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International Association of Classification Societies
(IACS)
FSA of Bulk Carriers
Fore-end Watertight Integrity
Annex 4
Hatch Cover Failure Scenarios
20010225_Annex 4Hatch Cover Failure Scenarios.doc,28 February, 2001
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Table of Contents
1
SUMMARY
1
2
INTRODUCTION
4
2.1
Background
4
2.2
Objective
4
2.3
Scope
4
2.4
Overall approach of study
5
2.5
Description of base case design
6
3
HAZARDS RELATED TO HATCH COVERS AND COAMINGS
7
4
RISK CONTRIBUTION FROM HATCH COVER FAILURES
9
4.1
Casualty data
9
4.2
Overall risk contribution based on casualty data
9
4.3
Casualty data broken down on location of water ingress and severity
11
4.4
Risk contribution from No. 1 hatch covers on capesize bulk carriers
12
5
COST EFFECTIVENESS ANALYSIS OF RISK CONTROL OPTIONS
13
5.1
General remarks
5.1.1
Risk model
1.1.2
Cost effectiveness calculations
1.1.3
Assumptions
13
13
14
14
1.2
New-building requirements for hatch cover design pressure
15
1.3
RCO No. 1; Replacement of ILLC 66 hatch covers
17
1.4
RCO No. 2; Replacement of IACS UR S21 hatch covers
19
1.5
RCO No. 3 Hydraulic hatch cover closure
20
6
DISCUSSION
23
7
RECOMMENDATIONS
24
8
REFERENCES
24
APPENDIX 1
INFORMATION FROM MACGREGOR
APPENDIX 2
BREAK-DOWN OF CASUALTY DATA
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APPENDIX 3
HATCH COVER HAZARDS
APPENDIX 4
LOADS
HATCH COVERS EXPOSED TO GREEN SEA IMPACT
APPENDIX 5
IACS UNIFIED REQUIREMENTS S21
APPENDIX 6
ANALYSIS OF HATCH COVER CAPACITY
APPENDIX 7
ANALYSIS OF HATCH COVER RELIABILITY
APPENDIX 8
MARGINAL COST EFFECTIVENESS
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1
1
SUMMARY
The present annex is a part of the IACS study Formal Safety Assessment of Bulk Carriers,
Fore end watertight integrity.
The objective of this annex is to evaluate, by Cost Effectiveness Analysis, risk control options
aiming at preventing water ingress scenarios due to hatch cover failures on bulk carriers.
In the present evaluation, a representative capesize bulk carrier was studied to quantify the
effects and costs related to hatch cover risk control options. A capesize carrier was selected
primarily because it was, based on experimental results (MSC72/4/1/Add.1, MSC 72/4/1),
suspected that the design loads for these bulk carriers might be in the lower range.
The scenarios considered in this annex consist of the following events:
1. Significant water ingress through an opening in No. 1 hatch cover. Given an opening of
the size of a hatch cover, the cargo hold may be completely flooded within matter of
minutes, see e.g. DETR (1998).
2. In some of the cases there are progressive flooding of cargo holds, leading to total loss of
ship and in most cases many fatalities.
3. In the remaining cases, the flooding is limited, resulting in serious casualty and not total
loss, and few, if any, fatalities.
In addition to evaluating the marginal cost effectiveness for hatch cover design loads, the
following risk control options were studied:
1. IACS UR S21 for hatch covers replacing ILLC 66 implemented on existing ships
2. Hatch covers designed to a 30% increase in IACS UR S21 design loads to replace IACS
UR S21 hatch covers on existing ships.
3. Hydraulic hatch cover closure system for No. 1 hatch cover
The cost effectiveness of the different risk control options have been assessed by their Gross
and Net Costs of Averting a (statistical) Fatality (CAF), defined as:
∆C
GrossCAF =
∆R
∆C − ∆B
NetCAF =
∆R
where ∆C is the total costs, ∆B is the economic benefits, and ∆R is the number of fatalities
averted by the risk control option.
MSC 72/16 recommends a decision criterion of US$ 3 million for risk control options
involving reduction in both number of fatalities and injuries, and US$ 1.5 million for risk
control options involving a reduction in number of fatalities only. The latter is likely to be the
case for casualties related to hatch covers. Higher values may be justified for risks that are just
tolerable, and MSC70/WP.12, paragraph 30, refer to criteria in the range from US$ 1 to 8
million.
Table 1 gives the results for the different hatch cover risk control options studied.
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Table 1
2
Cost Effectiveness results for hatch cover risk control options.
RCO description
IACS UR S21 hatch covers to
replace ILLC 66 hatch covers
on existing capesize carriers
10 year old ships
15 year old ships
20 year old ships
Hatch covers designed to a 30%
increase in IACS UR S21 design
loads to replace IACS UR S21
hatch covers on existing ships.
10 year old ships
15 year old ships
20 year old ships
Hydraulic hatch cover closure
system on No. 1 hatch cover
10 year old ships
15 year old ships
20 year old ships
New-buildings
30% increase in IACS UR S21
design loads for new ships
∆R
∆C
(fatalities
(US$)
averted per
ship)
∆B
(US$)
5.16E-02 - 105,000 – 175,000
1.71E-01
3.44E-02 - 105,000 – 175,000
1.14E-01
1.72E-02 - 105,000 – 175,000
5.71E-02
Gross CAF Net CAF
(US$
(US$
million)
million)
27,700 92,000
22,400 74,300
13,821 45,800
0.6 - 3.4
0.1 - 3.2
0.9 - 5.1
0.3 - 4.9
1.8 - 10.2
1.0 - 9.9
2.00E-03
1.34E-03
6.68E-04
120,000 – 200,000
120,000 – 200,000
120,000 – 200,000
1,100
870
540
60 – 100
90 – 150
180 – 300
60 - 100
90 – 150
180 - 300
1.95E-02
1.30E-02
6.49E-03
1.46E-02
-
58,000
58,000
58,000
58,000
-
10,500
8,450
5,210
8,550
-
2.97
4.46
8.91
3.97
86
2.43
3.81
8.11
3.39
When evaluating the risk control options involving retrofitting existing bulk carriers with
reinforced hatch covers, low cost estimates were used, not accounting for costs related to
design, strengthening of support structure, and replacing wheels, rails, wheel lifters, fittings,
and driving equipment. Colman (2000) described the cost of reinforcing No. 1 and 2 hatch
covers on existing capesize bulk carriers as unlikely to exceed £150,000, or US$225,000, and
as unlikely to exceed £100,000, or US$ 150,000, for new-buildings. This is in excess of the
costs used in the present evaluation, and would make the Gross and Net CAF estimates
higher. The evaluations do not give robust conclusions for the average bulk carriers with
hatch covers designed according to ILLC 66. However, it is possible that the hatch cover
capacity varies to such an extent that the replacement of hatch covers on a subset of the fleet
could be justified. If a screening of the relevant bulk carriers could identify individual ships
with particularly high hatch cover risks, then the replacement of their No. 1 and 2 hatch
covers could be justified from a cost-effectiveness perspective.
Risk control option No. 3 implying an hydraulic hatch cover closure system on No. 1 hatch
cover for existing ships and new buildings has been evaluated to have a Gross and Net CAF
in excess of the decision criterion.
The casualty data indicates that the capesize carriers are more at risk for hatch cover failure
than the smaller ships. A reason for this may be their lengths being comparable to the wave
lengths in severe sea states, giving increased sea loads on the hatch covers and hence
increased probability of water ingress. The handysize carriers appear to be more exposed than
the handymax and panamax, probably due to lack of damage survivability. In many cases, the
flooding of one cargo hold is sufficient for these ships to suffer from lack of buoyancy. The
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3
panamax bulk carriers may have an order of magnitude lower frequency of total loss due to
hatch cover failure compared to the average risk for bulk carriers. Given hatch cover
strengthening of panamax, handysize, and handymax carriers, the relative risk reduction is
expected to be similar to the risk reduction estimated for the capesize carriers. The absolute
risk reduction is expected to be 50% or less for the smaller bulk carriers compared to the
capesize carrier. The costs of strengthening hatch covers are also expected to be lower. Hence,
the cost effectiveness for smaller bulk carriers is expected to be of the same order of
magnitude as the cost effectiveness estimated for the capesize bulk carrier, giving similar
recommendations.
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2
2.1
4
INTRODUCTION
Background
The present annex is a part of the IACS study Formal Safety Assessment of Bulk Carriers,
fore end watertight integrity.
Serious concerns have been expressed about the safety of bulk carriers for some time
particularly following a spate of losses in the early 1990s. IMO prompted an international
programme of research and development culminating in the 1997 SOLAS Conference on
Bulk Carrier Safety. Revised rules and standards relating to the design and operation of bulk
carriers were included in a new Chapter XII to the 1974 International Convention for the
Safety of Life at Sea (SOLAS 74).
These measures were rightly targeted at those vessels seen as being most at risk. However,
concerns have been expressed that these safety improvements are piecemeal. The possibility
of unidentified or not fully identified hazards in bulk carriers has been acknowledged. It has
further been acknowledged that an overarching review of the many inter-related facets of bulk
shipping safety is required to assist IMO in the development of the international regulatory
framework.
2.2
Objective
The objective of this annex is to evaluate, by Cost Effectiveness Analysis, risk control options
aiming at preventing water ingress scenarios due to hatch cover failures on bulk carriers.
2.3
Scope
The scenarios considered in this annex consist of the following events, see Figure 1:
1. Significant water ingress through a hatch cover opening. Given an opening of the size of a
hatch cover, the cargo hold may be completely flooded within matter of minutes, see e.g.
DETR (1998).
2. In some of the cases there are progressive flooding of cargo holds, leading to total loss of
ship and in most cases fatalities.
3. In the remaining cases, the flooding is limited, resulting in serious casualty and not total
loss, and few, if any, fatalities.
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Escalation ?
YES
5
Total loss of
ship
Substantial water ingress in
forepeak or No. 1 cargo hold
NO
Serious casualty
OR
Side shell
failure
Figure 1
Hatch
cover
failure
Failure of deck
fittings
Scenario investigated in the present annex.
In the present evaluation, a representative capesize bulk carrier was studied to quantify the
effects and costs related to hatch cover risk control options. A capesize carrier was selected
primarily because it was, based on experimental results, suspected that the design loads for
these bulk carriers might be in the lower range.
In addition to evaluating the marginal cost effectiveness for hatch cover design loads, the
following risk control options were studied:
1. IACS UR S21 for hatch covers replacing ILLC 66 implemented on existing ships.
2. Hatch covers designed to a 30% increase in IACS UR S21 design loads to replace IACS
UR S21 hatch covers on existing ships.
3. Hydraulic hatch cover closure system for No. 1 hatch cover.
2.4
Overall approach of study
In order to be able to cover the totality of the complex problem to a sufficiently detailed level,
the work was carried out by a multi-disciplinary team covering:
• environmental loads,
• hydrodynamics,
• structural analysis,
• structural reliability,
• FSA.
In order to be able to evaluate the risk control options, the analysis was divided into the
following steps:
•
•
•
•
•
analysis of historical casualty data to establish an average annual probability of hatch
cover failure and the consequences given hatch cover failure, see Section 4
analysis of environmental loads and dynamic response, in order to establish a realistic
probability distribution for the loading for No. 1 hatch cover, see Appendix 4.
formulation of hatch cover capacity, see Appendix 6
structural reliability analyses to estimate the probability of failure for each hatch cover
design, see Appendix 7
cost effectiveness analyses, given different costs of the risk control options, different
probabilities of failure, and hence different risk, see Section 5.
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The structural reliability analysis was carried out for three different hatch cover designs. The
three designs were established by deterministic design analyses:
1. Initial design, according to ILLC 66.
2. Modified design, according to UR S21.
3. Modified design, according to UR S21, but with a 30% increase in load.
The reliability analysis uses probabilistic models for both load and capacity. The probabilistic
loading model is the same for all three designs. The design modifications are reflected in the
distribution for the capacity. The relative difference in probability of failure for the three
different designs is used as input to the cost effectiveness analysis.
2.5
Description of base case design
A typical hatch cover configuration for the first cargo hold of a capesize bulk carrier has been
selected as base case. The length of the hatch cover is 15.59m and the breadth is16.68m. The
structural system of the hatch consists of a top plate with longitudinal L-stiffeners, transversal
and longitudinal girders. Typical dimensions are:
−
Plate thickness, 8mm and 10mm near the centre
−
Stiffeners, L profile 100x75x7, spacing 550mm
−
Transversal girders, 7 in total, height approximately 850mm at centre and ends,
450mm in between. Typical web thickness of 7mm, with lower flange of 240x25mm2
and 100x15 mm2 for the large and small size girders respectively. Girder spacing 2700
mm.
−
Longitudinal girders around 850mm height, web thickness of 7mm and lower flange
of 375x40 mm2 for the two most heavily loaded ones. Girder spacing 2750mm.
A finite element plot of the hatch cover is included in Figure 2, seen from the side underneath.
The green and red colours are used the surface plate of 8 mm and 10 mm thickness
respectively. Blue colour is used for the webs of longitudinal and transversal girders, and
yellow is used for the vertical surfaces that rest on the coaming. Please note that the lower
flanges of the girders and the stiffeners (parallel to the main girders) are omitted in the figure.
Ship particulars are included in Table 13.
The base case design has been found to comply with the ILLC 66, where load and capacity
can be summarised as:
• Design load given by 1.75 ton/m2 . (This represents cargo loading rather than water
pressure)
• Stress criterion defined as the ultimate stress divided by 4.25.
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Figure 2 Finite element plot of hatch cover, seen from the side underneath (excluding
lower flanges and stiffeners)
3
HAZARDS RELATED TO HATCH COVERS AND COAMINGS
Different Hazard Identification studies have been conducted, and Table 12 in Appendix 3 lists
the hazards related to hatch covers and coamings as collected by MCA (2000).
The causes of water ingress and major contributors to the risk are represented in Figure 3.
Substantial water ingress due to failure of hatch covers or coamings
OR
Failure in heavy
weather due to
wastage caused
by lack of
maintenance
Figure 3
Hatch covers
open due to
rolling in
heavy weather
Failure due to severe
sea loads caused by
inappropriate heading
and speed for sea
condition
Failure in
heavy weather
due to design
compromise
Causes of hatch cover failure giving substantial water ingress
Wastage of hatch covers and coamings is easily detected by the crew, during classificatio n
surveys, and in port state controls, and is assumed to be a minor contributor to the risk. In the
LMIS database, 20 serious casualties were identified involving hatch cover failure and water
ingress, see Appendix 2. In one of these cases, poor maintenance is indicated as a cause.
Hence a rough estimate of the fraction of serious casualties due to hatch covers caused by
“Failure in heavy weather due to wastage caused by lack of maintenance” is taken as 5%.
5 of the casualties found in
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the LMIS casualty database included
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8
statements like “hatch cover lost”. 3 of these were also checked in Lloyd’s Casualty Archive.
In one of these cases it was specified that No. 1 hatch cover was blown off by ferocious
winds, and in another that the hatch cover was lost at sea. The event “Hatch cover opens due
to rolling in heavy weather” is taken to include the cases where the hatch covers are lost at
sea, including other causes than merely rolling in heavy weather. It is then estimated that
approximately 25% of the hatch cover failures resulting in severe flooding is related to “Hatch
cover opens due to rolling in heavy weather” in the above figure.
The background for looking at risk control options related to hatch covers is the international
discussion that have been going on regarding design loads for hatch covers, see e.g.
