in service [1], all major classification societies
issued or revised rules and guidelines for explicit fatigue analyses. Failures in side
shell longitudinals were generally used for calibrating the load assumptions and
fatigue damage calculations.
It is quite natural that several investigations in the following years were focussed
on the complex load process at the ships sides (e.g., [24]) and on the fatigue
resistance of the affected structural details (e.g., [5]). Additional studies
were performed to review various procedures of the classification societies for
fatigue strength assessment, which showed large differences in the assumptions and
results [6].
In order to further investigate the different procedures for fatigue strength
assessment of ship structural details, the members of ISSC Committee III.2, Fatigue
and Fracture, decided to perform another comparative study in which the
procedures of the different classification societies were applied by the individual
participants. It was decided, however, not to investigate structural details at the
ships side again for following reasons:
* Side shell longitudinals are not the only ship details prone to fatigue. Longitudinalmembers in the upper and lower flanges of the hull girder require at least the same
attention given side shell longitudinals in view of the cyclic vertical and horizontal
wave bending moments. Recent casualties of ships breaking apart (MSC Carla,
Flare and Erika) underline the possible consequences of such failures,
although the contribution of fatigue to the cases mentioned still needs to be
clarified.* The load process at the ships sides is very complex, being characterized by
uncertainties in the combination of local and global hull girder loads and in non-
linear effects particularly related to local pressure loads. These uncertainties
consequently affect the assumptions contained in the different fatigue assessment
procedures and, hence, the spread of results.
Therefore, a detail in the upper longitudinal structural members was selected. The
loading for this detail is relatively well defined in the rules of the classification
societies. With respect to criticality, a detail in a relatively large container vessel has
been considered, where the use of higher-tensile steel and the presence of high tensile
stillwater stresses increase the failure probability. Frequent recommendations to pay
attention to outfitting details at the coaming of container vessels indicate that fatigue
cracks have occurred there.
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2. Selected example
The example ship chosen is a 3800 TEU Panamax container vessel. The principal
particulars of the vessel are given as follows:
Length between perpendiculars, Lpp 242.000 m
Rule length, L=0.97LWL 234.740 m
Breadth moulded, B 32.250 m
Depth moulded, D 21.300 m
Draught moulded, T 14.000 m
Block coefficient, Cb 0.670
Maximum service speed, v 24.000 kn
Web frame spacing 3.144 m
Max still water bending moment, hogging 210000 t m
A welded pad detail on top of the longitudinal hatch coaming bar, where the hatch
covers are supported for vertical loads, was selected for the analysis. The
longitudinal elements in the midship section and the pad detail are given in Fig. 1.
The vessel has higher-tensile steel, HTS 36, in the upper part, noted AH and EH in
Fig. 1, with the rest of the ship made of mild steel, noted A. The pad detail for the
comparative study is located at mid-ship, midway between the fore and aft end
transverse coaming. The effects of torsional loading and stresses due to warping
distortion are therefore disregarded. The shape of the pad detail was designed toreduce the notch effect in the longitudinal direction. The size of welding is specified
by a throat thickness of 10.0 mm. Cracks are expected to initiate at the weld toes on
the top of the coaming plate.
The midship section properties of the hull grider were calculated and given as:
Moment of inertia about the vertical axis 659.765 m4
Moment of inertia about the horizontal axis 258.740 m4
Height of neutral axis above base line 8.496 m
Section modulus at coaming (rule z=22.781 m) 18.114 m3
3. Applied rules and guidelines
The fatigue analyses were performed according to current rules and guidelines
from the following classification societies:
* American Bureau of Shipping (ABS)* Bureau Veritas/Registro Italiano Navale (BV/RINA)* Det Norske Veritas (DNV)* Germanischer Lloyd (GL)* Korean Register (KR)
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Fig. 1. Midship section with longitudinal elements for a Panamax container vessel. Details of the pad
plate on the longitudinal hatch coaming bar are also seen.
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* Lloyds Register of Shipping (LR)* Nippon Kaiji Kyokai (NK)* Russian Register of Shipping (RS)
Direct calculations using the spectral method and statistical approach were also
performed.
The SN approach was generally applied, mostly in a simplified way, by
prescribing the shape of the stress histogram together with the number of load cycles.
The stresses used were either nominal stresses in the top of coaming, hot-spot stresses
at the critical weld toe, or else notch stresses. The hot-spot stress concentration
factor (SCF) was either calculated using a local finite element model or estimated on
the basis of tabulated SCFs, while the notch factor for the weld toe was taken from
the rules or guidelines, where applicable. Special aspects of the different rules and
guidelines are given in the following.