MSC/72/4/1, and DETR (1998). “Inappropriate heading and speed for sea condition” may be
a contributing factor to hatch cover failure in heavy weather. However, to investigate the
hypothesis that the design loads for hatch covers are too low, the remaining 70% of the
serious casualties involving flooding and hatch cover failure were assumed related to “Failure
in heavy weather due to design compromise” in the above figure.
The assumed relative importance of the causes to hatch cover failure given in the above figure
hence is:
• “Failure in heavy weather due to wastage caused by lack of maintenance” accounts for 5%
of the serious casualties attributed to hatch cover failure
• “Hatch covers open due to rolling in heavy weather” is taken to include all cases where
hatch covers are lost to sea, also for other reasons than rolling, and is assumed to account
for 25% of the serious casualties
• “Failure due to severe sea loads caused by inappropriate heading and speed for sea
condition” is taken as negligible
• “Failure in heavy weather due to design compromise” is taken to include the cases of
hatch cover collapse assumed to result from excessive green sea loads on the hatch covers.
The remaining 70% of the cases are assumed attributed to this failure mode.
Several other failure mechanisms exist for hatch covers and coamings, but these are expected
to cause smaller openings, and less water ingress. This water ingress may be sufficient to
jeopardise the cargo but do not pose immediate threats to safety. A brief summary of some of
these failure modes is given in Appendix 3.
The causes of progressive flooding is believed to be one of the following:
• Failure of bulkhead separating a flooded and not flooded cargo hold
• Failure of hull girder
• Failure of hatch cover of not flooded cargo hold
• Cargo liquefaction and loss of stability
• Side shell failure of not flooded cargo hold
The implementation of SOLAS XII will prevent progressive flooding due to failure of
bulkheads between any cargo holds on new ships and between No. 1 and 2 cargo holds on
existing ships.
It is in the risk analysis below assumed that progressive flooding in the future will take place
with the same probability as in the past. This is clearly conservative, and causes the risk
contribution from hatch cover failure to be over estimated.
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4
4.1
9
RISK CONTRIBUTION FROM HATCH COVER FAILURES
Casualty data
In the LMIS casualty database, April 1999 version, 195 serious casualties including total
losses involving water ingress have been found in the period from 1978 to 1998. The data
foundation represents 73,600 ship years for bulk carriers of 20,000+ DWT. In addition, there
are 27 cases where it is not clear whether the casualties involved water ingress or not. In 75 of
these 195 water ingress cases, no detailed information has been found regarding the cause of
the water ingress to the cargo holds, whereas some of these contain information about which
cargo holds suffered from water ingress.
27 cases were found where failure of hatch covers and water ingress or possible water ingress
was reported. 20 of the cases were recorded as serious casualties, of which 9 were total losses.
8 accidents involved in total 246 fatalities. 7 cases were recorded as non-serious casualties.
4.2
Overall risk contribution based on casualty data
The Potential Loss of Life (PLL) is taken as the average number of fatalities per ship year.
The contribution to PLL from hatch cover failures is estimated to:
PLLhatch covers =
n
246fatalities
=
= 3.34 ⋅ 10 −3 fatalities per ship year
m 73,600 ship years
In Annex 2, Table 9 gives a loss matrix for generic bulk carrier accidents. Here, a generic
serious casualty is estimated to cost US$ 5,608,000, and a total loss US$ 24,808,000. Given 9
total losses due to hatch cover failures and 11 serious casualties not leading to total losses, the
economic losses due to serious casualties and total losses due to failure of hatch covers is
estimated as:
ELhatch covers = f totalloss due to hatch covers ⋅ C totalloss + f seriouscasualtydue to hatch cover ⋅ Cseriouscasualty
=
9
11
⋅ 24,808,000 +
⋅ 5,608,000 = $3,900 per ship year
73,600
73,600
This constitutes a lower bound of the risk contribution from hatch cover failure, as only the
cases where this is explicitly stated are included in the data. At the same time, the bound is
also considered as a best estimate, since hatch cover failure is one of the failure mechanisms
easier to detect and hence more likely to be reported.
Assuming that all 75 cases found, where the source of water ingress was not accounted for,
are related to hatch cover failure, an upper bound of the frequency of water ingress due to
hatch cover failure is found. 37 of the casualties were total losses, and in total the casualties
involved 522 fatalities. Upper bounds for the risk contribution from hatch covers hence are
given as:
PLLu hatch cover =
246 + 522
= 1.04 ⋅10 − 2 fatalities per ship year
73,600
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10
The upper bound hence represents an increase of a factor of 3.1 in the annual fatality rate due
to hatch cover failure.
If cost estimates are combined with the frequencies for total loss and serious casualty, the
economic losses due to serious casualties and total losses due to failure of hatch covers is
estimated as:
ELu hatch cover =
9 + 37
11 + 38
⋅ 24,808,000 +
⋅ 5,608,000 = USD19,200 per ship year
73,600
73,600
The upper bounds may be used to evaluate the robustness of the recommendations for risk
control options considering hatch covers and coamings. However, hatch cover failure is
believed to be a failure mechanism relatively easy to detect, and the probability that this is
reported is hence believed to be larger than e.g. the probability that side shell failure leading
to water ingress being reported. The risk estimates based on the reported cases are thus taken
as best estimates.
Figure 4 and Figure 5 show the distribution of the number of total losses, serious casualties
and fatalities on age of ship at the time of the casualty. When looking at the number of total
losses and fatalities, there is no apparent and strong correlation with the ship age. When
looking at the number of serious casualties, the age category from 15 to 19 years appear to be
somewhat over-represented. These results are highly uncertain due to statistical uncertainty
caused by the low number of events. Due to the limited data and the vague trends, ageing
effects have not been considered in the present study of water ingress scenarios due to hatch
cover failure, i.e. hatch cover risks have been assumed as constant over the ship lifetime.
9
Number of casualties
8
Number of serious
casualties
Number of total losses
7
6
5
4
3
2
1
0
0-4
5-9
10-14
15-19
20-24
> 25
Age of ship
Figure 4
Number of serious casualties and total losses vs. the age of the ship at the
time of the casualty.
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Annex 4
11
80
Number of fatalities
70
60
50
40
30
20
10
0
0-4
5-9
10-14
15-19
20-24
> 25
Age of ship
Figure 5
casualty
4.3
Number of fatalities plotted against the age of the ship at the time of the
Casualty data broken down on location of water ingress and severity
In Appendix 2, casualty data has been evaluated in detail. The casualty data analysis, based on
a limited amount of data, gave strong indications that the hatch cover of No. 1 cargo hold is
the most troubled. The No. 1 hatch cover may be tied to 244 out of 246 fatalities, 9 total
losses, and approximately 8 serious casualties.
The risk contributions from the No. 1 hatch cover as deduced from historical data are:
PLLhatch cover1 =
ELhatch cover1 =
244
= 3.32 ⋅ 10 −3 fatalities per ship year
73,600
9
8
⋅ 24,808,000 +
⋅ 5,608,000 = US $3,600 per ship year
73,600
73,600
The annual fleet average probability of No. 1 hatch cover failure causing serious casualty or
total loss is estimated to:
0.77 ⋅ 20
p f , hatch cover 1 =
= 2.09 ⋅ 10 − 4
73,600
Upper bounds for the risk contribution are achieved by assuming that all 75 casualties not
accounted for are due to hatch cover failures, and that all of them are related to No. 1 cargo
hold:
244 + 522
PLLu hatch cover 1 =
= 1.04 ⋅ 10 − 2 fatalities per ship year
73,600
ELu hatch cover 1 =
9 + 37
8 + 38
⋅ 24,808,000 +
⋅ 5,608,000 = US $19,000 per ship year
73,600
73,600
The upper bound for the annual fleet average probability of No. 1 hatch cover failure causing
serious casualty or total loss is estimated to:
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Annex 4
p u f , hatch cover 1 =
4.4
12
0.77 ⋅ (20 + 75)
= 9.94 ⋅ 10 − 4
73,600
Risk contribution from No. 1 hatch covers on capesize bulk carriers
The base risk contribution from severe hatch cover failures for capesize bulk carriers was
estimated based on historical data and is summarised below.
In the below table, the serious casualties (including total losses) are distributed on bulk carrier
size.
Table 2
Comparison of different bulk carrier size categories
BC
category
Relative Number of serious Expected number of serious
part of casualties and total casualties if equal underlying
fleet
losses
frequency
handysize
0.48
11
9.6
handymax
0.24
4
4.8
panamax
0.19
1
3.8
capesize
0.09
4
1.8
Ratio of observations per
expected outcome
1
0.8
0.3
2
Based on this very limited data foundation, it is assumed that the capesize bulk carriers are
more at risk than the other size categories, and that the frequency of serious casua lty
(including total losses) is a factor of 2 higher than the average serious casualty frequency.
The estimated average probability of No. 1 hatch cover failure causing serious casualty or
total loss for capesize carriers, p capesize f , hatch cover1 , estimated according to:
p capesize f , hatch cover1 = p f , hatch cover 1 ⋅ k capesize
where
p f , hatch cover 1
was above found to be 2.09 ⋅ 10 −4 as a lower bound and 9.94 ⋅ 10 −4 as an upper
k capesize
bound
is a correction factor for capesize carriers found to be 2, see Table 2.
Splitting the risk on causes as assumed in Section 3, gives the results shown in Table 3.
Table 3
Hatch cover failure probabilities for capesize bulk carriers split on
causes.
Cause
Failure in heavy weather due to wastage
caused by lack of maintenance
Hatch covers open due to rolling in heavy
weather
Failure due to severe sea loads caused by
inappropriate heading and speed for sea
condition
Failure in heavy weather due to design
compromise
Relative
contribution
(%)
Annual
probability,
lower bound
Annual
probability,
upper bound
5
2.1 ⋅ 10 −5
1.0 ⋅ 10 −4
25
1.1 ⋅10 −4
5.0 ⋅ 10 −4
-
-
-
70
2.9 ⋅ 10 −4
1.4 ⋅10 − 3
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Annex 4
Total
5
13
100
4.2 ⋅ 10 −4
2.0 ⋅ 10 −3
COST EFFECTIVENESS ANALYSIS OF RISK CONTROL OPTIONS
5.1
General remarks
5.1.1
Risk model
A simple risk model is used to predict the effect of hatch cover risk control options, as shown
in the following event tree.
YES
NO
Casualty
severity
Number of
fatalities
per event
Total loss
of ship
n TL
Serious
casualty
nSC
Escalation?
pf
Serious water ingress
due to hatch cover
failure
PEsc
1 - PEsc
Figure 6
Risk model used to estimate risk reduction
The risk contribution, in terms of PLL, from hatch cover failure leading to serious water
ingress hence is given by:
PLL = p f ⋅ ( PEsc ⋅ nTL + (1 − PEsc ) ⋅ nSC ) fatalities per ship year
where pf is the probability of hatch cover failure, PEsc is the probability of the event escalation
to a total loss, nTL is the average number of fatalities per total loss, and nSC is the average
number of fatalities per serious casualty.
The annual probability of hatch cover collapse, pf, has been calculated by Structural
Reliability Analysis (SRA) for different hatch cover designs. The analyses are documented in
Appendix 7. Structural analyses, see Appendix 6, to establish a formulation of hatch cover
strength, and hydrodynamic analyses, see Appendix 4, to establish a loading formulation,
provided input to the SRA.
All risk control options evaluated in the following are aimed at reducing the probability of
failure of hatch covers. The risk reduction, ∆R, implied by a risk control option hence is taken
as:
∆R = ∆PLL ⋅ TExp = ∆p f ⋅ (PEsc ⋅ (1 − rSOLASXII ) ⋅ nTL + (1 − PEsc ⋅ (1 − rSOLASXII )) ⋅ n SC ) ⋅ TExp per ship
Here, TExp is the expected time interval when the risk control option is effective, and ∆pf is the
reduction in probability of hatch cover failure due to the risk control option. rSOLASXII is the
expected risk reduction due to the implementation of SOLAS Chapter XII and was in Annex
Revision 0
Annex 4
14
2 estimated to 22% for existing bulk carriers and 65% for new-buildings.
The casualty data evaluation of Appendix 2 gives:
PEsc = 0.5 given No. 1 cargo hold flooding
nTL =
244 fatalities
= 31.7 fatalities per total loss given No. 1 cargo hold flooding
0.5 ⋅ (20 ⋅ 0.77 ) total losses
n SC = 0 fatalities per serious casualty (excluding total loss), given No. 1 cargo hold flooding
The economic benefits, ∆B, from a risk control option is estimated according to:
∆B = ∑
∆Bi
(1 + r )n − 1
=
∆
B
⋅
i
(1 + r )i
r (1 + r ) n
= ∆p f
(
1+ r) n − 1
⋅ ( PEsc ⋅ (1 − rSOLASXII ) ⋅ CTL + (1 − PEsc ⋅ (1 − rSOLASXII )) ⋅ C SC ) ⋅
r (1 + r )n
where ∆Bi, is the reduced economic losses due to serious casualties and total losses per ship
year, r is the discount interest rate, representing a corporate rate of return taken as 10% in the
calculations below. CTL and CSC are the estimated costs of a total loss and serious casualty
respectively, as given in Table 9 of Annex 2.
5.1.2
Cost effectiveness calculations
In the following evaluations, the Gross and Net Cost of Averting a Fatality (CAF) have been
used. They are defined as:
∆C
∆R
∆C − ∆B
NetCAF =
∆R
GrossCAF =
where ∆C is the total costs, ∆B is the economic benefits, and ∆R is the number of fatalities
averted by the risk control option.
Since the Gross CAF only compares the risk control option costs and the implied risk
reduction in terms of fatalities averted, it clearly identifies the risk control options, which can
be justified from a safety perspective alone. If an risk control option implies large economic
benefits but has no safety implications, it would display a high Gross CAF.
As the Net CAF also accounts for economic benefits, it may also be used to identify which
risk control options are justifiable from a commercial point of view or combined commercial
and safety point of view.
5.1.3
Assumptions
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Annex 4
15
In the following, the different assumptions made during the risk control option evaluations are
described.
The base risk as obtained from casualty data is assumed representative for bulk carrier hatch
covers designed according to the ILLC 66. This is based on the assumption that bulk carriers
with hatch covers designed according to the ILLC 66 constitute the majority of the total
population in the period from 1978 to 1998. The part of the fleet sailing with stronger hatch
covers during the period is assessed to less than 10%.
The Enhanced Survey Programme (ESP) is assumed not to affect the base risk contribution
from hatch cover collapse (“Failure in heavy weather due to design compromise”) or from
hatch covers being washed away (“Hatch covers open due to rolling in heavy weather”).
The objective of part of the study was to investigate risk control options related to hatch cover
strength. In order to evaluate these risk control options, the starting hypothesis was that a 70%
of serious casualties involving hatch covers were related to hatch cover collapse due to
buckling. This is a conservative assumption, since some of the casualties may be related to
other failure modes giving slower flooding. This gives too high risks related to the collapse
(due to buckling) failure mode.
For new buildings, the difference in the costs of hatch covers is assumed to be a function of
steel weight alone. The number of working hours and other production (e.g. welding bar,
oxygen, scaffolding, etc) costs are assumed to remain the same independent of whether the
design is a UR S21 or a ILLC 66 design. The costs of maintenance are also assumed to be
equal. Hence the added costs of UR S21 is related to the extra steel weight alone. The cost
model is given as:
∆C = ∆T ⋅ c1
where ∆C is the added cost due to strengthened hatch covers, ∆T is the added amount of steel,
and c1 is the cost per metric tonne of steel, including welding.