3.1. American Bureau of Shipping (ABS)
ABS supports both the nominal and the hot-spot stress approaches in their
simplified fatigue strength assessment method [7]. The fatigue stress ranges are
assumed to follow a Weibull probability distribution. For the present pad detail, the
design stress range is the same as that specified by IACS UR S11 [8]. The Weibull
shape parameter is a function of both ship length and location, and under some
assumptions, is estimated to be 0.81. The effect of mean stress has been ignored. DEn
SN curves were used to describe the fatigue strength of the details. In order to
account for corrosion, a net ship concept was used together with a stress reduction
factor of 0.95. For more detailed description using ABS approach to analyze this pad
detail, please refer to [9].
3.2. Bureau Veritas/Registro Italiano Navale (BV/RINA)
By the BV rules [10], which are comparable to RINA rules, two load cases should
be taken into account: half-time in head sea conditions and half-time in oblique sea
conditions. The stresses are based on the rule bending moment specified by IACSUR S11 [8]. The stresses and SN curve are based on the notch stress approach. The
hot-spot stress concentration factor was obtained from a finite element calculation,
while the notch stress concentration factor was obtained from a table depicting the
type of weld and quality of welding. The calculation included the stress component
from the IACS head sea vertical bending moment, transferred to a probability of
105 using a Weibull distribution with a shape parameter of 0.943 combined with
a horizontal bending moment. Another combination of vertical and horizontal
bending moments provided the oblique seas component. A reduction of life for
thickness above 16 mm is included as well as a reduction of stress amplitude when the
notch stress amplitude corresponding to a probability of 10
5 is above the yieldstress. The part of stress above the yield stress was weighted with a factor of 0.6
(compressive stress factor).
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3.3. Det Norske Veritas (DNV)
The procedure from DNV [11] for fatigue strength assessment is based on the hot-
spot stress approach. The stress concentration factor was obtained from a table ofstandard details which contains doubler plates. However, the stress range was
multiplied by a stress concentration factor for the weld (notch stress) before entering
the SN curve. The calculations only include the stress component from the vertical
bending moment, per IACS UR S11 [8], transferred to a probability of 104 by the
Weibull distribution with a shape parameter of 0.93. The stress range was further
reduced by a factor of 0.80 to account for worldwide operation. A reduction in life
for a plate thickness above 22 mm was included.
3.4. Germanischer Lloyd (GL)
The simplified approach of GL [12] assumes a long-term histogram of stress ranges
characterised (a) by a maximum stress range of 75% of the IACS UR S11 [8] design
value for vertical hull girder bending in order to account for the effect of varying
loading conditions and worldwide service, (b) by a total of 50 million load cycles,
and (c) by a Weibull shape parameter x 1: Both, the hot-spot as well as the
nominal stress approach are supported; the latter based on the recommendations of
the International Institute of Welding [13]. Here, a relatively conservative detail
category of FAT 50 is assumed, which may be substantially increased in case of a
reduced weld flank angle. Mean stress is taken into account by a correction factor.
3.5. Korean Register (KR)
The guidance for the fatigue strength assessment of ship structures of KR [14] is
based on the hot-spot stress approach. Since the mean stress effect is to be considered
in the case of compressive mean stress only, it was ignored in this calculation. The
thickness effect is to be considered for plate thicker than 22 mm. In the calculation of
the SCF, if the weld is not considered in the finite element model, the stresses at the
distance t/2 and 3t/2 from the weld toe are determined by the Lagrange interpolation
using the element stresses in the region. The hot-spot stress was then obtained bylinear extrapolation to the weld toe using the stresses determined at distances of t/2
and 3t/2 from the weld toe.
3.6. Lloyds Register of Shipping (LR)
From Lloyds Register, the ShipRight Fatigue Design Assessment Level 3
procedure [15] was applied. FDA Level 3 is a spectral approach where, in this case, a
scaled hull form and weight distribution from a similar ship was used for generation
of the vertical bending moment response amplitude operator from 2D strip theory. A
typical container trade pattern was assumed and a 20-year simulation period with27% non-sailing days was chosen. The fatigue stress was obtained from a detailed
FE-model composed of 4-noded shell elements, representing the geometric stress
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(hot-spot stress) with some embedded notch stress effects. The analysis was
performed in-house by LR.
3.7. Nippon Kaiji Kyokai (NK)
NK [16] guidance supports the hot-spot stress approach. The stresses are normally
based on direct calculation using a spectral method and statistical approach. The
hot-spot stress is determined with the aid of a zooming FE-model with element sizes
of approximately t t in the critical area. Several standard SN curves for each
welded joint were prepared and the effects of mean stress with residual stress were
considered. In the present case, the stress range was based on 100% of the rule
bending moment according to IACS UR S11 [8] since direct calculation using the
spectral method was not carried out. The SN curve for non-load carrying type fillet
welded joints in NK guidance was used.