For existing ships it is assumed that the hatch covers simply are replaced with strengthened
hatch covers. The operation is assumed to take place during survey and hence does not imply
extra off hire. The total costs thus are assumed equal to the costs of the new hatch covers,
taken as the total steel weight of the hatch covers, T, multiplied by an average cost per tonne
steel, c2 , including work, welding bar, oxygen, scaffolding etc:
∆C = T ⋅ c2
The life expectancy of an average bulk carrier is in the following taken as 25 years.
5.2
New-building requirements for hatch cover design pressure
In order to estimate the marginal cost effectiveness of different design pressure levels for
hatch cover, hatch cover costs, as achieved from the bulk carrier department of MacGregor
(see Appendix 1), and annual probabilities of hatch cover collapse were used. Details are
given in Appendix 8. The structural reliability results and cost data have been used to
establish linear regression models for annual probability of hatch cover failure and hatch
cover costs, as functions of the design pressure. Based on these, the marginal cost
Revision 0
Annex 4
16
effectiveness has been estimated as a function of the design pressure.
1.00E+08
IACS UR S21
Gross CAF (USD)
1.00E+07
1.00E+06
1.00E+05
1.00E+04
1.00E+03
Gross CAF
1.00E+02
Criterion
1.00E+01
1.00E+00
0
20
40
60
80
100
Hatch cover strength (kN/m2)
Figure 7
Marginal Gross CAF as a function of hatch cover strength, prior to the
implementation of SOLAS Chapter XII.
The results are applicable for changes in new-building requirements at the time of the
implementation of IACS UR S21, which shows that IACS UR S21 from a cost effectiveness
perspective seems to be close to optimal.
To evaluate an increase in the design loads today, the effect of SOLAS Chapter XII has to be
incorporated, which has been done in Figure 8.
1.00E+09
Gross CAF (US$)
1.00E+08
UR S21
1.00E+07
Gross CAF
1.00E+06
Criterion
UR S21+30%
1.00E+05
1.00E+04
1.00E+03
1.00E+02
1.00E+01
1.00E+00
0
20
40
60
80
100
Figure 8
Marginal Gross CAF as a function of design pressure, given the
implementation of SOLAS Chapter XII.
Presently, an increase in the design loads and hence the strength of hatch covers of e.g. 30% is
estimated to imply a Gross CAF of US$86 million, by far exceeding the recommended
decision criterion.
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Annex 4
5.3
17
RCO No. 1; Replacement of ILLC 66 hatch covers
Risk control option No. 1 implies replacing No. 1 and 2 hatch covers designed according to
ILLC 66 with hatch covers designed according to IACS UR S21.
Table 4 refers the estimated annual probabilities of hatch cover collapse as estimated by SRA,
see Appendix 7.
Table 4
Annual probabilities of hatch cover collapse. (For SRA
estimates, see Appendix 7)
Case
Initial design, ILLC 66
UR S21
Probability of hatch cover
collapse, estimated by SRA
9.35⋅10-4
1.16⋅10-5
Probability of hatch cover
collapse, estimated from
casualty data
2.9⋅10-4 - 1.4⋅10-3
-
Note that the estimated annual probability of hatch cover collapse for the ILLC 66 design
(9.35⋅10-4 ) is high compared to the estimates based on the casualty data (2.9⋅10-4 - 1.4⋅10-3 ), of
which the lower bound is assessed to be the closest to the true hatch cover collapse
probability. The relatively high SRA estimate may e.g. be due to:
• uncertainty in the hatch cover loading model
• the assumption about a uniform pressure level on the hatch cover
• the weather routing in the data period being more effective than the weather routing
during the time period when wave data was collected
• the world-wide wave data not being representative for the typical capesize bulk carrier
trades.
In the calculations below, the lower probability of ha tch cover collapse as estimated from
casualty data have been used to generate the lower limit for the risk reduction. The estimated
annual probability of hatch cover collapse from SRA has been used to generate an upper limit
for the risk reduction.
To establish the lower estimate for the risk reduction and economic benefits, the difference
between the lower estimate for the annual failure probability for the ILLC 66 design and the
SRA estimate for the IACS UR S21 design is utilised:
∆p l f = 2.9 ⋅ 10 −4 − 1.15 ⋅10 − 5 = 2.8 ⋅ 10 −4
The lower bound for the number of fatalities averted per ship year by the measure compared
to IACS UR S21 is estimated according to:
∆R l = ∆p l f ⋅ (PEsc (1 − rSOLASXII ) ⋅ nTL + (1 − PEsc (1 − rSOLASXII )) ⋅ n SC ) ⋅ TExp
= 2.8 ⋅ 10 − 4 ⋅ (0.5 ⋅ (1 − 0.26 ) ⋅ 31.7 + 0 ) ⋅ TExp = 3.4 ⋅ 10 − 3 ⋅ TExp
The lower bound for the economic benefits are estimated to:
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Annex 4
∆B = ∆p
l
l
f
18
n
(
1+ r) −1
⋅ (PEsc ⋅ (1 − rSOLASXII ) ⋅ CTL + (1 − PEsc (1 − rSOLASXII )) ⋅ C SC ) ⋅
n
r (1 + r )
= 2.8 ⋅ 10 - 4 ⋅ (0.5 ⋅ (1 − 0.22) ⋅ US$24,808,000 + (1 − 0.5(1 − 0.22 )) ⋅ US$5,608,000 ) ⋅
1.1 Exp − 1
T
0.1 ⋅ 1.1 Exp
T
1.1 Exp − 1
0.1 ⋅1.1TExp
T
= US $3,600 ⋅
Similarly, the upper estimates for the risk reduction and economic benefits are estimated from
the difference in the annual failure probabilities as estimated from the SRA:
∆p u f = 9.35 ⋅ 10 −4 − 1.15 ⋅10 − 5 = 9.2 ⋅ 10 −4
The upper bound for the number of fatalities averted per ship year by the measure compared
to IACS UR S21 is estimated according to:
∆R u = ∆p u f ⋅ (PEsc (1 − rSOLASXII ) ⋅ nTL + (1 − PEsc (1 − rSOLASXII )) ⋅ nSC ) ⋅ TExp
= 9.2 ⋅ 10 − 4 ⋅ (0.5 ⋅ (1 − 0.26) ⋅ 31.7 + 0 ) ⋅ TExp = 1.1 ⋅10 −2 ⋅ TExp
The upper bound for the economic benefits are estimated to:
∆B u = ∆p u f ⋅ (PEsc ⋅ (1 − rSOLASXII ) ⋅ CTL + (1 − PEsc (1 − rSOLASXII )) ⋅ C SC ) ⋅
(1 + r )n − 1
n
r (1 + r )
= 9.2 ⋅ 10 ⋅ (0.5 ⋅ (1 − 0.22 ) ⋅ US $24,808,000 + (1 − 0.5(1 − 0.22)) ⋅ US $5,608,000 ) ⋅
-4
1.1 Exp − 1
T
0.1 ⋅1.1
TExp
1.1 Exp − 1
= US $12,100 ⋅
0.1 ⋅1.1TExp
T
A cost per metric tonne steel in the range of US$1,500 to US$2,500 as suggested by
MacGregor, see Appendix 1, is used to estimate the risk control option cost. MacGregor also
indicated that hatch covers No. 1 and 2 designed according to UR S21 would weigh
approximately 35 tonnes each. A lower and upper cost estimate hence is given by:
∆Clower = T ⋅ c 2 = 2 ⋅ 35tonnes ⋅ US $1,500 / tonne = US $105,000
∆Cupper = T ⋅ c2 = 2 ⋅ 35tonnes ⋅ US $2,500 / tonne = US $175,000
Even the highest cost estimate may be in the lower range, as costs related to design, spare
parts, operating equipment, transport and strengthening of the coaming etc, are excluded.
The risk control option Gross and Net CAFs, which depend on the age of the ship at the time
of implementation, are shown in the Table below.
Table 5
Cost effectiveness results for exchanging LLC hatch covers with IACS UR
S21 hatch covers
Revision 0
Annex 4
Age of ship at ∆ R
time of
(per ship)
implementation
∆C
(US$)
19
∆B
(US$)
10 years 5.16E-02 - 1.71E-01 105,000 - 175,000 27,700 - 92,000
15 years 3.44E-02 - 1.14E-01 105,000 - 175,000 22,400 - 74,300
20 years 1.72E-02 - 5.71E-02 105,000 - 175,000 13,821 - 45,800
Gross CAF
(US$ million)
0.6 - 3.4
0.9 - 5.1
1.8 - 10.2
Net CAF
(US$ million)
0.1 - 3.2
0.3 - 4.9
1.0 - 9.9
The above results indicate that the uncertainties in the results are large, and that no firm
conclusions can be drawn for whether ILLC 66 hatch covers on existing bulk carriers should
be replaced or not. However, the higher values are assessed to be the more realistic.
Colman (2000) describes the cost of reinforcing No. 1 and 2 hatch covers on existing capesize
bulk carriers as unlikely to exceed £150,000, or US$225,000. This is in excess of the costs
used in the present evaluation, and would make all the Gross and Net CAF estimates become
higher.
5.4
RCO No. 2; Replacement of IACS UR S21 hatch covers
Risk control option No. 2 implies replacing No. 1 and 2 hatch covers designed according to
IACS UR S21 with hatch covers designed to withstand a 30% increase in the design loads of
IACS UR S21.
Table 6 gives the estimated annual probabilities of hatch cover collapse for the relevant
designs, as estimated by SRA, see Table 22 of Appendix 7.
Table 6
Annual probabilities of hatch cover collapse, as estimated by SRA
Case
UR S21
30% increase in UR S21 design loads
Probability of hatch cover collapse, pf
1.16⋅⋅10 -5
7.85⋅⋅10 -7
The relative difference in annual failure probability between the IACS UR S21 design and the
design involving 30% increase in UR S21 design loads is estimated to:
∆p f = 1.15 ⋅ 10 −5 − 7.85 ⋅10 −6 = 1.07 ⋅ 10 −5
The number of fatalities averted per ship year by the measure compared to IACS UR S21 is
estimated according to:
∆R = ∆p f ⋅ ( PEsc (1 − rSOLASXII ) ⋅ nTL + (1 − PEsc (1 − rSOLASXII )) ⋅ n SC ) ⋅ TExp
= 1.07 ⋅ 10 −5 ⋅ (0.5 ⋅ (1 − 0.26 ) ⋅ 31.7 + 0 ) ⋅ TExp = 1.3 ⋅10 −4 ⋅ TExp
The economic benefits are estimated to:
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Annex 4
∆B = ∆p f
20
n
(
1 + r ) −1
⋅ ( PEsc ⋅ (1 − rSOLASXII ) ⋅ CTL + (1 − PEsc (1 − rSOLASXII )) ⋅ CSC ) ⋅
n
r (1 + r )
= 1.97 ⋅ 10 -5 ⋅ (0.5 ⋅ (1 − 0.22 ) ⋅ US $24,808,000 + (1 − 0.5(1 − 0.22)) ⋅ US $5,608,000) ⋅
1.1 Exp − 1
T
0.1 ⋅ 1.1
TExp
1.1 Exp − 1
0.1 ⋅1.1TExp
T
= US $140 ⋅
Upper and lower bounds for the risk control option cost are:
∆Clower = T ⋅ c 2 = (2 ⋅ 40tonnes )US$1,500 / tonne = US $120,000
∆Cupper = T ⋅ c 2 = (2 ⋅ 40tonnes )US $2,500 / tonne = US $200,000
The resulting Gross and Net CAFs are shown in the table below.
Table 7
Cost effectiveness results for replacing IACS UR S21 hatch covers with
hatch covers designed to 30% increase in IACS UR S21 loads
Age of ship at time ∆ R
of i mplementation (per ship)
10 years
15 years
20 years
2.00E-03
1.34E-03
6.68E-04
∆C
(US$)
∆B
(US$)
120,000 - 200,000
120,000 - 200,000
120,000 - 200,000
Gross CAF
(US$ million)
1,100
870
540
60 - 100
90 - 150
180 – 300
Net CAF
(US$ million)
60 – 100
90 – 150
180 - 300
In the above table, it is indicated it is not cost effective to exchange No. 1 and 2 hatch covers
on capesize carriers designed according to IACS UR S21 with new hatch covers designed
according to a 30% increase in the IACS UR S21 design loads. The cost estimates used are
also lower bounds of the true costs, giving a robust recommendation that IACS UR S21 hatch
covers should not be exchanged on existing ships.
Colman (2000) describes the cost of reinforcing No. 1 and 2 hatch covers on existing capesize
bulk carriers as unlikely to exceed £150,000, or US$225,000. This is in excess of the costs
used in the present evaluation, and would make the Gross and Net CAF estimates higher.
5.5
RCO No. 3 Hydraulic hatch cover closure
Risk control option No. 3 implies fitting No. 1 cargo hold with a hydraulic hatch cover
closure system.
In Table 3, the annual probability of the failure mode “hatch cover No. 1 being washed
overboard, or opening due to extensive rolling” was estimated to:
p f = 1.05 ⋅ 10 −4
Assuming
that
hydraulic
Revision 0
hatch cover closure system on No. 1
Annex 4
21
hatch cover only is able to prevent all such failures from happening, the risk reduction over
the remaining lifetime of an existing bulk carrier is estimated to:
= 1.05 ⋅10 − 4 ⋅ (0.5 ⋅ (1 − 0.22 ) ⋅ 31.7 + 0 ) ⋅ TExp = 1.3 ⋅10 −3 ⋅ TExp
∆B = ∆p f
n
(
1+ r) −1
⋅ ( PEsc ⋅ (1 − rSOLASXII ) ⋅ CTL + (1 − PEsc (1 − rSOLASXII )) ⋅ C SC ) ⋅
n
r (1 + r )
= 1.05 ⋅10 - 4 ⋅ (0.5 ⋅ (1 − 0.22 ) ⋅ US $24,808,000 + (1 − 0.5(1 − 0.22)) ⋅ US $5,608,000 ) ⋅
1.1 Exp − 1
T
0.1 ⋅1.1
TExp
1.1 Exp − 1
= US $1,380 ⋅
0.1 ⋅ 1.1TExp
T
For bulk carrier new-buildings, the risk reduction from SOLAS Chapter XII was estimated to
65%, giving:
= 2.61 ⋅ 10 − 4 ⋅ (0.5 ⋅ (1 − 0.65) ⋅ 31.7 + 0) ⋅ 25 = 1.46 ⋅10 −2 fatalities averted
∆B = ∆p f
n
(
1 + r) − 1
⋅ ( PEsc ⋅ (1 − rSOLASXII ) ⋅ CTL + (1 − PEsc (1 − rSOLASXII )) ⋅ CSC ) ⋅
n
r (1 + r )
= 2.84 ⋅10 ⋅ (0.5 ⋅ (1 − 0.65) ⋅ US$24,808,000 + (1 − 0.5(1 − 0.65)) ⋅ US $5,608,000 ) ⋅
-4
1.1 Exp − 1
T
0.1 ⋅1.1
TExp
1.1 Exp − 1
= US $940 ⋅
0.1 ⋅ 1.1T Exp
T
A gross estimation of a hydraulic securing system for hatch covers envisages eight hydraulic
pistons per hatch. Assuming the modification is made on a single hatch, the additional
mechanical equipment, hydraulic piping, dedicated hydraulic unit and signalling system
amounts to about US$ 24,000, not including the additional inspections during the periodical
hatch cover surveys. In general, the maintenance costs of such systems may be significant,
and these systems are also regarded as less flexible than quick-acting cleats, which are
generally chosen. Assuming that the maintenance costs add up to 10% of the new-building
cost per year, the cost of this risk control option is estimated to:
25
US $2,400
= US $58,000
1.05i
i =1
∆C = US $24,000 + ∑
In Table 8, the Gross and Net CAF for implementing hydraulic hatch cover closure on new
and existing bulk carriers have been estimated.
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Annex 4
22
Table 8
Cost effectiveness results for introducing hydraulic hatch cover closure
systems on No. 1 hatch cover
Age of ship at time of ∆ pf
implementation
10 year old ships
15 year old ships
20 year old ships
New-buildings
1.05E-04
1.05E-04
1.05E-04
1.05E-04
∆R
∆C
(per ship) (US$)
1.95E-02
1.30E-02
6.49E-03
1.46E-02
∆B
(US$)
58,000
58,000
58,000
58,000
10,500
8,450
5,210
8,550
Gross CAF
(US$ million)
2.97
4.46
8.91
3.97
Net CAF
(US$ million)
2.43
3.81
8.11
3.39
If hydraulic hatch cover closure systems are implemented on several hatch covers, costs of
order of magnitude US$ 100,000 are anticipated, excluding maintenance costs. The risk
reduction would not increase accordingly since the majority of the fatalities are related to
failure of No. 1 hatch cover. Hence, fitting all cargo holds with hydraulic hatch cover closure
systems is expected to have Gross and Net CAFs by far exceeding the recommended decision
criterion.
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Annex 4
6
23
DISCUSSION
MSC 72/16 recommended a decision criterion of US$ 3 million for risk control options
involving reduction in both fatalities and injuries and US$ 1.5 million for risk control options
involving a reduction in fatalities only. The latter is likely to be the case for casualties related
to hatch covers.
The marginal Gross and Net CAF evaluation of hatch cover design load levels indicated that
the IACS UR S21 design load levels are close to optimal compared to the recommended
decision criterion of MSC 72/16.
When evaluating the risk control options involving retrofitting existing bulk carriers with
reinforced hatch covers, low cost estimates were used, not accounting for costs related to
design, strengthening of support structure, and replacing wheels, rails, wheel lifters, fittings,
and driving equipment. Colman (2000) described the cost of reinforcing No. 1 and 2 hatch
covers on existing capesize bulk carriers as unlikely to exceed £150,000, or US$225,000, and
as unlikely to exceed £100,000, or US$ 150,000, for new-buildings. This is in excess of the
costs used in the present evaluation, and would make the Gross and Net CAF estimates
higher. The estimated risk reductions are optimistic due to the assumptions that 70% of the
serious casualties and total losses were caused by hatch cover collapse. In reality, some of
them may have been caused by smaller leaks over a longer period of time, for which other
risk control options than presently studied would have been relevant.
Due to the costs being under-estimated and the risk reductions over-estimated, the risk control
options of replacing hatch covers on existing ships tend to look more cost-effective than they
actually are. Even under these assumptions, the evaluations do not give robust conclusions for
whether the ILLC 66 hatch covers should be replaced on the average bulk carriers. However,
it is possible that the hatch cover capacity varies to such an extent that the replacement of
hatch covers on a subset of the fleet could be justified. If a screening of the relevant bulk
carriers could identify individual ships with particularly high hatch cover risks, then the
replacement of their No. 1 and 2 hatch covers could be justified from a cost-effectiveness
perspective.
The uncertainties in the results are considerable, for several reasons. The base case risk level
was based on the assumption that the historical risk level was applicable for ILLC 66 hatch
cover designs, whereas parts of the bulk carrier fleet have had hatch covers designed to
stricter standards by individual class societies 1 . However, throughout the period from 1978 to
1998, less than 10% of the fleet is believed to have had stronger hatch covers. The statistical
data is open to interpretation, and risk intervals were based on casualties reported as related to
hatch covers and the SRA results. In addition, casualties have occurred for which little
information is available, e.g. 6 total losses recorded as “missing”. If assigning all casualties
with insufficient information available to the flooding scenarios due to hatch cover failure, the
risks were estimated as a factor 3 above what is presently used. Under such assumptions, it
would be cost-effective to replace ILLC 66 hatch covers.
Risk control option No. 3 implying an hydraulic hatch cover closure system on No. 1 hatch
cover was also found to display Gross and Net CAFs exceeding the recommended decision
criterion.
1
DNV has had stricter requirements for hatch covers since the early 1970s.
Revision 0
Annex 4
24
In the present evaluation, a representative capesize bulk carrier was selected, primarily
because based on experimental test results it was suspected that the design loads for these
bulk carriers were in the lower range.
The casualty data indicates that the capesize carriers are more at risk for hatch cover failure
than the smaller ships. A reason for this may be their lengths being comparable to the wave
lengths in severe sea states, giving increased sea loads on the hatch covers and hence
increased probability of water ingress. The handys ize carriers appear to be more exposed than
the handymax and panamax, probably due to lack of damage survivability. In many cases, the
flooding of one cargo hold is sufficient for these ships to suffer from lack of buoyancy. The
panamax bulk carriers may have an order of magnitude lower frequency of total loss due to
hatch cover failure compared to the average risk for bulk carriers. Given hatch cover
strengthening of panamax, handysize, and handymax carriers, the relative risk reduction is
expected to be similar to the risk reduction estimated for the capesize carriers. The absolute
risk reduction is expected to be 50% or less for the smaller bulk carriers compared to the
capesize carrier. The costs of strengthening hatch covers are also expected to be lower. Hence,
the cost effectiveness for smaller bulk carriers is expected to be of the same order of
magnitude as the cost effectiveness estimated for the capesize bulk carrier, giving similar
recommendations.
7
RECOMMENDATIONS
No risk control options related to flooding scenarios caused by hatch cover failure were
identified, which were associated to Gross and Net CAFs clearly below the decision criterion
as recommended in MSC72/16. Regarding the risk control option involving the replacement
of ILLC 66 hatch covers, the Gross and Net CAF intervals are wide, reflecting considerable
uncertainties. Hence no firm conclusion can be given for the implementation of this risk
control option. However, it is possible that the variation in hatch cover capacity of ILLC 66
design is so large that the replacement of hatch covers on a subset of the fleet could be
justified. If a screening of the relevant bulk carriers could identify individual ships with
particularly high hatch cover risks, then the replacement of their No. 1 and 2 hatch covers
could be justified from a cost-effectiveness perspective. Other alternative risk control options
may also exist, like only replacing No. 1 hatch cover, and not both No. 1 and 2. Hence, it is
recommended that such a risk control options be evaluated further.
This study has assessed IACS UR S21 hatch cover designs to be significantly more reliable
than ILLC 66 hatch cover designs with respect to hatch cover collapse caused by green seas
on deck. The cost effectiveness assessment indicated that the design loads of IACS UR S21
are close to optimal. At present, IACS model tests are being carried out, increasing the
knowledge about green loads on deck. When obtained, new information about hatch cover
loading is recommended be included in UR S21, but as the requirements seem to have close
to an optimal level of reliability, this is recommended be maintained. This would mean that
the increased knowledge, when implemented in UR S21, would give less uncertainty and
probably less variability between UR S21 hatch cover designs.
8
REFERENCES
Bitner-Gregersen,
E.
M.,
Revision 0
“Environmental
Description
for
Annex 4
25
Long-Term Load Response of Ship Structures”, DVN Research Report No.: 94-2054, Rev.
No. 1, 1995.
Colman, The Honourable Mr. Justice Colman (2000), “Report of the re-opened formal
investigation into the los of the MV Derbyshire”. In the High Court of Justice (Admirality
Court). November, 2000.
DETR (1998), “M.V. Derbyshire Surveys, UK/EC Assessors’ Report, A summary”,
Department of the Environment, Transport and the Regions, UK, March 1998.
DNV (1992), “PROBAN – General Purpose Probabilistic Analysis Program, Theory
Manual,” Det Norske Veritas, report No. 92-7049. rev. No.01, Høvik.
DNV (1994), “Wadam, Wave Loading by Diffraction and Morrison Theory”, User’s Manual,
DNV Report No. 94-7100, December 1994.
DNV (1995a), POSTRESP, Interactive Postprocessor for General Response Analysis, User’s
Manual, DNV Report No. 95-7014, August 1995.
DNV (1995b), DNV Classification Note 30.1, “Buckling Strength Analysis”, July 1995.
Faulkner, D., Corlett, B. J., Romeling, J. U. (1996): “Design of hatch covers and coamings for
abnormal waves”, the international conference on " Watertight Integrity and Ship
Survivability" organised by the Royal Institute of Naval Architects (RINA) in London 21-22
November 1996.
DNV (1997), “SWAN-2 Theory and Numerical methods”, DNV Research report No. 942030, revision 1, 18th of April 1997.
DNV (1998), DNV Classification Note 30.7, “Fatigue Assessment of ship Structures”,
September 1998.
IACS (1997) “Evaluation of Scantlings of Hatch Covers of Bulk Carrier Cargo Holds”, IACS
Requirements 1997, Volume 1, S21.
International Conference on Load Lines (1966), “Final Act of the conference with
attachments including the International Convention on Load Lines, 1966”, IMO London,
1981.
MCA (2000), Hazard identification as circulated by MCA on 25 April 2000 (IACS message
0032_IAe).
MSC 70/4, “Sensitivity of wetness and deck loads to bow height and forward buoyancy
reserves in extreme weather conditions”, submitted by the United Kingdom.
MSC 70/4/6, “Operational measures for avoiding dangerous situations in extreme weather
conditions”, submitted by Greece.
MSC 72/4/1 “Green sea loads on hatch covers and deck wetness derived from seakeeping
model tests on a range of bulk carriers”, submitted by the United Kingdom.
MSC 72/4/1/Add. 1: “Further green sea loads results of seakeeping model tests on a range of
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26
bulk carriers”, submitted by the United Kingdom.
MSC 72/16, “Decision Parameters including Risk Acceptance Criteria”, submitted by
Norway.
Nestegård, A., Krokstad, J.R. (2000), “Non- linear wave loading”, DNV Report No. 20003401, Høvik, 2000.
Skjong, R., Bitner-Gregersen, E., Cramer, E., Croker, A., Hagen, Ø., Korneliussen, G.,
Lacasse, S., Lotsberg, I., Nadim, F., Ronold, K.O. (1996), “Guideline for Offshore Structural
Reliability Analysis – General”. Det Norske Veritas Research Report No. 95-2018. DNV,
Høvik, Norway.
Østvold, T. K., Steen, E. (2001), “Non- linear finite element analyses of hatch covers”, DNV
Report No. 2001-0391, DNV, Høvik, Norway.
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Appendix 1 Information from MacGregor regarding hatch cover costs,
email of 8 October 2000.
Dear Ms. Eknes,
Thank you for involving us and giving us the opportunity to comment on your study. Both the
LLC and UR S21 are everyday issues for us and those are easy to comment upon, but the new
40 % higher load seems quite a big challenge. We have no information of hatch cover
failures for ships where the UR S21 is already considered. There might be reasons to increase
the load, but this issue is not for us to comment upon.
You do not explain which stress levels are considered for the 40 % higher loads so any
estimates from our side is pure guess-work.
You have only mentioned the load influences on Hatch 1. For most cape size bulk carriers
also Hatch 2, being within the forward 25 % of the ship, is also subject to increased loads.
However not of the same magnitude, but the corrosion additions applies here in the same way
as for Hatch 1.
The UR S21 loads are typically:
for Hatch 1 6-6,5 t/m2
for Hatch 2 4-4,5 t/m2
Typical steel weights for an open web structure (square metre weights calculated on the area
of top plate) are:
Hatch cover 1, according to LLC only
185 kg/m2
Hatch cover 1, according to UR S21
270 kg/m2
180 kg/m2
220 kg/m2
We have assumed that H.1 would be 15 x 17 m and H.2 would be 15 x 20 m. The
corresponding panel weights would be:
24 tonnes (net, without fittings)
The average corrosion addition (included in the above weights) is some 17 % for an open web
structure, slightly higher for a double skin structure. The fittings weight (within the normal
load range) is some 10 % of the steel structure weight.
A 40 % higher load would mean approximately a 10-15 % addition to above (UR S21)
weights.
As you can see from the above, the weight increase can be considerable and this would mean,
that also the fittings and drive equipment would be bigger. Thus the size of wheels, rails,
wheel lifters, hydraulic motors, support pads, guides and wedges would need to be increased.
If we wish to keep the operation times the same, hydraulic pump unit and the hull piping
should also be bigger.
Today we try to harmonise the fitting range by using a higher amount of high tensile steel for
the forward hatch covers in order to save weight. Thus we are able in most cases to use for
instance hydraulic motors of one size throughout the whole vessel. We doubt this would be
possible for an increased load giving the shipowner an additional range of fittings and thus an
increased spare parts handling cost.
For side rolling hatch covers
to a newbuilding, the fabrication cost
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28
is in the magnitude of 1.200-2.000 US$/tonne, depending on where the hatch covers are
produced (Asia-Europe).
We have not even tried to estimate these costs for the bigger fittings, but the cost could be
considerable.
The cost for existing ships follows the above, but you need to consider, that in most cases the
existing fittings cannot be reused. It is also doubtful, that the coamings are sufficiently strong.
Have you considered the higher transversal forces due to increased vertical loads? There
might also be a need to increase the size of transversal stoppers to such a magnitude, that
today’s standard solutions cannot be used, requiring even a completely new hatch cover
system.
The cost for the steel structure when replacing one hatch for an existing ship would be
approximately 1.500-2.500 US$/tonne, excluding design, spare parts, operating equipment,
transport and strengthening of the coaming, of course depending on where the hatch covers
are produced (Asia-Europe), and depending on the delivery time.
We do not believe it feasible to strengthen existing hatch covers as most of the top plate
should be renewed due to the stresses and due to buckling.
We do hope the above clarifies the situation. Should you have any further queries, please do
not hesitate to contact us again. We would be very interested in following the development of
your studies. If a report would be published, please send one also to us.
Best regards
MacGREGOR (FIN) Oy
Hatch Covers
Torbjörn Dahl
Senior Naval Architect, Bulk Ships
Hallimestarinkatu 6
FIN-20780 KAARINA, Finland
Telephone: +358-2-4121 313
Mobile phone: +358-40-524 4601
Fax: +358-2-4121 508
E-Mail: torbjorn.dahl@macgregor-group.com
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Appendix 2
29
Break-down of Casualty Data
Below the hatch cover casualty data are further investigated, in order detect trends and to attribute
the casualties to different circumstances.
The tree below illustrates the breakdown of 20 serious casualties on water ingress location, total
losses, and fatal accidents.
When trying to quantify the different branches of the tree, in many cases there are only 1 or perhaps
0 observations. This is of course not sufficient to make any reliable estimates for the different
branch probabilities. The results from the tree is mainly used to estimate the risk contribution from
No. 1 and 2 hatch covers, and this is also the branches to which the majority of the cases relate.
Frequency
Failure of at least no.
1 hatch cover
P1
Loss of ship
P4
Serious
casualty
P5
Flooding
due to hatch
cover failure
Failure of at least no. 2
hatch cover, but not no.1
hatch cover
P2
Loss of ship
P6
P3
Fatalities
P12
No fatalities
P13
Fatalities
P14
No fatalities
P15
Fatalities
P16
P7
No fatalities
P17
Fatalities
P18
No fatalities
P19
Serious
casualty
P9
Figure 9
No fatalities
P11
Serious
casualty
Loss of ship
P8
Failure of any other
hatch cover, and not
n.1or n.2 hatch cover
Fatalities
P10
Fatalities
P20
No fatalities
P21
Breakdown of casualties on location of failed hatch cover, and consequences of
casualty.