3.8. Russian Register of Shipping (RS)
RS [17] supports both the nominal stress approach for standard details and the
hot-spot stress approach for other details, including the present pad detail. For this
calculation, only the stress component from the vertical bending moment was
included and transferred to a probability level of exceedance 103 by a Weibull
distribution with a shape parameter of 0.88. Mean stress was taken into account by a
correction factor. Corrosion effect may be considered.
4. Calculated results
The calculated results have been derived independently by the responsible
participants and, in most cases, in co-operation with the respective society. The
details of the applied approaches are very different, ranging from the nominal stress
approach to spectral load analysis and local FE analysis of the patch detail.
However, the collected results are assessed by comparing the loading, the local stress,
and the fatigue life. The results are presented in Table 1.The third column in Table 1 (after the classification society and type of approach)
contains information on the loading, i.e., the stress histogram as explained in the left-
hand diagram of Fig. 2. The highest stress range Dsmax was generally based on
nominal stresses in order to allow a better comparison. Although the loading was
governed by vertical hull girder bending only, significant differences can be observed
in Dsmax assumed, ranging from 199 to 319 MPa.
Also, the Weibull shape parameter, as well as the number of stress cycles, vary to a
certain extent. It is interesting to note that the Weibull shape parameter apparently
decreases with increasing maximum stress range, indicating some calibration behind
the load assumptions.The stress concentration factors in the centre column are partly derived from finite
element analyses with models as shown in Fig. 3. The hot-spot SCF derived varies
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Table 1
Comparison of results from comparative study
Rules and
guidelines
Type of stress
approach
Histogram of nominal stress ranges SCF
(hot spot/weld)
Design S
Dsmax (MPa) nmax x sma(MPa) DsR (MP
ABS [7] Nominal 318.7 5107 0.81 F F/F 49.9
Hot spot 318.7 5107 0.81 F 1.736/F 80.3
BV/RINA [10] Notch 278b 2.8107 0.943 F 1.63/1.84 142.6
136c 2.8107 0.943 F
DNV [11] Notch 233.0 6.65107 0.93 F 1.47/1.5 142.2
GL [12,18] Nominal 209.2 5107 1.0 104.9 F/F 50
Hot spot 209.2 5107 1.0 104.9 1.9/F 110f
KR [14] Hot spot 278.8 5.64107
0.943 1.66/F 91.3 LR [15] Hot spotg 210e 5.7107 F F 1.81g 124g
NK [16] Hot spot 281.5 108 1.0 108.9 2.15/F 95
RS [17] Hot spot 199.0 5107 0.88 113.8 1.80/F 100
aMean stress only given if it affects life.bPart from head seas.cPart from oblique seas.dEstimated from usage factor.eRead from 108 probability of exceedence.fIncluding 10% increase for exact stress analysis.gHot spot with some embedded notch stress effects.h
Life 13.2 years for corrosion wastage by 0.5% of section modulus every year.
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considerably from 1.47 to 2.15. Table 2 gives some additional information about thefinite element calculations, i.e., element types and sizes used and the stress evaluation
technique applied. The assumed notch factors vary as well.
The SN data in Table 1 are further explained in the right-hand diagram of Fig. 2.
The reference stress range DsR is related to the type of stress mentioned in the second
column. Again, the numbers vary considerably, while the knuckle points of the SN
curves and the slope exponents show less scatter.
As a consequence, the resulting fatigue lives in the last column differ by a factor of
more than ten, i.e. from 1.8 to 20.7 years. The low fatigue lives are somewhat
surprising because the opinion of some experienced designers was that the detail
analyzed is very good and should not show any problems in service. Also, the serviceof container vessels, with even worse details for more than 20 years, has
demonstrated a better fatigue life than predicted here.
Fig. 2. Definitions for Table 1.
Fig. 3. Typical FE-model for calculation of hot-spot stress concentration factor.
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Table 2
Calculation of hot-spot SCFs
Guidelines Element at hot-spot Stress Evaluation
Elem.
type
Size
long./transv./vert.
Weld
modeling
Stress
type
Extrapolation to
surface hot-s
ABS Solid20 30 3030mm3 Triang. v. Mises Linear Linea
BV/RINA Solid8 30 30 10mm3 No Principal Linear Linea
GL [18] Solid20 30 30 30mm3 Prism. Perpend. Linear LineaKR [14] Solid8 30 30 30mm3 No Principal Linear Linea
LR [15] Shell 30 30mm2 Shella Normalb Linear None
NK [16] Solid8 30 30 15mm3 Prism. Principal Linear Linea
RS [17] Solid20 10 10 7.5 mm3 Prism. Perpend. Nonlinear Linea
a45 mm thick elements.bStress component normal to crack plane within 7401 from principal crack plane.