Frequency of flooding due to hatch cover failure
In the LMIS casualty database, for bulk carriers of 20,000 DWT and larger, 20 cases were found
involving failure of hatch cover and water ingress. In 19 of the cases, water ingress was reported,
while one case probably involved water ingress although not specified. An estimate of the
frequency of serious casualty involving water ingress due to hatch cover failure hence is given as:
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30
g
20
=
= 2.7 ⋅ 10 −4 annual frequency of serious casualty or total loss involving water
m 73,600
ingress due to hatch cover failure.
f1 =
Location of water ingress
In 13 of the 20 serious casualties and total losses involving water ingress and hatch cover failure
events, the location for the water ingress is given. The results are shown in the below table.
Table 9
Number of casualties split on reported
location of water ingress
No. 1 cargo hold
n. 1 and 2 cargo holds
n. 2 cargo hold
n. 4 cargo hold
Total reported
unknown cases
Total
Number of occurrences
7
3
1
1
1
13
7
20
% of total
54
23
7.7
7.7
7.7
100
Based on the above table, the probability of the water ingress being related to No. 1 cargo hold is
taken as:
g 7 +3
P1 = =
= 0.77
m
13
The probability of the water ingress involving No.2 cargo hold and not No. 1 is taken as:
g 1+ 1
P2 = =
= 0.15
m 13
Finally, the probability of the water ingress involving any other cargo hold, but not No. 1 or No.2 is
taken as:
g
1
P3 = =
= 0.08
m 13
Probability of total loss given water ingress location
Out of the 13 cases where the location of the water ingress was recorded, 5 were total losses while
8 were serious casualties. The table below gives the distribution between total losses and serious
casualties.
Table 10
Number of casualties split on serious
casualties and total losses.
No. 1 cargo hold
n. 2 cargo hold
n. 4 cargo hold
Total known cases
Unknown
Serious casualties
3
2
1
1
1
8
4
Total losses
4
1
0
0
0
5
3
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Annex 4
Total
12
31
8
Given water ingress in No. 1 cargo hold, the probability of total loss is taken as:
P4 =
g
4 +1
=
= 0.5
m 3 + 2 + 4 +1
Based on the information extracted, it also seems that all the total losses may be attributed to water
ingress in No. 1 cargo hold.
The probability of serious casualty, given water ingress in No. 1 cargo hold is estimated to:
P5 =
g
3+2
=
= 0.5
m 3+ 2 + 4+1
For water ingress in other cargo holds than No. 1, the probabilities of total loss based on the
casualty data are estimated to:
P6 = P8 =
g 0
= =0
m 1
Consequently, the probabilities of serious casualty given water ingress in other cargo holds than
No. 1 are estimated to:
P7 = P9 =
g 1
= =1
m 1
Probability of fatalities given water ingress location and casualty severity
In total, there are 8 total losses among the identified relevant cases. 7 of the identified hatch cover
and water ingress casualties involved fatalities. In 6 of these cases, the location of the water ingress
was recorded. The below table gives the distribution of fatal and non- fatal casualties.
Table 11
Number of fatalities split on serious casualties and total losses
n. 1 cargo hold
n. 2 cargo hold
n. 4 cargo hold
Unknown
Total
Number of fatal accidents
among the serious casualties
0
0
1
0
0
0
1
Number of fatal accidents among
the total losses
4
1
0
0
0
1
6
If assuming that the fatal accident with no details regarding water ingress location, involved
flooding of No. 1 cargo hold, the probability of fatal accident given total loss due to water ingress
in No. 1 cargo hold caused by hatch cover failure is to:
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P10 =
32
g 6
= = 0.75
m 8
It is also assumed that the 2 remaining total losses with no information about water ingress location,
follow the pattern of the 5 reported cases, hence involving water ingress in No. 1 cargo hold. In
total there were 227 fatalities in the 6 total losses involving fatalities.
The probability of non- fatal accident given total loss due to water ingress in No. 1 cargo is
estimated to:
P11 =
g 2
= = 0.25
m 8
One of the serious casualties related to No. 2 cargo hold involved 2 crew members being swept
over board, and the probability of fatal accident given serious casualty due to water ingress in No. 2
cargo hold is estimated to:
P16 =
g 1
= = 0.5
m 2
Similarly, the probability of non- fatal accident given serious casualty due to water ingress in No.2
cargo hold is estimated to:
P17 =
g 1
= = 0.5
m 2
For the remaining scenarios, no fatal accidents are recorded giving:
P12 = P14 = P18 = P20 = 0
and
P13 = P15 = P19 = P21 = 1
The above estimates based on very few or no observations are obviously encumbered with large
uncertainties. However, this does not influence on the evaluations of the risk control options, since
these are directed at No. 1 and 2 hatch covers, and the majority of the cases may be related to No. 1
hatch cover.
In Figure 10 the above results are fitted into the tree as given above.
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Annex 4
Failure of at least no.
1 hatch cover
0.77
Loss of ship
0.5
Serious
casualty
0.5
2.7·10-4
Serious
casualty
involving
water
ingress due
to hatch
cover failure
Failure of at least no. 2
hatch cover, but not
involving no.1 hatch cover
0.15
Loss of ship
0
Fatalities
0.75
Frequency Fatalities per Total losses Serious
ship year
per ship year casualties per
ship year
7.8·10-5
7.8·10-5
3.1·10-3
No fatalities
0.25
2.6·10-5
2.6·10-5
Fatalities
0
No fatalities
1
1.0·10-4
1.0·10-4
Fatalities
0
No fatalities
1
Serious
casualty
Fatalities
0.5
2.0·10-5
1
No fatalities
0.5
2.0·10-5
Loss of ship
0
Failure of any other
hatch cover, and not
no.1or no.2 hatch
cover
0.08
33
2.7·10-5
2.0·10-5
2.0·10-5
Fatalities
0
No fatalities
1
Serious
casualty
1
Fatalities
0
No fatalities
1
2.2·10-5
2.7·10
-4
2.2·10-5
3.1·10
-3
1.0·10
-4
1.6·10-4
Figure 10
Quantified breakdown of casualty data on water ingress location, loss of ship
versus serious casualty, and whether the accidents involved fatalities.
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34
Appendix 3 Hatch cover hazards and failure modes
Different Hazard Identification studies have been conducted, and Table 12 below lists the hazards
related to hatch covers and coamings as collected by MCA (2000).
Table 12
Ser.
No.
1.2.1
1.2.2
1.2.3
1.2.4
1.2.5
1.2.6
1.2.7
1.2.8
1.2.9
1.2.10
1.2.11
1.2.12
1.2.13
1.2.14
1.2.15
1.2.16
1.2.17
1.2.18
1.2.19
1.2.20
1.2.21
1.2.22
1.2.23
Hazards related to hatch covers and coamings (MCA, 2000)
Hazard
Pha Effects
Cause
se
Damage to cargo All Potential for ingress of Grab striking structure
hatch cover and
water
Heavy item of cargo dropped
coamings
Cargo gear wires cutting grooves in coaming
Low freeboard
Port Potential for ingress of Damage due to operating in frozen
water
conditions
Sea Potential for ingress of Poor design of cleaning systems
water
Hatch may become open Design compromise - insufficient hatch
allowing ingress of water cover protection from breaking seas. Sea
(substantial).
Stability loads in excess of design criteria
impaired,
Stresses Opening hatch covers when ship rolling
increased; Increased local Inappropriate heading and speed for sea
loads due to sloshing
conditions
Wastage due to lack of maintenance
Ingress of water into Design compromise - vessel type
single hold vessel
Design compromise - hatch cover rigid
vessel twists in seaway
Inappropriate loading techniques causing
coamings to distort
Low freeboard
Foundering
Loads in imposed by seas
Design compromise - insufficient hatch
cover protection
Inappropriate heading and speed for sea
conditions
Wastage due to lack of maintenance
Failure of cargo MI Potential for ingress of Inadequate maintenance management
hatch
closing
water
Faulty operation (crew error)
mechanism
Inadequate testing
Failure of cargo MI Ingress of water
Inadequate maintenance management
hatch seals and
Inadequate testing
individual cleats
1.2.24 Individual
cargo All Ingress of water
hatch cleat or seal
failure
Mechanical
Mechanical
Mechanical
Design
Mechanical
Design
Design
Operator
Operator
Wastage
Design
Design
Operator
Design
Mechanical
Design
Operator
Wastage
Maintenance
Operator
Operator
Maintenance
Operator
Design compromise - criteria does not meet Design
operational requirements
1.2.87 Stress concentration All Crack on upper deck Design compromise - inattention to detail
plate
1.2.88 at hatch opening
Lack of maintenance
corner
1.2.89 Stress concentration All Fatigue cracking at hatch Cyclic stress by wave load
coaming end
1.2.90 at hatch coaming
Lack of maintenance
end bracket
Cause Type
Design
Maintenance
Mechanical
Maintenance
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35
Quick acting cleats are used to keep the cover down as shown in Figure 11. These are manually
operated. The following modes of failure are realistic for these cleats:
1. The cleats are often not used, since they may be damaged if they are locked and the hydraulic
opening device (auto cleating) is used.
2. The flexible rubber ring is less flexible after a time in use and may fail due to fatigue.
3. The flexible rubber ring may be damaged due to high tension.
4. Pirate manufactures of the rubber ring exist and the quality varies.
The washer or spunch shown in Figure 11 should keep the hatch cover watertight at 25%
indentation. These spunches may also fail due to point 4 above.
SPUNCH
Figure 11
Cleat
Some side-rolling hatches are mounted with closing cleats (dock bolts) on top of the hatch cover at
centre line. These are often not used for the same reason as number 1 for the quick acting cleats.
Horizontal forces on the front part of the hatch coaming/cover and vertical forces on deck may give
a gap between the coaming and the cover itself as shown in Figure 12, Figure 13 and Figure 14.
Large horizontal forces on the hatch cover may move the hatch cover if the stoppers on both front
and aft part of the hatch cover are damaged, however this mode of failure is more relevant for plate
covers (which do not have stoppers) on smaller vessels.
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36
Figure 12
Gap due to horizontal forces on the hatch coaming.
Figure 13
Gap due to vertical forces on deck.
Figure 14
Gap due to horizontal forces on the hatch cover itself, moving the hatch cover.
Due to rolling seas and large container loads on top of the hatch cover fatigue cracks may be
initiated at the side brackets as presented in Figure 15. This is often observed on open bulk carriers,
but this is not regarded as a critical failure mode for ordinary bulk carriers.
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Annex 4
Figure 15
37
Observed cracks on open hatch bulk carriers.
The hatch covers are often dimensioned for bending at the middle span, and the height at the girder
ends may sometimes be reduced as shown in Figure 16. It has been indicated that the ends may not
have enough shear capacity, but we do not have any documentation verifying this failure mode.
Figure 16
Bending moment and shear force distribution for a simply supported girder
exposed to even load distribution.
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Annex 4
Appendix 4
38
Hatch covers exposed to green sea impact loads
Introduction
The present section considers hatch cover loads. Following the Derbyshire investigation, there has
been an international discussion concerning hatch cover loads, and in the following, common
procedures and recent model tests on the topic have been reviewed, and hatch cover loads have
been estimated. The green sea loading on deck is of a highly complex nature. The estimation of the
impact loads on the hatch cover in this project has been performed utilising the state-of-the-art
hydrodynamic tools for predicting seakeeping response and results from previous research.
Common procedures
International Convention on Load Lines, 1966 (ILLC 66)
The load and strength criterion according to ILLC 66 is given in regulation 16 pt. 2, and is referred
below:
(2) Where weathertight covers are of mild steel the strength shall be calculated with assumed
loads not less than 1.75 metric tons per square metre (358 pounds per square foot) on
hatchways in position 2, and the product of the maximum stress thus calculated and the factor
of 4.25 shall not exceed the minimum ultimate strength of the material. They shall be so
designed as to limit the deflection to not more than 0.0028 times the span under these loads.
Mild steel plating forming the tops of covers shall be not less in thickness than one per cent of
the spacing of stiffeners or 6 millimetres (0.24 inches) if that be greater.
IACS Unified Requirements S21 (UR S21)
The UR S21 (IACS, 1997) apply to bulk carriers contracted for construction on or after 1 July
1998, and are for hatch covers on exposed decks. Other types of loading are also to be considered,
if necessary.
A load model is defined, taking into account a number of ship parameters, and leads to pressure
loads significantly higher than that of ILLC 66. Furthermore, a strengt h criteria is given for normal
and shear stress; i.e. no to exceed 0.8 and 0.45 times the yield strength respectively. In computing
the section modulus of primary supporting members, a definition of the effective with of
compression panel flanges is included.
UR S21 is attached in Appendix 5.
Review of recent model tests
Investigations sponsored by the United Kingdom authorities
The ship model tests reported in MSC 72/4/1 were performed at the Denny Tank, the SSRC model
testing facility in Dumbarton. The Denny Tank is regarded as a conventional towing tank equipped
with the latest computer systems for data acquisition and analysis and a range of sophisticated
instrumentation. The investigation was initiated during 1998 and involved physical and numerical
model testing of a model of the hull of the MV Derbyshire with different bow configurations in
storm and hurricane sea states. The main aim was to estimate the sensitivity of deck wetness and
green water loads to variations of forward buoyancy reserve and bow height. Further, the effects of
flooding the bow spaces, forward speed, different heading angles, forecastles and breakwaters were
also investigated. The following conclusions were reached:
1) deck wetness and green seas loads are very sensitive to bow height and forward speed
2) the standards of the ILLC 66 are clearly inadequate for the conditions tested
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39
3) breakwaters tested proved to be ineffective in protecting the fore end of the ship.
These model tests were presented at the IMO MSC 70, see MSC 70/4. Based on the findings from
1998, and the recommendations from similar research, it was decided to continue the study during
1999, and to extend the study to cover a range of sizes of bulk carriers of more modern design. The
primary aim was to establish relationships between bow height and the level of wetness, and deck
and hatch cover wave loading as a function of sea state. Further, the study should provide
recommendations for minimum bow height and hatch cover design.
A representative prototype of three bulk carrier types: Handy size, Panamax and Cape size, were
investigated (32 case studies for each vessel size). The same Handy and Cape size prototype were
investigated numerically by the National Technical University of Athens (NTUA) and are reported
in MSC 70/4/6.
The vessels were tested in:
• Head sea (180° heading), bow sea (165° and 150°) and beam sea (90°).
• 2 forward speeds (3.75 and 7.5 knots) and 0 speed.
• Different bow height. The bow height was varied by fitting the bow with forecastles of different
size while keeping the freeboard constant at midship.
• Four sea states characterised by different combination of significant wave height, Hs, and
spectral peak period, Tp, were tested: [Hs=7.50m, Tp=9.60s], [Hs=10,00m, Tp=11.51s],
[Hs=12.78m, Tp=12.69s], and [Hs=13-14.99m, Tp>12.69s (exact value of Tp not given)].
• For the sea state 1 and 2 the JONSWAP spectrum for limited fetch was used. The sea state 3
was created by using a modified JONSWAP spectrum with the high frequency side
proportional to ω−4 instead of commonly used ω−5 , while the parameters α, γ, and ωp were
determined for the hurricane weather system as given in the literature (γ=5.05). The sea state 4
(sea state 3 plus a regular wave component) was designed to model hurricane waves containing
“abnormal” individual waves of over 25 metres in height.
The main conclusions of the study are as follows:
1) In the more severe sea conditions and at 7.5 knots forward speed, the No.1 and No.2 hatch
cover peak impact loads for all vessels exceed the design loads assumed by the UR S21.
2) The experimental results indicate that Cape size vessels will experience peak impact loads at
No.1 hatch cover larger than the design loads assumed in the UR S21. This situation is
worsened if the vessel is making way at any significant speed.
3) The occurrence of wetness and green sea loads systematically increases with increasing vessel
size. Only in the moderate storm (Hs=7.5m), the Handy size bulk carrier is more affected by
wetness and green sea loads than the larger ships. Further, wetness and green sea loads
systematically increases with severity of the sea, decreasing forward freeboard and increasing
ship speed.
4) Forward speed is the most critical factor in terms of wetness and associated green water loads
on deck.
5) During a wetness event associated with extreme peak loads, the measured load cell pressure on
the No. 1 hatch cover of the Cape size ship was, generally, about 40% higher than the depth of
green water measured by wave probes (with null forward speed).
6) Forecastles reduce deck wetness and frequency of hatch loads, but do not reduce peak loads on
forward hatch covers sufficiently to bring the loads within current design standards.
Comparison of MARIN and SSRC tests
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40
As mentioned in MSC 72/4/1 a series of tests were carried out at MARIN (Wageningen, the
Netherlands) in order to validate the work done at the SSRC, and in general the comparison has
shown good agreement, indicating the validity of seakeeping model tests at the SSRC. These results
will be reported to IMO separately. A preliminary documentation discussing correlation between
the results was received by DNV and is summarised below.
The basis (Derbyshire) hull and bow used by the SSRC was the same as tested in MARIN as well
as the test setup (It is not specific mentioned what type of the wave spectrum was used by
MARIN.). However, the size of the tank was different. Further, the MARIN tests and the SSRC
tests had different duration. The MARIN test was longer than each of the SSRC tests. For the
hurricane conditions the SSRC adopted a sea state Hs=13-14.99m, while MARIN used a sea state
Hs=15m.
The SSRC carried out 4 tests with the bow in intact condition for each of two sea states:
Hs=12.78m (heading 180°) and Hs=13-14.99m (heading 180°). In order to obtain a longer record
the tests for each sea state were added together prior to the analysis and compared with a MARIN
test carried out for Hs=12.78m (heading 180°) and Hs=15.00m (heading 180°), respectively.
Further, 2 tests of the hull with the bow in damaged condition (flooded) were carried out by SSRC
in the sea state Hs=12.78m (heading 180°), and two tests in the sea state with Hs=13-14.99m
(heading 180°). The same as for the intact condition for each sea state in order to obtain a longer
record the tests were added together before the analysis and compared with the MARIN bow in
damaged test with Hs=12.78m done at 225° heading and with Hs=15m (heading 180°).
All the facts mentioned above can explain some small discrepancies between the SSRC and the
MARIN test results.
Summary Comments
•
•
•
•
•
In general a series of tests carried out at MARIN as well as numerical calculations carried out
by NTUA (for the same Handy and Cape size prototype) show good agreement, indicating the
validity of seakeeping model tests at the SSRC. The hull and bow used by MARIN as well as
the test setup was the same as tested in the SSRC. However, number of tests is limited. It seems
that further tests are necessary in order to reach a firm conclusio n concerning discrepancies
between the model test results and the design loads assumed by the UR S21.
The model test conditions should be further checked in order to confirm the validity of the
investigations, e.g. wave reflections in the basin.
The Cape size bulk carrier model used in phase 2 (1999) had a large bulb. Other bulk carrier
designs should also be investigated.
The highest sea state applied by the SSRC: Hs=13-14.99m (Tp>12.69s) is lower than the 20year extreme for the North Atlantic according to the DNV Classification Note 30.5
“Environmental Conditions and Environmental Loads” , Hs=15.63m (Tz=12.61s, Tz denotes
zero-crossing wave period).
The severe sea state 3 was created by using a modified JONSWAP spectrum with the high
frequency side proportional to ω−4 instead of the high frequency side proportional to ω−5 ,
γ=5.05, while the parameters α, and ωp were determined for the hurricane weather system as
given in the literature. Both the Pierson-Moskowitz spectrum and the JONSWAP spectrum
decays according to ω−5 . Further, the standard JONSWAP spectrum uses γ=3.3 being less
peaked than the spectrum with γ=5.05. The modified JONSWAP spectrum with the high
frequency side proportional to ω−4 has much more energy in the high frequency tail. For
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41
structures sensitive to the high frequency range a response is significantly increased by using
the wave spectrum with the high frequency side proportional to ω−4 instead of ω−5 . For
responses sensitive to the body of the spectrum the JONSWAP spectrum decaying according
to ω−4 will lead to lower response values than obtained by using the standard JONSWAP
spectrum. Furthermore, the model tests were carried out for unidirectional wave systems.
Unidirectional waves tend to be larger than short crested waves leading to higher impact loads.
It should be noticed that the UR S21 is based on the conventional JONSWAP spectrum, and
uses short crested seas where the wave energy is spread around a dominant angle.
The hydrodynamic analysis
The green sea loading on deck are caused by water flooding, due to the fact that the water (wave)
surface is above deck (freeboard). The relative motion between the water surface and the vessel has
been calculated using existing models of a capesize bulk carrier. The main ship particulars and
characteristics are presented in Table 13.
Table 13
Main ship particulars in the hydrodynamic analysis (and structural
analysis in parenthesis).
Length between perpendiculars, Lpp [m]
Water line length, L [m]
Breadth moulded, B [m]
Draught, T [m]
Depth moulded, D [m]
Displacement, ∆ [tons]
Block Coefficient, CB [-]
Metacentre height, GM [m]
Radius of gyration in roll about COG, r44 [m]
275 (271)
283 (283)
47 (45)
16.8 (18.15)
24.8 (24.6)
184 800 (189197)
0.83 (0.85)
10.27
15.23
An existing capesize bulk carrier has been used as a case study in the structural analysis. This
vessel is slightly different from the vessel used in the hydrodynamic analysis. The main ship
particulars of the vessel used in the structural analysis are written in parenthesis in Table 13.
The hydrodynamic models used in this analysis consist of a DNV-SWAN model (DNV, 1997) as
shown in Figure 17 and a WADAM model (DNV, 1994).
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Figure 17
42
The DNV-SWAN mesh used in the linear analysis of the relative motions.
The WADAM model has been used to calculate:
• Transfer functions for relative motion of a few points in the bow area based on an undisturbed
incoming wave (input to PROBAN, (DNV, 1992))
• Transfer functions for relative motion of a few points in the bow area based on a disturbed wave
• Acceleration in the bow area for use in a fatigue evaluation.
The DNV-SWAN model has been used to calculate:
• Transfer functions for relative motion of a few points in the bow area based on a disturbed wave
used to establish design waves.
• Time series of linear relative motion in head sea at No. 1 hatch cover based on a specific design
wave
• Time series of non- linear relative motion in head sea at No. 1 hatch cover based on the same
specific design wave.
The analysis is performed with zero speed. Head sea conditions are regarded to be worst with
respect to deck wetness according MSC 72/4/1/Add. 1. In head sea, the realistic maximum speed in
extreme weather conditions will be low (and even negative) and is here assumed to be zero. Less
extreme weather conditions with significant forward speed may also give significant deck loads,
but this is not further treated herein.
According to MSC 72/4/1: “Forward speed is the most critical factor in terms of wetness and
associated green water loads on deck”. Based on this conclusion, the DNV-SWAN model is used to
briefly investigate the effect of forward speed as presented in Section 0. The effect of vessel size is
investigated briefly in Section 0.
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43
Relative motion
The locations where the relative motions are calculated are presented in the Table 14.
Table 14
Point
1
2
3
4
5
6
Points for relative motion calculations
Description
Bow
Middle of hatch, starboard
Middle of hatch, port side
Aft of hatch, port side
Forward of hatch, port side
Middle of hatch at CL
x from aft perpendicular [m]
271
247
247
239
254
247
y from centre line [m]
0
-19.5
19.5
21.0
16.5
0
WADAM (DNV, 1994) and DNV-SWAN (DNV, 1997) have been used to assess the relative
motion. In DNV-SWAN the relative motion has been calculated directly, while WADAM
calculates the surface elevation. POSTRESP (DNV, 1995a) is then used to establish the transfer
function of relative motion based on the difference between the vessel motion and the surface
elevation. POSTRESP may also be used to calculate the relative motion based on an undisturbed
wave. The latter is done for WADAM for the sake of comparison at the side of the vessel.
Results for different environmental conditions
The extreme relative motions at the different locations are presented in the Table 15. Relative
motions with a return period of 20 years are given. Different analysis conditions have been used
with respect to heading and directional wave energy spreading represented by cos2 function (also
referred to as short crested sea, alternatively long crested sea if no spreading function has been
used).
In the long term joint environmental description, the direction variable is discretified into 12 sectors
from 0° to 330°. The probability of occurrence for each sector is the same (1/12). UDD is an
abbreviation of this uniform directional distribution. UDD with cos2 spreading is consistent with
the way main design loads normally are derived by Classification Societies. For extreme sea states,
wave spreading represented by cos2 is not likely. However, e.g. head sea sector with cos2 may be
interpreted as a directional distribution with head sea as the main direction.
Table 15
Relative motions with a return period of 20 years in the North
Atlantic wave environment using Pierson-Moskowitz wave
spectrum.
Condition
WADAM, head sea, long crested, disturbed
WADAM, head sea, long crested, undisturbed
SWAN, head sea, long crested, disturbed
WADAM, head sea, cos2 , disturbed
WADAM, head sea, cos2 , undisturbed
SWAN, head sea, cos2 , disturbed
WADAM, UDD, cos2 , disturbed
WADAM, UDD, cos2 , undisturbed
SWAN, UDD, cos2 , disturbed
Point 1
27.8
23.5
28.0
27.5
23.2
27.6
25.2
21.3
25.3
Point 2
19.8
17.7
21.0
20.7
18.2
21.6
23.2
18.9
22.5
Point 3
19.8
17.7
21.0
20.7
18.2
21.6
23.2
18.9
22.5
Point 4
17.1
15.8
17.1
18.7
16.6
18.6
22.9
18.3
22.0
Point 5
22.3
19.6
22.9
22.8
19.8
23.2
23.7
19.6
22.9
Results for both disturbed and undisturbed relative motion are presented. Disturbed/undisturbed
means that the diffracted wave due to the presence of the ship is/is not taken into account as shown
schematically in Figure 18.
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Figure 18
44
Disturbed versus undisturbed wave.
Table 15 is presented for the purpose of comparison between disturbed and undisturbed relative
motion and WADAM and DNV-SWAN. The results of the extreme relative motions in the bow
area show that:
• Results from WADAM agree well with DNV-SWAN for zero speed
• Relative motion is larger for the disturbed wave compared to the undisturbed wave
• Long crested head sea and short crested head sea gave approximately the same values
• Head sea gives larger relatives motion than the uniform directional distribution at centre line
• Head sea gives less relative motions than the uniform directional distribution at the ship side
(due to roll motion).
The relative motion at the ship side will overpredict the loads on deck especially at the centre line,
since the roll component and pile up of water at the side for short small waves becomes significant
in the statistical calculations. In addition, it will not give a good estimate of the average pressure
over the hatch cover at centre line. The undisturbed relative motion in centre line at middle of the
No. 1 hatch is used instead to determine the deck loads. The disturbed wave can not be used at this
position, since it is inside the vessel’s boundaries. The transfer functions for the undisturbed
relative motion are shown in Figure 19.
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Figure 19
45
Transfer function of relative motion at the centre of No. 1 hatch cover (point 6).
Head sea is 180°° .
The relative motion in point 6 has been calculated for Pierson-Moskowitz (PM), Jonswap (JS) and
two versions of the Gamma (G) wave spectrum (with ω-4 and ω-5 , which affects the tail behaviour).
According to the Annex to MSC 72/4/1, ω-4 was found to fit relevant data for hurricane extreme
waves better than ω-5 . With ω-5 , the Gamma wave spectrum should be practically the same as the
PM for fully developed sea (provided that peakness parameter γ = 1).
Undisturbed relative motion at the middle of the hatch cover has been calculated for the North
Atlantic (NA) and World-Wide (WW) scatter diagram taken from (DNV, 1998). The applied wave
spectrum has been used for all sea states in the scatter diagrams. The relative motion is calculated
in short crested and long crested head sea conditions. The uniform directional distribution is
calculated with short crested sea. The middle of the No. 1 hatch cover for point 6 is located
approximately 0.91⋅Lpp from aft perpendicular.
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Table 16
46
Undisturbed relative motion at the centre of No. 1 hatch cover (point 6).
Condition
Point 6
PM
16.68
17.70
17.83
15.15
15.52
15.71
WADAM, UDD, cos2 , NA
WADAM, head sea, long crested, NA
WADAM, head sea, cos2 , NA
WADAM, UDD, cos2 , WW
WADAM, head sea, long crested, WW
WADAM, head sea, cos2 , WW
Point 6
JS, γ = 3.3
17.81
18.78
18.98
16.44
16.42
16.78
Point 6
Gamma, ω -5
16.78
17.76
17.91
15.25
15.57
15.78
Point 6
Gamma, ω -4
14.47
15.63
15.54
13.36
14.02
14.03
It is seen from Table 16 that:
• The head sea condition is more severe than the uniform directional distribution with respect to
water on deck (roll motion is not contributing at centre line).
• Jonswap gives higher values than PM and Gamma
• Gamma with ω-5 gives higher values than Gamma with ω-4
• Gamma with ω-5 gives the approximately the same values as PM
The difference in relative motion with respect to the choice of wave spectrum is seen to be
significant, and the relative difference of water head on deck is then even more significant. The
Jonswap wave spectrum used in Table 16 is not realistic for all sea states. A better, but more time
consuming process is to establish a specific γ parameter for each sea state. According to in-house
experience with other scatter diagrams, the results are expected to be close to the PM results. In the
development of the UR S21 (IACS, 1997), ω-5 has been used. Based on the above, it has been
decided to use the PM wave spectrum herein.
Bulk carriers trade all over the world, and the bulk carriers, which have been reported lost due to
collapse of the hatch cover, have sunk in different parts of the world. To verify observed collapse
frequencies of hatch covers, the WW scatter diagram has been chosen as the most representative for
the extreme load calculations. UDD, PM and cos2 have been assumed. The possibility for the
master to go up against the waves in extreme weather conditions has not been taken into account.
The resulting load will be referred to as the verification load.
To verify the observed accidents may be a different matter than establishing design loads. In order
to establish design loads, the NA wave environment is recommended. The long-term distribution of
the undisturbed relative motion at point 6 for the NA wave environment, UDD, PM and cos2 is
presented in Figure 20. The long-term distribution may be represented by a 2-parameter Weibull
distribution with:
•
•
•
Long term response period, TR, = 9.253 seconds
Weibull slope parameter, h, = 0.9613
Weibull scale parameter, q, = 0.8233
The relative motion at No. 1 hatch cover with a return period of n years can then be calculated as:
1
 n ⋅ 365 ⋅ 24 ⋅ 3600  h

RELU 6 = q ⋅  ln
TR


n
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Figure 20
47
Long-term distribution of relative motion at the centre of No. 1 hatch cover.
North Atlantic wave environment, UDD (blue line) and short crested waves are
used. Short crested head sea is presented for comparison.
Non-linear correction factor to relative motion in the bow area.
A design wave approach is used to estimate a non- linear correction factor for the relative motion.
The design wave amplitude is defined as the ratio between the 20 year design value and the
maximum peak in the transfer function of the disturbed relative motion in head sea. The design
wave amplitude was calculated for point 2 in Table 15 (disturbed relative motion at side shell at
middle of No. 1 hatch, NA, PM, UDD and cos2 ):
Adesign =
rel 20 y 22.48
=
= 12.42
trf
rel max
1.810
The corresponding period was 11.0 seconds for the peak in the transfer function. A check of the
steepness of this design wave is made. The steepness should be less than 1/7 = 0.14.