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5. Direct calculation using a spectral method and statistical approach
The fatigue life of the coaming pad detail was further predicted using a linear
spectral method and long-term wave statistics. The analysis is briefly described in thefollowing. A detailed description is given in [19]. The following assumptions were
made:
* hydrodynamic calculations based on strip theory for 50 wave frequencies between
0.2 and 2.5 rad s1
* ISSC wave spectrum and cosine-squared spreading function* Long-term wave data from Global Wave Statistics [20] from relevant North
Atlantic areas* duration at sea 85% of 20 years
* speed profile (24 kn up to Hs=5 m, 20 kn up to 7 m, 15 kn up to 9 m, 10 kn up to11 m and 0 kn above 11 m)
The resulting long-term distribution of nominal stress ranges is characterized by a
Weibull distribution with the following data:
Ds=274.1 MPa
n =5.37 107
x =1.0205
The first two values agree quite well with the assumptions of IACS R 56 [21],whereas the shape parameter x of the Weibull distribution obtained by a linear least-
squares procedure is much higher than that proposed by IACS. Assuming the
rainflow correction factor in [22] as well as FAT 50 together with the associated
design SN curve according to IIW [13], a fatigue life of 5.3 years is obtained.
The effect of several input parameters on fatigue life was investigated in a
parametric study, which is summarized in Table 3. None of these parameters, i.e. sea
area, ships speed, rainflow correction and SN curve with or without change in
Table 3
Results of direct calculations with parameter variation
Sea area Speed Knuckled
SN curve
Rainflow
correction
Dsmax (MPa) nmax (107) x Fatigue life
(yr)
N. Atlantic Profile Yes Yes 274.1 5.37 1.0205 5.3
N. Atlantic Profile No No 274.1 5.37 1.0205 4.0
N. Atlantic Profile No Yes 274.1 5.37 1.0205 4.7
N. Atlantic Profile Yes No 274.1 5.37 1.0205 4.5
Northern N. A. Profile No Yes 290.8 5.18 1.0677 2.9
Pacific Profile No Yes 255.0 5.55 0.9991 5.3
N.A.+Pacific No Yes 267.6 5.46 1.0040 4.6
N. Atlantic Zero No Yes 274.6 5.62 0.9663 4.8N. Atlantic 24 kn No Yes 298.3 5.37 0.9893 3.5
N.A.+Pacific Profile Yes No 267.6 5.46 1.0040 6.2
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slope, had a large effect on the calculated fatigue life, ranging from 2.9 to 6.2 years.
The only major effect found was when the maximum significant wave height was
reduced, i.e., the wave climate became more moderate.
6. Conclusions
Approaches for fatigue strength assessment have been developed during the recent
years and implemented in several classification societies rules and guidelines. These
have been applied, within a comparative study, to the fatigue assessment of a pad
detail on the coaming of a Panamax container ship. Although the loading is quite
well defined in the upper part of the hull girder, the predicted lives vary considerably;
between 1.8 and 20.7 years. It has been shown that the high scatter is attributable todifferent assumptions regarding load effects as well as local stress analyses and SN
curves. In most cases, a simplified approach has been applied using standardized
stress histograms (Weibull distributions) together with the SN approach. The type
of stress varies between nominal, hot spot and notch stress, each coupled with certain
uncertainties regarding the detail classification or determination of stress concentra-
tion factors.
In addition, a direct calculation of the loading has been performed. It showed a
relatively short fatigue life of 5.3 years, assuming the North Atlantic wave climate
and a certain speed profile. In a parametric study several assumptions and input
parameters have been modified, resulting in calculated fatigue lives, ranging from 2.9to 6.2 years, which is in the lower end of the results based on the rules and guidelines
of the classification societies. In conclusion, the situation appears to be
unsatisfactory, particularly in view of the large scatter of result, but also because
of the fact that designers regard the detail as unproblematic with respect to fatigue.
This indicates that several procedures including the direct load calculation yield
overly conservative results.
Acknowledgements
The comparative study was performed as part of the work of Committee III.2,
Fatigue and Fracture, of the International Ship and Offshore Structures Congress
(ISSC) during the working period 19972000. The authors wish to thank those
members of the classification societies, who have actively contributed to the
calculations, as well as the remaining members of the Committee, who have
contributed valuable comments to the study.
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