S=
2 ⋅ Adesign
2 ⋅ 12.42
=
= 0.131
g
9.81
2
2
⋅T
⋅ 11
2 ⋅π
2 ⋅ 3.1415
< 0.14
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48
The DNV-SWAN model was then run with this regular wave both linearly and non- linearly and the
relative motions for point 2 (middle of the hatch at side) were compared. The non- linear correction
factor was estimated to be:
k nl =
rel 2 nl
= 0.91
rel 2 l
This factor derived for point 2 will be used in the deck load calculations of the No. 1. hatch cover at
point 6. Point 2 was used since this relative motion could be taken directly from the time series in
DNV-SWAN, while relative motion in point 6 had to be established based on combined motions
and waves in the non- linear analysis (which is more time consuming).
Design load calculation
The load model presented below is a simplified one. The green water impact pressure on top of the
hatch cover is found by adjusting the static pressure based on relative vertical distance between the
water surface and the hatch cover. Physics of green water motion is more complex, and further
studies of the water impact on deck are discussed in (Nestegård and Krokstad, 2000).
The pressure on top of the hatch cover may be calculated in the following way:
(
)
P d = k dyn ⋅ ρ ⋅ g ⋅ k nl ⋅ relu 6 − (h − T ) [kN/m2 ]
Where:
kdyn
knl
ρ
g
h
T
relu6
= dynamic factor relative to static pressure height
= non- linear factor on relative motion for point 2
= 1.025 tons/m3
= gravity, 9.81m/s2
= height from base line to hatch cover
= draught at the hatch cover
= linear relative motion in head sea at point 6 (middle of No. 1 hatch cover at CL)
The kdyn is a dynamic factor, which takes into account the impact of water on the deck/cover. This
factor is taken from comparison between loads on the No. 1 hatch cover based on model tests and
the measured height of water on deck. The UK investigations (MSC 72/4/1) indicate a factor of 1.4.
However, this factor will be dependent on the configuration in the bow area (e.g. with or without
large bulb).
It has been assumed that the impact pressure is evenly distributed over the whole hatch cover
simultaneously. This is not correct due to the complicated nature of the water impact on deck of a
pitching vessel, however this model should be suitable for this investigation.
A comparison between the directly calculated pressure above and the UR S21 is presented in
Section 0.
Comparison study between directly calculated pressure and UR S21
The comparison is made for the capesize vessel chosen for the structural analysis. This vessel is
regarded as typical. The vessel’s characteristics are given in Table 13 in parenthesis. The hatch
coaming is 1.20m high at centre line, the hatch cover itself is 0.89m high and the deck chamber is
0.90m. The vessel has B-60 freeboard of 6.483m and a spring of 0.855m at the middle of the No. 1.
hatch cover. This gives a height from base line to hatch cover of 28.478m.
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49
Using the expression for the dynamic pressure in Section 0 and assuming UDD, PM wave spectrum
and short crested sea, the verification pressure is calculated to:
P d = 1.4 ⋅1.025 ⋅ 9.81 ⋅ (0.91 ⋅ 15.15 − (28.48 − 18.15) ) = 49 kN/m2
for World-Wide
While the design pressure is calculated to:
P d = 1.4 ⋅ 1.025 ⋅ 9.81 ⋅ (0.91 ⋅16.68 − ( 28.48 − 18.15) ) = 68 kN/m2
for North Atlantic
The UR S21 gives 52 kN/m2 for the same configuration.
Even if these loads are calculated for one specific capesize vessel based on a slightly different
hydrodynamic model, it indicated that the direct calculations may give significant higher pressures
than the UR S21 standard based on the North Atlantic wave environment.
The dynamic factor of 1.4 is connected with some uncertainties, however the pressure of 68 kN/m2
is proposed as the new design pressure in the structural assessment. This corresponds to an
increase of ∼30% compared to UR S21. If e.g. the dynamic factor is reduced to 1.2 based on new
model tests of capesize vessels, the proposed design pressure is reduced to 58 kN/m2 , which is an
increase of 13% compared to UR S21. For ships wit h totally flat deck it is likely that the dynamic
factor may be as low as 1.0.
Horizontal loads on coamings from green sea
The green sea loads on deck may cause large horizontal forces on the hatch coaming. In the paper
by Faulkner et al (1996) following the Derbyshire accident, the horizontal load is given as:


g ⋅T 2
2


1
1
λ
1


p h = ⋅ ρ ⋅ v 2 = ⋅ ρ ⋅ 1.2 ⋅ + U  = ⋅ ρ ⋅ 1.2 ⋅ 2 ⋅ π + U 


2
2
T
2
T






1
g
= ρ ⋅ (0.6 ⋅ T + U ) 2
2
π
Where:
ρ
= The density of water, 1.025 [tonns/m3 ]
T
= The wave period [s]
λ
= The wave length [m]
v
= The relative velocity [m/s]
λ/T
= Crest velocity [m/s]
U
= Vessel speed [m/s]
2
kN/m2
The expression inside the parenthesis represents the relative velocity between the wave and the
vessel. The 1.2 factor “is to allow for flow channelling and wind augment” according to Faulkner et
al (1996). The first term represents the propagating wave profile velocity, while the next term
represents the speed of the vessel heading into the waves. This expression is rather theoretical
assuming that the vessel and the object hit by the wave do not disturb the wave itself. It is
interpreted as the average pressure due to the drag force for a small horizontal bracing moving into
the waves, neglecting slamming effects. However, to our knowledge there exist no well established
procedures for predicting this horizontal green water impact forces, and the present approach by
Faulkner may be used as a first estimate. Schematically the impact is shown in Figure 21.
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Figure 21
50
Horizontal impact between vessel and a propagating wave.
From the design wave approach in the section 0, the peak in the transfer function was identified
with a wave period of 11 seconds for zero vessel speed. The ship speed in extreme wave conditions
will be low and is set to zero in this case, which gives a horizontal pressure of 218 kN/m2 based on
the formula.
The horizontal pressure may be applied to the front coaming and front part of the hatch cover in the
finite element model in order to evaluate if the linear elastic stress gives a significant stress zone
exceeding the yield stress. Yield followed by permanent deformations of the coaming and
supporting structure might lead to a gap between the coaming and hatch cover. This may be a
relevant failure mode related to water entry into the first cargo hold.
Structures in way of the front hatch coaming may reduce the horizontal pressure significantly.
Furthermore this pressure does not necessarily occur simultaneously with the vertical impact loads
on deck and hatch cover. When the hatch is designed for cargo on top such as containers, it is
common to use a wave breaker in front of the hatch to reduce the horizontal forces.
Inertia force of the hatch cover and possible fatigue damage.
POSTRESP is used to calculate the vertical acceleration from WADAM. The acceleration will
induce stresses in the hatch cover due to the inertia forces. The resulting stresses are used in a
simplified fatigue check.
The accelerations are calculated at the middle of the hatch on port side at a probability level of
exceedance equal to 10-4 (most probably largest response based on 10000 cycles, which
corresponds to a return period of 1 day). The position and results are shown in Table 17.
Table 17
Acceleration for fatigue evaluation calculated in NA wave
environment with UDD, cos 2 directional energy spreading and PM
wave spectrum.
x from AP [m]
y from CL [m]
247
5.0
Response
period, TR [s]
11.2
Weibull
slope
parameter, h [-]
1.01
Acceleration at 10-4
level [m/s 2 ]
2.37
The maximum principal stress in the hatch cover is taken from the finite element model. Based on
the eigenweight of the hatch cover times the acceleration above, the maximum principal stress was
calculated to 5.3 MPa in the plate flange at centre line at the middle of the hatch cover. Assuming a
local stress concentration factor of 2.5 (due to geometry and weld effects), the total stress range is
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51
26.5 MPa. The maximum allowable stress from DNV Classification Note 30.7 [6] is 163 MPa
based on:
•
•
•
•
probability level of exceedance equal to 10-4
0.7⋅108 cycles (20 year design life)
SN-curve for welded joints in corrosive environment
Weibull parameter of 1.01
We can conclude that fatigue in the hatch cover plate flange due to inertia forces of the eigenweight
is not a realistic failure mode.
Effect of forward speed
The Capesize model was run by DNV-SWAN in head sea for three different vessel speeds:
• 0 knots
• 3.75 knots
• 7.5 knots.
The transfer functions from the linear analysis was combined with PM wave spectrum and NA
wave environment, and the 20 year values were calculated. Point 2 representing the disturbed
relative motion at the side shell at the middle of the No. 1 hatch has been used as reference. The
peak of the transfer function was identified and the design wave established. For the design wave,
both a linear and a non- linear analysis was performed and the non-linear correction factor was
determined. The non- linear factor is multiplied with the linear 20 year value and finally an estimate
of the non- linear relative motion is derived as presented in Table 18.
Table 18
Forward speed effect on relative disturbed motion at point 2
Speed
Linear relative motion: NA, 20 year, head sea,
long crested [m]
Relative linear factor
Peak period in transfer function, T [s]
Peak relative motio n [m/m]
Design wave, A [m]
Non-linear factor
Non-linear relative motion:
Relative non- linear factor
0 knots
21.00
3.75 knots
24.25
7.5 knots
28.11
1.00
11.0
1.80
11.7
0.95
19.85
1.00
1.15
12.6
2.19
11.1
0.86
20.86
1.05
1.34
13.7
2.58
10.9
0.82
23.10
1.16
The relative motion increased with increasing speed. The design pressure on the hatch cover No. 1
will be affected accordingly. It is also noted that the linear analysis gave higher relative factors to
the speed effect than the non-linear analysis.
If we assume that the height from still water line to top of the hatch cover is the same as used
previously (10.33m) the pressure will be increased by:
• 11% increasing the speed from 0 to 3.75 knots
• 34% increasing the speed from 0 to 7.5 knots
This is a significant increase. According to the model test referred to in MSC 72/4/1 the increase
from 0 to 7.5 knots gave an increase of 25% for hatch No. 1 and 50% for hatch No. 2. It remains to
determine the realistic vessel speeds in different wave heights and different headings for typical
bulk vessels.
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52
The effect of vessel size
Four vessels with different dimensions have been run with WADAM in order to see if there is a
significant difference in the relative motions at the No. 1 hatch cover based on the approach used
for the Capesize vessel. The vessel’s characteristics are presented in Table 19.
Table 19
Type
Capesize
Panamax
Handymax
Handymax
Main ship particulars in the hydrodynamic WADAM analyses
Lpp [m] B [m]
271
217
180
146
47.00
32.25
30.40
22.86
T [m]
∇ [m3 ]
16.83
13.75
11.62
10.35
180465
81471
51648
28002
GMT
GML
r44 [m]
r55 [m]
[m]
[m]
10.27
349.6
15.2
67.9
6.47
292.9
9.2
57.3
3.75
223.6
11.3
43.6
5.00
167.3
8.6
36.0
The relative motion is calculate along centre line of the ship based on both North Atlantic and
World-Wide scatter diagram. Pierson-Moskowitz wave specrum, cos2 directional wave energy
spreading and uniform directional distribution is used. The results are shown in Figure 22 to Figure
25. In addition to the relative motions, the vertical distance between the still water line and the deck
at centre line including forecastle, poop, shear and bulwarks are indicated and referred to as
“freeboard”.
Based on the figures the following is observed:
• The relative motions all along the vessel increase with increasing vessel length.
• The difference between North Atlantic and World-Wide wave environment increase with
increasing vessel size.
Based on the formula in Section 0, the pressure on top of the No. 1 hatch cover based on the North
Atlantic wave environment is calculated to:
• 68.2 kN/m2 for the Capesize (middle of No. 1 hatch cover at 0.91Lpp )
• 59.2 kN/m2 for the Panamax (middle of No. 1 hatch cover at 0.90Lpp )
• 46.5 kN/m2 for the large Handymax (middle of No. 1 hatch cover at 0.86Lpp ) and
• 34.1 kN/m2 for the small Handymax bulk carrier (middle of No. 1 hatch cover at 0.84Lpp )
Hence, the No. 1 hatch cover pressure increases for increasing vessel size.
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Linear relative motion at centre line with a return period of 20 years, uniform
2
directional distribution and cos spreading for a CAPESIZE
22,0
20,0
18,0
16,0
14,0
North Atlantic
12,0
World Wide
10,0
Freeboard
8,0
6,0
4,0
2,0
0,0
0,0
0,1
0,2
0,3
0,4
0,5
0,6
0,7
0,8
0,9
1,0
1,1
Distance from AP/L pp
Figure 22
Relative motion for a Capesize bulk carrier.
2
directional distribution and cos spreading for a PANAMAX
22,0
20,0
18,0
16,0
14,0
North Atlantic
12,0
World Wide
10,0
Freeboard
8,0
6,0
4,0
2,0
0,0
0,0
0,1
0,2
0,3
0,4
0,5
0,6
0,7
0,8
0,9
1,0
1,1
Distance from AP/Lpp
Figure 23
Relative motion for a Panamax bulk carrier.
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2
directional distribution and cos spreading for a large HANDYMAX
22,0
20,0
18,0
16,0
14,0
North Atlantic
12,0
World Wide
10,0
Freeboard
8,0
6,0
4,0
2,0
0,0
0,0
0,1
0,2
0,3
0,4
0,5
0,6
0,7
0,8
0,9
1,0
1,1
Figure 24
Relative motion for a “large” Handymax bulk carrier.
2
directional distribution and cos spreading for a small HANDYMAX
22,0
20,0
18,0
16,0
14,0
North Atlantic
World Wide
12,0
10,0
Freeboard
8,0
6,0
4,0
2,0
0,0
0,0
0,1
0,2
0,3
0,4
0,5
0,6
0,7
0,8
0,9
1,0
1,1
Figure 25
Relative motion for a “small” Handymax bulk carrier.
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Appendix 5 IACS Unified Requirements S21
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Appendix 6 Analysis of Hatch Cover Capacity and Approximate Formulation
for Pressure Capacity of Hatch Cover
7.1
Finite Element Model
A linear finite element analysis of the hatch, including coaming and supporting structure has been
carried out. The hatch cover is divided at the ship centre line, and rolls sideways to open. When
closed, it rests on support pads along the coaming. The support pads are assumed to transfer
vertical loads only, whereas fixation stoppers are used to prevent horizontal movement of the hatch.
Two stoppers are used in the transverse direction relatively near the centre line of the ship, whereas
a single stopper is used in the longitudinal direction aft position. With such support conditions,
each half of the hatch can be assumed simply supported along three edges, and free along CL.
Shell elements (8-nodes) have been used in combinations with beam elements representing plate
stiffeners and flanges. Symmetry has been assumed, and only the port side has been modelled.
Boundary conditions ensuring symmetry has been applied at the centre line. The model is shown in
Figure 26, where the direction of the coordinate system is included, positive x-directions towards
the bow. Blue colour is used for the hatch itself, and yellow is used for the coaming and adjacent
part of the ship structure; i.e. deck plating and transverse bulkhead. The main purpose of this
model is to:
−
−
−
−
Enable stress checks in the hatch cover for different combinations of loading.
Enable stress checks in the hatch coaming and adjacent deck structure for different
combinations of loading. Due to model simplifications, such stresses should only be used
as an indication of potentially highly utilised areas.
Estimation of transversal stresses in the hatch cover, to be used as input for the computation
of effective flange in the capacity evaluations.
Estimation of fatigue stresses.
A linear finite element analysis is, however, not considered adequate for ultimate capacity
calculations, and has not been used directly with regards to the capacity formulation in the
reliability analysis. This is because considerable load redistribution due to plastic deformation can
take place, and because buckling needs to be considered. Capacity formulations are further
discussed in section 0.
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Seen from above
Hatch cover
Coaming
Ship deck
Bukhead,
aft, beam
model
z
"Cantilever",
supporting
coaming
Bukhead,
fore
x
y
Hatch cover,
webs of girders
z
y
Seen from underneath
Figure 26
down).
x
Finite element model (if the figures are hard to read, try to turn them upside
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7.2
60
Capacity model
In the assessment of the hatch cover girder strength, the contribution from the panel stiffeners,
including the attached flange, to the girder section modulus was deducted. Further in the
assessment, the strength contribution from the girder adjacent to the hatch side was disregarded.
The effectiveness of the attached flange of the top plate has been determined in accordance with
UR S21.
The hatch cover strength was assessed by a simplified grillage model where the transverse girders
of the transverse girders of the hatch covers are assumed rigid. The strength contribution by the
axial girders has been determined as an equivalent girder located at the center line.
q1
q2
q3
p
s1
Figure 27
s2
s3
Hatch cover model used
The capacity has been estimated as outlined below.
An equivalent section modulus, Zeq, was estimated as:
2
Z eq = Z1 + Z 2
3
where Z1 , is the section modulus of girder No. 1, and Z2 is the section modulus of girder No. 2.
The hatch cover loading was represented as an equivalent line load model acting at the center line:
q=
11
2
1
q1 + q 2 + q3
12
3
3
where
q1 = p ⋅ s1
q2 = p ⋅ s2
q3 = p ⋅ s3
where p is the pressure acting on the hatch cover, and s1 , s2 , and s3 are indicated in the above figure.
The required section modulus is given by:
q ⋅l2
Z=
8 ⋅σ y
where l is the span length between the supports, found to be 15.7 m, and σy is the yield stress.
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An estimated collapse line load and hence critical hatch cover pressure, pc, may be achieved
according to:
pc =
8 ⋅ Z ⋅σy
l
2
⋅
1
11
2
1
s1 + s 2 + s 3
12
3
3
The model gave ultimate capacities for the investigated design as shown in the table below.
Table 20
Structural analysis results
Strength (kN/m2 )
Case ID
Base case (ILLC 66)
UR S21
UR S21 + 30%
35.8
65
84.5
The base case design was assessed by non- linear structural analysis, which gave an ultimate
capacity of 35 kN/m2 (Østvold, 2001), see the figure below.
2
35
p [kN/m ]
30
25
20
15
10
5
0
0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
w (mid span)/0.0056L
Figure 28
Lateral pressure versus scaled mid-span lateral displacement of the centre
hatch cover longitudinal girder (Østvold, 2001).
7.3
Capacity Evaluations, Initial Design
7.3.1
ILLC 66, capacity check
Although the initial design presumably complies with the ILLC 66, a check based on loads and
allowable stress has been performed.
The design load is 1.75 tonn/m2 . This is representative for cargo loads on top of the hatch, and is
not specified as a water on deck load case. The eigenweight of approximately 200 kg/m2 has been
included.
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62
A stress check is specified, where the calculated stress multiplied by the factor of 4.25 should not
exceed the ultimate stress. The ultimate stress for the steel is between 440MPa and 570MPa.
Assuming the ultimate stress in the middle between these values, the calculated stress should
(conservatively) be checked against 440/4.25=104MPa.
A simple hand-calculation, considering a simply supported beam, confirms that the design comply
with ILLC 66. (A Von Mises stress plot from the finite element analysis for the top plate is
included in Figure 29 for information only. The plot indicates a slight violation of the stress
criterion in a local area at the centre of the hatch.)
Figure 29
7.3.4
Stress check, ILLC 66, allowable stress of 104 MPa.
Evaluation of stresses in sub-structure
The finite element model has been used to evaluate the overall stress level in the hatch coaming and
adjacent deck plating. Both horizontal and vertical loading has been considered. A vertical load
has been applied on the hatch and deck plating, and a horizontal load of 218 kN/m2 (see Appendix
4) has been applied to the front coaming and over the height of the hatch cover front. The model is
not sufficiently refined to give detailed stress results, but is used to get an indication of the stress
level. The Von Mises stress has been considered and evaluated against a criterion of 0.8 times the
yield stress. The overall stress level is well below this threshold, but relatively high utilisation is
seen at the lower flange of the cantilever spanning from the transverse bulkhead at the front to the
coaming. High stresses are also obtained at the single hatch stopper for horizontal load in the
longitudinal direction, located at the aft coaming. Plated buckling has not been evaluated.
A potential increase in the load requirement must be considered also in the design of the supporting
structure.
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Appendix 7 Analysis of hatch cover reliability
Introduction
The reliability analysis as presented here is intended to be representative for typical Capesize bulk
carriers operating world-wide and designed according to the ILLC 66 (International Conference on
Load Lines, 1996). The failure criterion is formulated as a limit state function:
g = pc − pd
Where the failure domain is g < 0 . Failure occurs when the pressure load p d on the hatch exceeds
the pressure load capacity p c related to moment failure at the centre of the longitudinal girders.
The main purpose of the analysis is to quantify the relative change in probability of failure as a
consequence of a design modification. As indicated in section 2.4, the reliability analysis is carried
out for three different designs as basis for the cost effectiveness analysis:
1. Base case, initial design according to ILLC 66
2. Design according to Unified Requirements S21.
3. Design according to Unified Requirements S21, with a load increase of 30%.
The annual probability of failure has been calculated for each case, and effort has been made in
order to apply appropriate physical models and realistic uncertainty modelling. It is evident that the
absolute level of the failure probability is rather sensitive to the distribution input and the analysis
models used. In our case, however, only the relative change in probability of failure is used as
input to the cost effectiveness analysis, and the accuracy of the absolute values of failure
probability therefore becomes somewhat less critical as long as the results are at a realistic level.
The calculated annual probability of failure for the base case has been compared with the empirical
probability of failure from historical data, and this comparison serves to some extent as a
verification of the model. Although the failure data has been sorted to the extent possible, it is not
realistic to believe that all incidents are related to the failure mode applied in the reliability analysis.
There may be failures due to gross errors during design, fabrication, installation and operation, and
distribution models representative for ordinary conditions cannot model such failures. The
empirical data may also include other failure modes than the one applied in the reliability analysis,
but this may not necessarily be a problem if the load and capacity models for such failures are
comparable with the ones used. The robustness of the conclusions from the cost effectiveness
analysis should be interpreted with this in mind.
Load pressure model
The load model is essentially based on the conclusions in Appendix 4. The probabilistic analysis
program PROBAN (DNV, 1992) uses the transfer function for the relative motion between the
wave elevation and the centre of the hatch cover. The short term response is calculated using the
Pierson-Moskowitz wave spectrum and short-crested sea with a cos2 distribution. All headings are
considered, with equal probability of occurrence. A sensitivity analysis has been carried out
assuming head sea only and long-crested sea.
A Gumbel extreme value distribution in a random sea state is formulated based on the assumption
of Rayleigh distributed maxima. Further, the annual extreme value distribution of the relative
motion between the wave elevation and the centre of the hatch cover is calculated assuming
independence between sea-sates. The contribution to the failure probability is limited to the loaded
condition, with minimum freeboard, which is assumed to be representative for 25% of the time in
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the average case (DNV, 1998). The relationship between the relative motion and the pressure load
on the hatch cover is taken from section 0, and it is assumed uniformly distributed over the whole
hatch surface.
(
)
P d = k dyn ⋅ ρ ⋅ g ⋅ k nl ⋅ relu 6 − (h − T ) [kN/m2 ]
Where:
kdyn
= dynamic factor relative to static pressure height, set to 1.0 in the reliability
analysis, see note.
ρ
= 1.025 tons/m3
g
= gravity, 9.81m/s2
nl
k
= non- linear correction factor (=0.91, ref. section 0)
relu6 = linear relative motion in head sea at middle of No. 1 hatch cover at CL
h
= height from base line to hatch cover (=28.48m)
T
= draught at the hatch cover (=18.15m)
Note:
The factor of 1.0 corresponds to ships with flat deck, and factors as high as 1.4 may also be relevant, see
section 0. The factor of 1.0 may be optimistic, however, for the present purpose where quantifying the relative
difference in failure probabilities for different designs, one would like the calculated probability of failure for
the base case to be near the observed empirical value. This is obtained using a factor of 1.0. To be physically
more realistic, one could chose to increase kdyn , and introduce other factors accounting for such as "bad
weather avoidance" and possibly some reserve strength in the capacity formulation. This has not been done.
The probability of failure increases with an order of magnitude is kdyn is set to 1.4, however, the relative
difference in probability of failure between designs is only slightly reduced compared results for k dyn =1.0.
Capacity Model
The capacity formulation is discussed in Appendix 6.
Uncertainty Modelling
The uncertainty modelling is summarised in Table 21. It should be emphasised that the capacity
formulation involves interaction between several failure modes, and the associated uncertainty is
difficult to quantify. Similarly, the physics of the water on deck loading is very complex, and the
uncertainty in the predictions will be a matter of discussion. At the same time, one sho uld
remember that the main result of the present reliability analysis is the relative change in failure
probability as a consequence of an implemented risk control option in terms of structural
strengthening of the hatch cover. The confidence in this rela tive change is higher than in the
absolute magnitude of the calculated annual failure probabilities.
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Table 21 Uncertainty modelling
Variable
Distribution type
Comment
Significant wave height
3-parameter Weibull,
(see below for parameters)
Zero-crossing wave period
Lognormal (see below for
parameters)
Representative for worldwide conditions (BitnerGregersen, 1995)
Conditional on the
significant wave height
(Bitner-Gregersen, 1995)
Conditional on sea state
Extreme
relative
vertical Gumbel, derived from short
motion, hatch – sea surface
term statistices
Model uncertainty, pressure load
Normal:
(Mean=1.0, CoV=0.05)
Yield stress
Lognormal:
(Mean=360MPa,
CoV=0.08)
Model uncertainty, pressure
Normal:
capacity
(Mean=1.0, CoV=0.05)
Assumed
(Skjong et al, 1996)
Assumed
The parameters used in the formulation of the joint environmental model are included below, based
on Bitner-Gregersen (1995). The parameters are representative for a random 3-hour sea-state.
Significant wave height, Hs :
Modelled by a 3-parameter Weibull-distribution
β−1
  hs − γ  β 
β  hs − γ 
f Hs (h s ) = 
exp  − 
 
α  α 
  α  
with scale parameter α=2.387, slope parameter β=1.470 and location parameter γ=0.385.
Zero-crossing wave period:
Modelled with a lognormal distribution, conditional on the significant wave height
 (ln t z − µ) 2 
1
f Tz | Hs (t z | hs ) =
exp 

2
2πσt z
 2σ

where
and
µ = E (ln T z ) = −1.010 + 2.847 hs 0.075
σ = Var (ln Tz ) = 0.161 + 0.146 e −0.683hs
Additional comments to the uncertainty modelling:
− Uncertainty in material strength, yield stress to be modelled with a Lognormal distribution with
a CoV of 8%. The yield stress is specified to 315 MPa, and assuming this to be 5% fractile, the
mean value is 360 MPa.
− The uncertainty in the capacity formulation is difficult to quantify due to complexity in the
interaction between failure modes. The uncertainty in the combined plate/stiffener buckling
stress is applied in terms of an uncertainty factor with a mean value of 1.16 and a CoV of 16%,
(Skjong et al, 1996). The uncertainty factor has been applied to σun .
− Uncertainties in geometrical measures have been ignored
Results
Results in terms of the annual probability of failure are included for the three different designs.
Results using a uniformly distributed wave heading together with short-crested sea should be used,
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and the results considering long-crested head sea are included for comparison. The results are
summarised in Table 22.
Both First and Second Order Reliability Method (FORM and SORM) have been used in the
reliability analysis. The difference between FORM and SORM is important when the heading is
modelled as a uniformly distributed variable; i.e. The FORM probability of failure is typically a
factor of 4 times the SORM result. FORM does not properly account for the response variation
with heading when integrating the probability. The difference between FORM and SORM is small
when the heading is fixed.
Table 22 Annual probability of hatch cover failure using SORM
Case
Initial design
UR S21
UR S21 + 30% load
pf
9.35⋅⋅10 -4
1.16⋅⋅10 -5
7.85⋅⋅10 -7
Colman (2000) reports the re-opened formal investigation into the loss of the M/V Derbyshire in
1980. Derbyshire was a capesize OBO carrier with hatch covers designed to ILLC 66. The report
gives a No. 1 hatch cover capacity of 42 kN/m2 and predicts a capacity of 83 kN/m2 if the hatch
covers had been designed according to IACS UR S21. If the 35.8 kN/m2 capacity of the ILLC 66
design used above was replaced by the 42 kN/m2 in the above reliability analyses, the probability of
failure of No. 1 hatch cover would be reduced. The estimated probability of No. 1 hatch cover
failure would then become lower the probability of hatch cover failure conservatively deduced from
casualty data, which would not be unreasonable. If replacing the estimated strength of the UR S21
design, of 65 kN/m2 with 83 kN/m2 , the estimated probability of failure of the UR S21 design
would be lower than 10-6 .
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Appendix 8 Marginal Cost Effectiveness
When deciding upon design pressures for hatch covers on new buildings, infinitely many outcomes
exist, since the design pressure is a continuous variable. To be able to evaluate different design
pressure levels by Cost Effectiveness Analysis, the marginal costs and risk reductions need to be
evaluated. This implies that the cost effectiveness of a design pressure is found by looking at
marginal changes of the design pressure.
Structural reliability analyses were carried out for a capesize No. 1 hatch cover design, believed to
be representative for capesize bulk carriers. 5 points were established as shown in the Table below.
Table 23
Structural Reliability results
Capacity (kN/m2 )
Case ID
Base case (ILLC 66)
UR S21
UR S21 + 30%
35.8
65
84.5
Annual probability of failure, Pf
9.35E-04
1.16E-05
7.85E-07
Figure 29 below shows the negative logarithms of the probabilities of failure, together with a linear
regression, as function of the design pressure.
7
Linear regression
6
Structural reliability
results
-log(pf)
5
4
UR S21+30%
UR S21
3
ILLC 66
2
1
0
0
20
40
60
80
100
Figure 30
Annual probability of failure (-log(Pf)) as a function of hatch cover strength
The linear regression is described by the following equation:
− log ( p f ) = 0.0628 ⋅ p c + 0.796
where pf is the annual probability of hatch cover collapse and pC is the hatch cover strength in
kN/m2 .
Similarly, a linear regression was made based on the cost data, as shown in Figure 30.
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180000
160000
UR S21+30%
Costs (US$)
140000
UR S21
120000
100000
Linear fit
ILLC 66
80000
Cost data
60000
40000
20000
0
0
20
40
60
80
100
Figure 31
Cost data shown together with linear regression line.
The linear regression for the costs is described by the following equation:
C = 1166 ⋅ pc + 62,720
where C is the cost in US$ and pc is the hatch cover strength in kN/m2 .
Based on the regression lines, the marginal Gross CAF was estimated as a function of the design
pressure, as given in Figure 31.
1.00E+09
Gross CAF (US$)
1.00E+08
1.00E+07
UR S21
Gross CAF
Criterion
1.00E+06
UR S21+30%
1.00E+05
1.00E+04
1.00E+03
1.00E+02
1.00E+01
1.00E+00
0
20
40
60
80
100
Figure 32
Marginal Gross CAF as a function of hatch cover strength, new-building
requirements.
For a further 30% increase in IACS UR S21 loads, the marginal Gross CAF is estimated to US$86
million.
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